NIST SPECIAL PUBLICATION 813

ICLASS-91

1

ICLASS-91

Proceedings

of the

Fifth International Conference on

Liquid Atomization and Spray Systems

Institute for Liquid Atomization and Spray Systems

National Institute of Standards and Technology

Hratch G. Semeijian, Editor

July 15-18, 1991

National Institute of Standards and Technology

Gaithersburg, MD, U.S.A.

U.S. Department of Commerce

Robert A. Mosbacher, Secretary

National Institute of Standards and Technology

John W. Lyons, Director

Organized by:

(ILASS-Americas)

and
National Institute of Standards U.S. Government Printing Office

and Technology Washington: 1991

Special Publication 813

Natl. Inst. Stand. Technol.

Spec. Publ. 813

869 pages (July 1991)

CODEN: NSPUE2

For sale by the Superintendent

of Documents

U.S. Government Printing Office

Washington, DC 20402
FOREWORD

The Fifth International Conference on Liquid Atomization and Spray Systems, ICLASS91, is hosted by the Institute for Liquid Atomization and Spray Systems (ILASS-Americas) and

the National Institute of Standards and Technology. Previous international ICLASS conferences

were held in Tokyo, Japan (1978), Madison, WI, USA (1982), London, UK (1985), and Sendai,

Japan (1988). Regional ILASS meetings are held annually in the USA, Japan, and Europe.

Proceedings have been published of papers presented at all ICLASS conferences. There has been

considerable growth in interest in the subject of Atomization and Sprays as seen in the increased

membership of ILASS world wide, and the considerable increase in the number of papers

presented at mechanical, chemical, aeronautical and aerosol engineering societies. A new

journal, Atomization and Sprays, the journal of the International Institutes for Liquid Atomization

and Spray Systems, was founded in 1991. It publishes papers of specialist interest, with a wide

range of applications, in an archival journal with easy access for librarians.

The papers presented at ICLASS-91 have been classified into areas of fundamental interest

involving hydrodynamic instability, wave mechanics, and liquid disintegration as part of the

process of atomization. New insights and information are leading to an improved understanding

of the interactions between fluid dynamics and aerodynamics in two-phase flow. An important

step forward in the characterization of sprays has been made by measurements with laser-based

instruments; these instruments allow measurement of drop size, velocity, number density and

liquid flux. Laser-based particle diagnostic techniques are being used in hundreds of laboratories

worldwide, and a significant fraction of the papers presented at ICLASS-91 report on

measurements made with such systems.

Important developments are being made in theoretical analyses and computations.

Classical theories on wave mechanics, instability, and amplitude growth are being used to analyze

liquid disintegration. Theories have been developed, based on particle-fluid mechanic

interactions, to explain dispersion, collision, coalescence and evaporation of drops in sprays.

Following the general trend in engineering science of rapid growth in computational fluid

dynamics (CFD), there is a signiflcant increase in the number of papers presented at ICLASS91 on numerical analyses, computations, predictions and modeling using complex codes that

account for two-phase, turbulent, three-dimensional interactions between particles and gas under

conditions of heat, mass and momentum transfer. The use of supercomputers is providing closer

agreement between predictions and experimental measurements.

Applications research has, in the past, been dominated by liquid fuels and combustion.

The influence of sprays on combustion and emissions will become of much greater importance

in the design of the next generation of aircraft engine combustors. Research on diesel engine

sprays has flourished with the realization of the strong influence of atomization and mixing in the

cylinder on emission of soot. The largest number of papers presented at ICLASS-91 falls into

the category of internal combustion engines. Liquid metal spraying for coating, powder

metallurgy and manufacturing of tools is a fast growing field in the aerospace industry. The

fundamental processes of atomization in liquid metal sprays are not so different from those in

spray combustion. Rapid solidification processes, which result in three-phase flows, provide

improved material properties. Impingement of drops on a substrate is critical for the ultimate
strength and metallurgical properties of the end product. Many millions of gallons of paint are sprayed onto automobiles and many other commercial products. Forty percent of the spray

misses its target. Blistering and nonuniform film thickness are major problems in the industry.

ICLASS-91 shows how scientific analysis and detailed instrumentation can lead to important

improvements for an industry that is huge and very old. Agricultural spraying is resulting in

large scale pollution of the environment. Society could benefit greatly from utilization of

knowledge acquired in other applications. Agricultural industry is poorly represented at ICLASS91.

Summarizing the state of the art as represented by the 100 papers presented at ICLASS91 and printed in these proceedings, we see significant advancements in elevating atomization and

sprays to a significant specialty in the field of science. The increased use of theory and

computational fluid mechanics is providing a framework for generalization and prediction which

should ultimately lead to more accurate computations and fewer measurements. Atomization, a

subject declared by many experts, in the past, to be too complex for study, is being tackled by

theoreticians, computationalists and experimentalists using novel techniques to measure

wavelength, frequency, amplitude growth rate, and wave propagation speeds on liquid surfaces.

This research is showing that atomization is a highly non-linear process with steady input of

liquid to the nozzle resulting in highly unsteady, pulsating flow in the spray. As a result of very

detailed measurements in sprays, providing information of distributions of drop size, velocity,

number density and liquid flux distributions, designers of nozzles, guns, spray chambers,

combustion chambers, and chemical processing plants are beginning to realize that a scientific

approach can lead to a more effective solution of design and practical problems than the

traditional trial and error methods which are currently used in most industries. All of these

changes are reported in the wealth of information contained in the papers published in these

proceedings.

Norman Chigier

Conference Chairman, ICLASS-91

NOTE:

Opinions expressed at the Conference, or published in the Conference Proceedings, are those of

the authors and contributors, and do not necessarily represent the views of the Organizing

Committees. Certain commercial equipment, instruments, or materials are identified in the

papers in order to adequately specify experimental procedures. Such identification does not imply

recommendation or endorsement by ILASS-Americas or the National Institute of Standards and

Technology, nor does it imply that the materials or equipment identified are necessarily the best

available for the purpose.

Iv
ICLASS-91

Fifth International Conference on

Liquid Atomization and Spray Systems

ORGANIZING COMMITTEES

Conference Committee

Norman A. Chigier - Conference Chairman (ILASS-Americas)

Hratch G. Semerjian - Conference Co-Chairman (NIST)

Arthur H. Lefebvre - Chairman, Papers Committee

Harold C. Simmons - Treasurer

G. Scott Samuelsen - Advisor

Publication Committee

Hratch G. Semerjian (Chairman)

Ashwani K. Gupta

Deborah J. Musumeci

Gary Presser

James R. Zurlo

V
I

I

I

i

I
TABLE OF CONTENTS

Page

FOREWORD iii

ORGANIZING COMMITTEES v

Invited Lectures v

A. The Physics of Atomization - N.A. Chigier. 1

B. Experimental and Theoretical Studies on the Structure of Fuel Sprays

in Diesel Engines - H. Hiroyasu 17

C. Atomisation of Liquid Metals for the Manufacture of Metal Powder

and for Spray Forming - K. Bauckhage 33

D. Twin-Fluid Atomization: Factors Influencing Mean Drop Size - A.H. Lefebvre 49

Technical Papers

Session A-1 EVAPORATION AND MASS TRANSFER IN SPRAYS

1. Relaxation of a Polydisperse Spray Undergoing Vaporization and

Coalescence in a Flow Field Behind a Normal Shock Wave: A Sectional

Approach - S. Zahavi and Y. Tambour. 65

2. Effect of Interaction on the Rates of Evaporation of a Group of Drops

and Prediction of Void Fraction - /. Taniguchi and K. Asano 73

3. Influence of Turbulence on the Mass Transfer from Liquid Fuel

Droplets - /. Gokalp, C. Chauveau, O. Simon and X. Chesneau 81

4. Supercritical Droplet Evaporation - A. Umemura, X.-Y. Chang and

T. Fujiwara 89

5. Water Droplet Evaporation in Fire Plumes - G. Krishan, R.P. Gakkhar

and S. Prakash 97

vii
Session B-1 SPRAY COATING

6. New Research Approach to Air Spray Painting - K.-C. Kwok and

B. Y.H. Liu 105

7. Coatings Applied with the Plasma-Spray Process Using Nickel-Aluminum

Powders -D.J. Varacalle, Jr., J.R. Knibloe, V.L. Smith-Wackerle,

J.A. Walter and G. Irons 113

8. Atomization Characteristics of a High-Speed Rotary-Bell Paint

Applicator - P.L. Corbeels, D.W. Senser andA.H. Lefebvre 121

9. Measurements in Paint Sprays Using a Phase-Doppler Anemometer -

/. Domnick, C. Tropea and T.-H. Xu 129

10. Atomization, Charge and Deposition Characteristics of Dual Polarity

Electrostatically Charged Aircraft Sprays of Serrated Spinner Cup

Atomizers - K.D. Kihm, B.H. Kim andA.R. McFarland 139

Session A-2 BASIC ATOMIZATION PROCESSES

11. A Numerical Model of the Capillary Instability - R.W. Sellens 149

"12. Atomisation of Flat and Annular Liquid Sheets - Practical use of Linear

Theories - C. Dumouchel and M. Ledoux 157

13. Flow Characteristics of Atomizing Liquid Jet by Supersonic Flow and

Related Shock Wave/Boundary Layer Interaction - T. Arai, H. Sugiyama,

N. Uno and T. Takahashi 165

14. Shock Wave Interaction With Liquid: Liquid Atomization and Analysis

Schlieren, Light Sheet and Holographic Records - N.M. Fargdre 173

15. Dynamic Similarity Between Liquid Sheets and Hard Spring Systems -

A. Mansour and N. Chigier. 181

16. An Experimental Study of Air Entrainment by Multiple Spray Nozzles -

A^. Dombrowski and J. Singh 189

17. Behavior of Surface Waves on the Breakup Mechanism of a Liquid

Column - D.-J. Kim and C.-W. Lee 197

vili
Session B-2 PRESSURE ATOMIZERS

18. Liquid Atomization with Hollow-Cone Nozzles - H.D. Dahl and

E. Muschelknautz 205

19. Drop Size-Velocity Correlation and Drop Flux in the Spray Cone of a Fan

Spray Atomiser - G. Schulte and I. Ozdemir. 215

20.. Numerical Analysis of Pressure Swirl Atomizers - C.-P. Mao and

S.G. Chuech 225

21. The Spray Structure from Swirl Atomizers (Part 1: General Structure of

the Spray) - N. Tokuoka, Y. Yamaguchi, M. Takada and F. Zhang 233

22. The Spray Structure from Swirl Atomizers (Part 2: Effect of Injection

Pressure and Atomizer Characteristics on the Spray Structure) - F. Zhang,

Y. Yamaguchi, M. Takada and N. Tokuoka 241

23. Two-Component Phase-Doppler Measurements from the Spray of a SwirlPressure Burner-Nozzle in the Vicinity of the Nozzle - K. Lampe,

K. Bauckhage and B. Haase 249

24. Characterization of Industrial Air Atomized Liquid Sprays for Fluidized

Bed Granulators - T. Isenschmid 257

Session A-3 DIESEL SPRA YS - /

25. Correlations for Diesel Spray Penetration including the Effects of the

Break-up Zone - AJ. Yule, M.R. Mirza and I. Filipovic 267

26. Break-up Length of a Liquid Jet and Internal Flow in a Nozzle - H. Hiroyasu, M. Arai and M. Shimizu 275

27. An Experimental Study on Initial Behavior of Diesel Fuel Spray

Characteristics - J.-C. Huh, C.-Y. Lee and O.-Y. Yang 283

28. Modeling Diesel Engine Spray Vaporization and Combustion - M.A. Gonzalez and R.D. Reitz 291

29. Matching of Engine Chamber and Fuel Supply System Via KIVA-II

Modeling - K.C. Tsao, Y. Dong and Y. Xu 299

30. Experimental Investigation on Micro-Explosion of Emulsified Diesel Oil

by Holography - X. Hang, S. Yunbiao, Z. Chongji and M. Yuanji 307

ix
Session B-3 DROP SIZE DISTRIBUTIONS OF SPRAYS

31. The Three-Parameter Log-Hyperbolic Distribution and Its Application to

Particle Sizing - T.-H. Xu, F. Durst and C. Tropea 315

yi. A Drop Size Distribution Data Base for Forest and Agricultural Spraying:

Potential for Extended Application - M.E. Teske, P.J. Skyler and

J.W. Barry 325

33. Effects of Initial Droplet Size Distributions on Polydisperse Spray Jet

Far-Field Diffusion Flames - Y. Tambour, I. Silverman and

J.B. Greenberg 333

34. Application of the Singular Value Decomposition to the Inverse

Fraunhofer Diffraction Problem - S. Huzarewicz, G. W. Stewart and

C. Presser. 341

35. The Angular Pattern of the Scattered Light and the Lorenz-Mie Theory

for the Study of Oil Sprays - F. Ossler, F. Beretta, A. D'Alessio and

A. Tregrossi 349

36. Experimental and Numerical Studies on Droplet Distribution in Diesel Oil

Sprays - K.-J. Choi, J. Xin and Y.B. Shen 357

Session A-4 NOVEL ATOMIZERS

37. A New Atomization Mechanism for Airless Spraying: The Supercritical

Fluid Spray Process - K.A. Nielson, C.W. Glancy, K.L. Hoy and

K.M. Perry 367

38. Droplet Diameter, Flux and Total Current Measurements in an Electrohydrodynamic Spray - P.F. Dunn and S.R. Snarski 375

39. Induced Atomization - S.P. Lin and D.R. Woods 383

40. Production of Uniformly-Sized Dual Concentric Droplets from Coaxial

Smooth Jet under Applied AC Electric Field - T. Sakai, M. Sadakata,

M. Sato, K. Kimura and M. Saito 389

41. Atomization of Liquid Cylinders, Cones, and Sheets by AcousticallyDriven, Amplitude-Dependent Instabilities - J.L. Dressier. 397

42. Atomisation and Combustion Performance of a Stem-Jet Type Twin-Fluid

Acoustic Atomiser - N.R. Ramesh and R. Natarajan 407

X
Session B-4 DIAGNOSTICS -

1

43. Calibration Factors for Laser Diffraction Ring Detectors: Theoretical

Modeling - S.B. Kenney and E.D. Hirleman 415

44. Particle Diagnostics and Turbulence Measurements in an Isothermal

Spray Under the Influence of Stabilizing Bodies - A. Bre^ de la Rosa and

W.D. Bachalo 423

45. Atomization with Spill-Controlled Swirl Pressure-J€t NmzIcs - M. Ldffler-Mang and W. Leuckel 431

46. Multiple Scattering and Size Distribution Effects on the Performance of a

Laser Diffraction Particle Sizer - T. Paloposki and A. Kmkkunen 441

47. Pulsed Laser Diffraction Method for Measuring the Mean Drop Size of a

Diesel Spray in an Evaporation Process - M. Tabata, H. Fujii, M. Arai,

K. Nishida and H. Hiroyasu 449

48. Time-Based Polarization Ratio Measurements of Droplet Size - J.R. Zurlo,

C. Presser and N. Chigier. 457

Session A-5 ATOMIZATION OF SPECIAL LIQUIDS

49. Generation of Spherical and Monodisperse Polymer Particles by

Atomization of Monomers or Dissolved Polymer Precursors -

T. Panagiotou and Y.A. Levendis 465

50. Characterisation of Sprays of Bitumen-in-Water Emulsions - N. Marcano

and A. Williams 475

51. The Influence of Spray Parameters on Local Mass Fluxes jmd Deposit

Growth Rates During Spray Compaction Process - V. Uhlenwinkel,

U. Fritsching and K. Bauckhage 483

52. Numerical Modelling in the Metal Spray Compaction Process -

U. Fritsching, H. Liu and K. Bauckhage 491

53. Simulation of a Spray Combustion Behavior for Two Kinds of Slurry Fuels

(Effect of Two Stage Air Introduction and Spray ChanKJteristics on

Combustion) - H. Aoki, T. Furuhata, S. Tarmo, T. Miura, and

M. Daikoku 499

54. Cryogenic Liquid-Jet Breakup in Two-Fluid Atomizers - R.D. Ingebo 507

xi
Session B-5 SPRAY COMBUSTION

55. Effect of Air Swirlers on Atomization and Spray Flame Stability -

C. -P. Mao 513

56. Effect of Dodecanol Content on the Combustion of Methanol Spray

Flames - C. Presser, A.K. Gupta, C.T. Avedisian and H.G. Semerjian 521

57. Spray Characterization in Axisymmetric, Swirling Combustor Flows - M.J. Ward, J. A. Bossard, R.E. Peck and E.D. Hirleman 529

58. Spark-Ignited Spherical Flames Propagating in a Suspended Droplet

Cloud - K. Nakabe, Y. Mizutani, F. Akamatsu, M. Fuchihata and

S.H. El-Emam 539

59. Some Studies on Flame Interference Effects of Burning Methanol

Droplets - B. Roy, R. Chandrasekhar, S. Bhattacharyya and A.K. Ghosh... 547

60. Atomization and Combustion of Liquid Fuels in Counter-Flow - B.N. Raghunandan, G. Jagannath and D.S. Mehra 555

Session A-6 DIESEL SPRAYS - //

61. Similarity Between the Break-up Lengths of a High Speed Liquid Jet in

Atmospheric and Pressurized Conditions - M. Arai, M. Shimizu and

H. Hiroyasu 563

62. Effect of Nozzle Configuration on the Atomization of Steady Spray -

T. Karasawa, M. Tanaka, K. Abe, S. Shiga and T. Kurabayashi 571

63. A Study on Spray Flow Issued from Plain Jet Atomizer using Phase/

Doppler Particle Analyzer - T. Inamura, N. Nagai and T. Hirai 579

64. Spray Characteristics of a Pilot Injection System Operating in an Idling

D.I. Diesel Engine - F. Yoshizu and M. Nakayama 587

65. A Few Useful Relations for Cavitating Orifices - F. Ruiz 595

66. Holographic Diagnostics for Diesel Spray Research - Z. Huang,

L.-F. Zhang and B.-Z. Li 603

xli
Session B-6 DIAGNOSTICS - II

67. Results of the ASTM Interlaboratory Study on Calibration Verification of

Laser Diffraction Particle Sizing Instruments using Photomask Reticles - E.D. Hirleman, P. G. Felton and J. Kennedy 611

68. Correcting Malvern Particle Size Measurements for Phase Distortion - R.A. Pietsch, P.E. Sojka and G.B. King 621

69. A Coaxial-Beam Drop Sizing Instrument for Dense Sprays - W.M. Grissom 629

70. Measurement of Droplet Sizes by Liquid Nitrogen Freezing Method -

T. Karasawa, S. Shiga and T. Kurabayashi 637

71. LDV and PDA Analysis of Spray-Wall Interaction - G. Brunello,

S. Callera, A. Coghe, G.E. Cossali and F. Gamma 645

72. Application of Poisson Statistics to the Problem of Size and Volume Flux

Measurement by Phase-Doppler Anemometry - C.F. Edwards and

K.D. Marx 653

Session A-7 TWIN-FLUID ATOMIZATION

73. Classification of Round Liquid Jet Disintegration in Coaxial Air Streams -

Z. Faragd and N. Chigier. 661

74. Experimental Investigation of Two-Phase Flow in the Near Field of an

Airblast Atomizer - P. Hebrard, P. Trichet and X. Bardey 669

75. Liquid Sheet Instability in a Coflowing Air Stream - H. Eroglu and

N. Chigier. 679

76. Spray Gas-Phase Interaction Downstream of a Co-Axial CounterSwirling Dome Swirl Cup - H. Y. Wang, W. Sowa, V. G. McDonell and

G.S. Samuelsen 687

77. Air Blast Atomization: Instability of an Annular Liquid Sheet

Surrounded by Two Air Flows of Different Velocities - P. Camatte and

M. Ledowc 695

78. Liquid Jet Instability in Coaxial Air Flow - H. Eroglu and N.A. Chigier.... 703

III

xiii
Session B- 7 SPRAY MODELING

79. Efficient Spray Modeling on Various Applications using the Group

Approach - Q. Zhou and S. C. Yao 711

80. Axisymmetric Calculations of Three-Droplet Interactions - C.H. Chiang

and W.A. Sirignano 719

81. Solutions of Polydisperse Spray Sectional-Equations via a Multiple-Scale

Approach: An Analysis of a Premixed Polydisperse Spray Flame -

/. Silverman, J.B. Greenberg and Y. Tambour. 727

82. On the Evaporation of Binary-Fuel Clusters of Drops Having a Large

Volatility Differential - K. Harstad and J. Bellan 735

83. Modelling of Electro-Aerosol Deposition in Human Airways - W. Balachandran, C.N. Ahmad and S. A. Barton 745

84. Droplet Interaction in Spray Modelling - M.M. Elkotb, O.M.F. Elbahar,

T.I. Sabry and S.A. Wilson 753

Session A-8 AIR BLAST AND ELECTROSTATIC ATOMIZATION

85. Study of an Air Blast Coaxial Atomizer - Experiments and Modelisation -

/. Care and M. Ledowc 763

86. Characterization of a Low Flow Rate/High Charge Density Electrospray -

A. Gomez and K. Tang 771

87. Characterization of the Spray Structure of a Large Scale H.F.O. Atomizer -

G. De Michele, M. Graziadio, F. Morelli and G. Novelli 779

88. Characterization and Performance of Spray Nozzles for Water

Distribution in Cooling Towers - L. Tognotti, S. Zanelli, B. Bellagamba,

F. Mattachini and G. Lotti 787

89. Suppression of Smoke Particles using Electrostatic Spraying -

C.N. Ahmad and W. Balachandran 797

90. Atomization and Dispersion of Quasi-Monodisperse Electrostatic Sprays of

Heptane - A. Gomez and K. Tang 805

xiv
Session B-8 AIRBREATHING AND ROCKET ENGINES

91. Evaluating the Influence of Fuel Preparation on the Performance of a

Spark-Ignited Engine - C.H. Daniels, L.W. Evers and Z. Han 813

92. Effect of Flow Rate on the Spray Characteristics of Impinging Water

Jets - P. Vassallo and N. Ashgriz 821

93. Analytical and Experimental Study of Primary Atomization of Water Jets -

A.J. Przekwas, J.-G. Lee, K.W. Gross, N.A. Chigier and H. Eroglu 831

94. Disintegration Phenomena of Metalized Slurry Fuel Jets in High Speed

Air Stream - T. Inamura, N. Nagai, T. Hirai and H. Asano 839

95. Thermodynamics and Applications of a Spray Interacting with a

Compressible Flow - F. Ruiz 847

96. Spray Characteristics of Pintle Type Injector Under Low-Pressure Field -

J. Senda, M. Yamaguchi, T. Wdcashiro, T. Tsukamoto, K. Hojyo and

H. Fujimoto 857

INDEX OF A UTHORS. 865

XV

ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper A

Plenary Lecture

THE PHYSICS OF ATOMIZATION

N.A. Chigier

Department of Mechanical Engineering

Carnegie Mellon University

Pittsburgh, PA, U.S.A.

ABSTRACT

The fundamental physical processes of atomization of liquid sheets and jets are examined.

Disturbances are generated inside nozzles in both liquid and gaseous flows due to turbulence,

boundary layer growth, asymmetries, sharp bends, cavitation, pulsations in the flow, and

vibrations of the nozzle. When the liquid emerges from the nozzle, the initial disturbances grow as a function of time and distance from the nozzle. Wave amplitudes grow until the liquid sheet

erupts to produce a shower of drops. Pulses of liquid form clusters of drops with the same

frequencies as those of the waves on the liquid surface. Growth of the waves on the free liquid

surface is augmented by 1) pressure variation and shear generated by relative velocity at the

gas/liquid interface 2) turbulence intensity and scale in the gas and liquid streams and 3) relaxation

of the velocity profile in the liquid stream as it adjusts from the no-slip boundary condition at the

nozzle wall to the equilibrium condition at the gasAiquid interface. New information has been

obtained on frequency, wave length, and wave amplitude measured along the length of liquid jets

and sheets for a wide range of liquid and gas velocities.

Large scale structures in the air flow generate large, ordered cell structures within the liquid

sheet. These structures are bounded by large diameter ligaments with thin membranes stretched

between them. The larger diameter ligaments are the origin of the large drops in the spray, and the

stretched membranes are the origin of the smaller droplets. Co-axial round liquid jets, surrounded

by axisymmetric air streams form flat curling liquid sheets with subsequent breakup into droplet

clouds. Intact lengths were found to be strongly dependent on Reynolds and Weber numbers.

Final breakup mechanisms of flat sheets, round jets, and large drops are shown to be remarkably

similar.

INTRODUCTION

Even though atomization has been used very extensively in a wide variety of applications, the

fundamental physical mechanisms of liquid breakup into drops have not yet been established.

When the word atomization was coined in the 19th century, there was a belief that 5 (im was the

smallest possible diameter of droplets that could be physically atomized. Today, there are a number of atomizers that generate much smaller droplets of submicron size. Theoretical analyses

based on wave theory have postulated wave growth along surfaces but until recently there was little

experimental evidence on generation and growth of waves. The influence of the fluid properties,

1
viscosity, surface tension, and density have always been considered important though there is

conflicting evidence as to the quantitative significance of these properties. For two-fluid atomizers,

it has long been established that increasing the relative velocity between the air and liquid flow

streams has a very powerful and dominant influence on atomization. Fluid dynamic effects of

turbulence in the liquid and air streams, boundary layer growth, flow separation at sharp bends and

corners, cavitation, and pulsations in the supply lines all contribute to the generation of

disturbances. After emerging from the nozzle, wave amplitudes grow under the influence of

relative velocity between the air and liquid flows. Shear develops in the liquid and air flows

resulting in turbulence generation with changing intensity and scale. Velocity profiles in the liquid

flow relax from the no-slip wall condition inside the nozzle to the equilibrium condition at the

liquid-air interface.

Several mechanisms have been postulated for the origin of disturbances. Turbulence

generated inside the nozzle is the most likely source. Turbulence inside the nozzle can be generated

by flow separation at sharp comers, wall roughness, boundary layers, and shear flow. The radial

transverse component of the fluctuating velocity is directed towards the liquid surface and can be

expected to initiate instabilities by formation of protuberances on the liquid surface. Boundary

layer instability along the inner walls of the nozzle, cavitation inside the nozzle, sudden pressure

drop at the nozzle exit, sudden velocity redistribution as the velocity profile relaxes on emergence

from the nozzle and interfacial hydrodynamic instability are all additional sources for disturbance

initiation. Lin and Woods, based on their experiments, conclude that interfacial instability is the

real origin of atomization. Atomization usually produces a wide size distribution. A narrow size

distribution can be achieved by controlling the external excitation at a precisely fixed frequency.

Knowledge of the physics of the atomization process will lead to improvements in atomizer

design. The major objective in the design of atomizers is to generate a specified size, velocity and

number density distribution of drops resulting in sprays with specified angles and liquid mass flux

distributions. Sprays with radial and circumferential symmetry are generally required.

In the past, Sauter Mean Diameter has been used as a measure of atomization quality. Sprays

with large differences in size distribution can yield the same global SMD so that the global SMD is

an insufficient indicator of atomization quality. There has been very little information on drop

velocities in sprays. The performance and effectiveness of a spray is dependent on individual drop

momentum which determines individual drop trajectories. Knowledge is, therefore, required of

both size and velocity of individual drops in the spray. Many processes will prescribe an upper

limit of droplet size so as to prevent deposition on chamber walls. Other processes, such as

agricultural spraying, will prescribe a lower limit on droplet size to avoid drift and carry over of

small droplets by wind and cross currents of air flow. For processes such as coating of surfaces or powder metallurgy, uniform drop size may be required. The ultimate objective is to have

sufficient control of drop size, velocity, spacing and angle of flight. In order to obtain control of

the spray, all the factors that affect the atomization process must be controlled throughout the

spraying process.

In addition to the fluid dynamic processes, external vibration, acoustic, electrical and magnetic

forces can amplify or damp disturbances. Piezoelectric transducers generate monosize drop

streams. Ultrasonic atomization produces clouds of drops at the nozzle exit. Electric fields are used to deflect drop trajectories. Loudspeakers generate acoustic fields which interact with waves

on liquid surfaces to cause damping or enhanced atomization. Coupling of these external fields

with the fluid dynamic fields can lead to far greater control of atomization than has been achieved

up to now.

Special attention needs to be addressed to the problem of pulsations. Most studies in the past

have concentrated on the long-time averaged properties of the spray. More recent studies, using

high speed movies and submicrosecond high speed flash and laser photography, have revealed the
presence of pulsations which are present even under conditions where liquid and air flows supplied

to the atomizer are free of pulsations. The disturbances and waves that propagate along liquid

surfaces have specific frequencies. When the liquid sheets erupt, droplets are ejected in pulses

with the same frequencies as the waves on the surface. These pulsations seem to be present in all

sprays. The clustering of drops has important adverse effects resulting in 1) hot streaks impacting

on combustion chamber walls, 2) reduced evaporation of drops within clouds where vapor

becomes saturated, 3) increased emission of pollutants, 4) incomplete heat and mass transfer, 5)

reduced combustion efficiency, 6) combusticMi instability, and 7) patchiness and non-uniformities

in coatings. Since these pulsations are inherent in the break-up process, it will be very difficult to

eradicate this phenomenon. Application of extemsd acoustic and electrical fields will be needed to

reduce the impact of these pulsations and clusters.

CLASSICAL BREAKUP REGIMES

The Rayleigh regime appears at low velocity when the growth of axisymmetric waves causes

the detachment of drops whose size is larger than the jet diameter.

The First Wind Regime is characterized by the inception of aerodynamic interaction between

the surrounding stationary gas and the liquid jet. The pressure force, induced by the aerodynamic

interaction is added to the surface tendon resultii^ in an increase in tiie growth of axial symmetric

waves. Non-axisymmetric surface disturbances and helical deformation of the jet axis are also

present.

The Second Wind Regime appears at higher jet velocities. Small asymmetric unstable waves

form on the liquid surface resulting in the formation of droplets of much smaller size than the jet

diameter. Droplet detachment from the surface begins far from the nozzle exit. The length of the

cylindrical or undisturbed zone decreases when the jet velocity increases.

The Atomization Regime is characterized by droplet detachment close to the orifice exit. The

vertex of the spray angle is near the center of the nozzle exit section. Droplet sizes are much

smaller than the orifice diameter.

Ranz assumed that the diameters of droplets formed from ligaments are proportional to the

wavelength of superficial unstable disturbances. Schweitzer suggested that turbulence was

responsible for the production of superficial disturbances which are broken by the action of

aerodynamic forces. Bergwerk made the hypothesis that the disturbances in the liquid jet are due

to cavitation inside the orifice. Rupe stressed the role played by the rearrangement of the velocity

profile on the jet stability. A laminar flow inside the nozzle is more unstable than a turbulent flow

because the laminar profile requires a more drastic rearrangement after it leaves the nozzle than a

turbulent profile.

Reitz and Bracco suggested that the outer siaface of the liquid is disrupted by the interaction

with the ambient gas which leads to the r^id and sdective growth of surface waves whose initial

amplitudes are controlled by cavit^n and the rearnmgement of the velocity profile.

Lane concluded tiiat there are 3 stages of disintegration:

1) Initiation of small disturbances cki the liquid surface in the form of ripples or protuberances.

2) Action of air pressure and tangential forces on these disturbances, forming ligaments which

may breakup into drops.

3) Breakup of drops into smaller drqps.

Aerodynamic forces have a major effect on fineness of atomization. The aerodynamic

interaction between the liquid jet and the surrounding gas increases the instability resulting in

shorter breakup lengths and smaller drops.

3
De Juhasz and Schweitzer concluded from their experiments that turbulence in the liquid flow

was the main influencing factor in atomization. Mehlig stated that the radial components of

velocity in the turbulent liquid flow are the predominant cause of atomization. Wave growth theory

has been used with Rayleigh's instability analysis by Chuech and Krulle to calculate jet

disintegration.

Strazhewski, Scheubel and Weber concluded that primary disturbances and wave growth are

relatively unimportant compared to the dominant effects of air resistance and gas-liquid relative

velocities. Previous reviews of atomization by Giffen and Chigier have provided no clear and

distinct explanation of the dominant cause of liquid disintegration.

ANALYSES OF WAVE GROWTH

In 1878, Lord Rayleigh produced the first mathematical analysis of the break down of a simple liquid jet by surface tension. Squire carried out an analysis of wave growth on a thin

inviscid liquid sheet. The analysis indicated the basic principles by which the surrounding

atmosphere interacts with the sheet and predicted the existence of an optimum frequency at which a wave would grow most rapidly.

Dombrowski set up a simple model of the break-up process of a spray sheet. He assumed that

the sheet breaks at the crests and troughs of the waves. Each half-wave-length contracts under

surface tension to form a cylindrical thread which then breaks down to produce a string of

droplets. This physical model greatly oversimplifies the real process but nevertheless provides

general indications of the characteristics of sprays generated by a number of types of atomizers.

The model assumes that the sheet thickness remains constant as the wave amplitude increases so

that the point where the sheet breaks cannot be calculated. Also, the predicted wave frequencies

are at variance with the dominant frequencies found in practice. A more rigorous analysis of wave

growth included a mechanism by which the sheet thickness decreases as the wave amplitude

increases. Trains of ripples on the sheet surface cause localized thinning of the sheet which cause

it to break down earlier than sheets without ripples.

In swirl spray pressure nozzles which form hollow conical sheets, mean drop size decreases

with increase of liquid pressure and increases with increase in viscosity. However, when mean

drop size is plotted as a function of sheet velocity, which is a more fundamental parameter, then

drop size is found to be independent of viscosity. This is a surprising result since it can be

expected that atomization processes would be hindered by viscous forces as is found in some other

forms of atomizers. High speed movie films taken of the flow pattern within an atomizer

constructed of transparent material showed that high viscosity liquids produced the most disturbed

air core. Dombrowski provides the following explanation. Disturbances created in the nozzle will

tend to be damped by surface tension at the air core, but this action will be opposed by viscous

forces. Since the surface tension of oil is relatively low, viscous drag predominates and

disturbances propagate through the sheet where they offset the natural tendency of the viscous

sheet to oppose breakdown.

Even though there is a general recognition that atomizers should be rigidly supported, there is a question as to whether mechanical vibrations can be imparted to the nozzle and thereby affect the

wave initiation and disintegration processes. Dombrowski carried out an experiment to detect these

vibrations by mounting an accelerometer to a fan spray nozzle with the output fed to a spectrum

analyzer. A number of frequencies were found which were of the same order as those present on

the sheets. These frequencies were traced to resonance of the brass tube on which the nozzle was

located, to resonance of the experimental rig, and to disturbances arising from within the building.

It has even been suggested that if the nozzle could be isolated from all external disturbances,

aerodynamic disintegration would not take place. Mean drop size was found to diminish with

4
increase of frequency until a critical value is reached. It then increases to an effectively constant

value. Small upstream disturbances initiate the disintegration process, but the final stages of

breakup are effectively controlled by the interaction of the liquid and the surrounding gas. In the

presence of hot gas, small wave-like disturbances grow in amplitude rapidly leading to break down

of the liquid sheet close to the nozzle. Sheet perforations give rise to a network of threads.

Threads tend to break down in a similar manner to a single column of liquid. Pulsing the liquid

along the plane of the sheet produces symmetric waves.

Previous classifications of the modes of atomization developed by Ohnesorge, Miesse, and

Reitz have shown that modes of atomization are a function of the Reynolds and Ohnesorge

numbers without including the gas-liquid relative velocity. Wolfe and Buschulte used the

aerodynamic Weber number as the single variable for atomization. The breakup of spherical water

drops in an air stream was found by Kennedy to be a function of the air velocity and Wea/Rcw^-^.

Farago has shown that liquid jet disintegration modes are a function of only two parameters Rcj^

and WeaLiquid jet disintegration can be divided into three categories:

• Rayleigh breakup: Mean drop diameter is of the order of the jet diameter. Maximum drop

diameter is twice the jet diameter. There is no ligament formation. The flow may be i)

axisymmetric or ii) non-axisymmetric.

• Formation of membrane ligaments: Round jets develop into thin liquid membranes.

Kelvin Helmholtz waves develop along the membranes. Drop diameters are considerably

smaller than the jet diameter.

- • Fibers are formed on the liquid surfaces. Fibers are peeled off the liquid surface and break

into drops via the Rayleigh mechanism. Drop diameters are an order of magnitude smaller

than the jet diameter.

Pulsations usually develop during the process of atomization. The frequency of release of

pulses of drops is close to the frequency of wave disturbances on the liquid surface. Clusters of

drops are formed which move downstream into the spray.

In recent years, a series of experiments have been conducted at Carnegie Mellon University by

Chigier, Farago, Mansour, and Eroglu with the specific objective of identifying the basic

mechanisms that influence and determine liquid disintegration. Extensive studies have been made

of liquid sheets with co-flowing air jets as well as cyUndrical liquid jets with coaxial co-flowing air

jets. Air and liquid velocities were varied to provide aerodynamic Weber number variations over a range of six orders of magnitude and Reynolds numbers of two orders of magnitude.

TWO DIMENSIONAL LIQUID SHEET AIRBLAST ATOMIZER

The two dimensional liquid sheet airblast atomizer, designed by Parker Hannifin, used at Carnegie Mellon University, is shown in Fig. 1. A liquid sheet emerges from a central slit. High

velocity air impinges, at an angle, on both sides of the sheet at the nozzle exit. Liquid and air gap

sizes and flow rates can be varied. A two dimensional liquid sheet provides an initially flat, large

surface area for study of wave growth and frequency of disturbances travelling along the liquid

surface. It allows direct comparison of measurements with predictions from linear theories of

wave development.

A series of photographs were made of the sheet surface with progressive increases in liquid

flow rate and air supply pressure. With no air flow, the liquid sheet emerges from the slit and

converges under the influence of surface tension. Waves move along the length of the sheet but

also emanate from the rims on either side of the sheet to form criss-cross patterns. Waves from

5
each rim interfere with each other resulting in increases in amplitude when the waves are in phase.

These capillary waves are driven by surface tension. When the liquid flow rate is increased, small

distortions appear on the sheet surface indicating a transition from laminar to turbulent flow at

Reynolds numbers, based on nozzle exit conditions, around 1470.

With air flow, we can distinguish two modes of breakup: a "mechanical" effect due to the

action of liquid pressure inside the nozzle and an aerodynamic effect due to the action of air

friction. Large drops are formed from the rims of the sheet where the aerodynamic forces are much smaller than the surface tension and inertia forces (Fig. 2). The rims behave like columns of

liquid that breakup by the Rayleigh mechanism to produce a line of drops whose diameter is

slightly larger than the liquid column diameter. When the liquid flow rate (and hence the liquid

velocity) is doubled, regularly ordered, organized cell structures are formed across the sheet. Each

cell structure is bounded by large diameter rims with thin membranes of liquid stretched across each cell. As the sheet breaks up, the larger diameter rims generate the larger drops while the

membranes generate the smaller droplets.

Fig. 1 Two dimensional variable geometry air blast research atomizer

Fig. 2 Two modes of sheet breakup exist: A "mechanical" mode due to the action of liquid

pressure inside the nozzle and an aerodynamic mode due to the effect of air fricti(Hi

(Po = 0.1 psig [102.015 kPa]; muquid = 1 1.34 g/s)

6
When the liquid flow rate is increased further, the intact length is increased (Fig. 3). Waves,

increasing in amplitude, accompanied by foimation of several ordered cell structures, can be seen

to clearly develop along the length of the sheet. The cells are stretched and appear to blow up in a

manner similar to the bagging of large single drops during breakup. In Fig. 3, both the mass flow

rate of the liquid and the supply air pressure to the nozzle have been increased. The effects of

increased turbulence in the liquid and air streams are clearly evident. The initial sheet region is

more disturbed with higher frequency and short wavelength waves. Cells have smaller dimensions

and are less ordered. Membrane shattering precedes cell rim breakup. Just before final breakup,

the leading edge of the sheet is anchored to the remaining liquid only through the cell frames. The

high shear causes slight variations in the liquid sheet thickness to develop into major disturbances

resulting in holes being blown through the sheet.

Fig.3 Small incoherent "cell" structures appear inside the larger coherent ones on the liquid

siio^t su.rf3.cc

(Po = 0.2 psig [102.704 kPa]; muquid = 45.42 g/s

Side views of the edge of the sheet are shown in Figs. 4 and 5. The growth in thickness of

the sheet and the development of waves as the result of "flapping in the wind" are clearly seen. Wave amplitudes grow until critical conditions are reached where the liquid bursts into a shower of

droplets. It appears as though some internal force exceeds the external restraining surface tension

force resulting in a bursting explosion or eruption. The condition shown in Fig. 4 looks like a

blown up balloon with almost the whole sheet stretched as a thin membrane, held by a larger

diameter rim. A perforation is seen to form in the membrane resulting in bursting and formation of

a shower of drops. The internal forces within the liquid are associated with shear and turbulence.

If these forces can be shown to grow with distance from the nozzle, then this could provide an

explanation for the conditions for bursting. In Fig. 5, both the liquid and the air flow rates have

been increased. The initially full sinusoidal wave becomes distorted to yield two half waves that are very similar in form. The stretching of the half waves increases the free surface energy of the

already distorted sheet. At the point where the two half waves are joined, localized high tensile

forces pull the half waves in opposite directions and ultimately result in the thinning of the junction

point. When the local thickness of the sheet falls below a critical value, perforations start to

develop in the sheet which ultimately leads to separation of the half waves.

7
Fig. 4 Perforations which expand under Fig. 5 The distortion of a sinusoidal wave

the action of surface tension occur yields two half waves that are very

in the sheet sinular in form

(Same experimental conditions as in Fig. 3) (Po = -2 psig [102.704 kPa]; muquid = 63.09 g/s)

These results show that when the sheet leaves the nozzle it is already perturbed by

disturbances which were initiated within the nozzle. These perturbations appear as deformations of

the liquid-air interface. Initial instabilities are driven by the high velocity air streams causing rapid

growth of the waves. Disintegration occurs when the amplitude of these waves exceeds a critical

value and half waves are torn off the sheet. The half waves break down into ligaments which in

turn break down into drops.

Lines drawn along the crests of major perturbations describe the angle of spread of the

perturbations which may be considered as the initial spray angle. Figure 6 shows the effect of

increasing the liquid flow rate on spray angle for two different air supply pressures. The spray

angle decreases progressively with increase in liquid flow rate. This can be explained by the

reduction in specific energy of air per unit volume of liquid leaving the nozzle. Moreover,

increasing the liquid flow rate results in decreased air/liquid relative velocities. This decrease in

relative velocity reduces the amplitude of oscillations resulting in a corresponding decrease in spray

angle.

Increasing the air velocity increases the specific energy of air per unit volume of liquid and this

results in a substantial increase in the amplitude of oscillations of the liquid sheet. This is shown

as an increase in spray angle with increase in air supply pressure, in Fig. 6. Intact length (axial

length from nozzle exit to the location where liquid remains intact) increases with liquid sheet

velocity and decreases with air velocity, as shown in Fig. 7. At higher air velocities, the influence

of liquid sheet velocity on the intact length is less pronounced. The increase of intact length with

Hquid sheet velocity and its decrease with increase in air velocity are consistent with the variation of

spray angle with these velocities. Higher liquid flow rates result in higher momentum for the

liquid in the axial direction. The liquid remains intact over longer distances and the spray angle

decreases as liquid flow rate, liquid velocity, and liquid momentum increase.

The spray is formed as a result of the breakup of the sheet. We have seen evidence that the

final breakup has the appearance of an explosion producing a shower of droplets. We need to find

the connection between the sheet that is in the final stages of breaking up and the subsequent

formation of the spray.

8
ISO -I

100 -

.s

o

o

0 p. = .1 psig

= .2 psig

a. SO

o

o

o

10

I 20 30 T40

I50 60

I 70

Liquid mass flow rate (g/s)

Fig. 6 Effect of air pressure and liquid flow rate on the spray angle (liquid sheet)

r Pressure

101.574 kPa

102.015 kPa

102.74 kPa

104.312 kPa

108.22 kPa

1 1 1 .667 kPa

0 10 20

Liquid Velocity [mis)

Fig. 7 Variation of liquid sheet intact length with liquid velocity for different air pressures

As soon as drops are released from the liquid sheet, they are accelerated by the air stream.

The time taken for droplet velocities to reach equilibrium with the air stream depends on the

drag/momentum ratio which in turn depends on the drop diameter and relative velocity between

each drop and the local air velocity. Droplets are dispersed in the air flow according to their size

class. Large drops subjected to shear will undergo secondary breakup while collisions of drops

can lead to coalescence resulting in increase in drop size.

ROUND LIQUID JETS IN CO-AXIAL AIR STREAMS

For fundamental theoretical and experimental studies, the cylindrical liquid jet with a co- annular, co-axial, co-flowing air jet provides a base-line for determining the physics of

atomization. Of particular interest is the comparison of the cylindrical jet with the flat sheet which,

in theory, is a cylindrical jet with infinite diameter. We examine the generation of disturbances in

the nozzle and Uie development of waves along the air/liquid interface. Transition from laminar to

9
turbulent flow occurs in the Reynolds number (based on nozzle exit conditions) range 6000 < Re <

10,000 when the inlet flow to the nozzle is uniform, oscillation free , and there is no vibration in

the experimental system.

Axisymmetric Rayleigh breakup was studied for a series of horizontal liquid jets (with no air

flow) with liquid velocity progressively increasing from 0.246 m/s to 18.2 m/s corresponding to

an increase of Rcw from 240 to 17,650. The Ohnesorge number for this series was 0.0037.

When the jet velocity was very low, less than 1 m/s, the horizontal liquid jet bends downwards

under the influence of gravity. At higher liquid velocities, the jet has sufficient momentum to

remain horizontal. The intact length increases progressively with increase in liquid velocity. Drop

diameters are similar to those of the jet and do not change with increase in liquid velocity. Small

satellite drops are formed. Introduction of a co-axial air stream is shown in Fig. 8. The air stream

maintains the Uquid jet horizontal except for the lowest velocity, Uw = 0.246 m/s. The liquid jets

remain axisymmetric with a slight tendency, at higher velocities, to form a helix. The drops also

tend to remain symmetric about the jet axis. Intact lengths increase with increase in liquid

velocities. The length of undisturbed liquid decreases as the air velocity increases beyond 25 nVs.

U„ = II 246 iil/s

= 4.35 m/s Ua = 32.4 m/s

Fig. 8 Axisymmetric Rayleigh disintegration of water jet in a coaxial air stream

(Jet Diameter = 1 mm)

For cases where the air velocity is one or two orders of magnitude greater than the liquid

velocity, the liquid jet first accelerates rapidly in the high speed air stream with a resulting reduction

in the liquid jet diameter. Organized structures in the turbulent air flow penetrate through the liquid

jet. A series of photographs at short time intervals of a liquid jet with a velocity of 1.1 m/s and an

annular air jet velocity of 45.8 m/s shows that the cylindrical liquid jet is distorted and stretched by

the large eddy structures in the air stream to curl up in the shape of a "ladle". The liquid forms a membrane stretched between rims which provides a supporting frame. The stretched membrane

10
tn'eaks up into a shower of small drops while the rims breakup into larger drops. Despite the very

considerable differences in initial geometry, there is a remarkable similarity between the final

breakup mechanism of 1) flat sheets 2) cylindrical jets and 3) large drops. The same phenomena

can be detected even for cases of liquid velocity of 1.5 m/s and air velocity of 130 m/s.

Detailed photographs of the disintegration of cylindrical liquid jets are shown in Fig. 9 for a

liquid velocity of 1.1 m/s and an air velocity of 45.8 m/s. The photographs show a time sequence

of events. The cylindrical liquid jets are transformed into thin membranes which curl up to form

semi-circular cups. These large scale deformations of the liquid jet are most probably caused by

large eddy structures in the air. They cannot be due to instability growth in the liquid. Once the

deformation has taken place, the aerodynamic pressure forces change in the shape, very similar to

the "bagging" phenomenon of breakup of large drops. Impingement of air on the concave

upstream surfaces and wake effects on the downstream surface accelerate the stretching of the

membrane and the breakup into drops. The sheets set up a flapping motion. Detailed

measurcn^nts of size distributions in the sprays show direct connections between the thick rims

which form large drops and the thin membranes which form small droplets. Under the influence

of surface tension forces, thick rims are formed at the edges of the membranes. The thickness of

these rims is larger than the thickness of the membrane but is significantly smaller than the initial

dimneter (rf the liquid jet. The largest drops formed have the same diameter as the rim thickness

while the smallest drops have the same diameters as the membrane thickness. Final breakup

appears to be by the Rayleigh mechanism and can be classified as non-axisymmetric Rayleigh

Fig. 9 Disintegration of coaxial liquid jets

(Pw = 0.7 psig [4.826 kPa]; pair = 0.0433 psig [1.493 kPa])

11
breakup. This membrane type of disintegration of round jets is very similar to that of thin liquid

sheets as described by Dombrowski. It is also very similar to the disintegration of spherical liquid

drops in an air stream as reported by Lane, Wolfe, and Kennedy.

It is quite remarkable that despite the important differences in geometry between cylindrical

jets, thin sheets and large spherical drops, in each case thin membranes are formed with

surrounding rims leading to the formation of large and small drops. In each case, it is the

penetration of large eddy structures from the turbulent air flow that distorts the liquid and stretches

it into thin membranes. The thin membranes and the spray of drops provide a visualization of the

organized large structures in the air.

Under conditions of high liquid velocity and high air velocity, the liquid jets maintain their

symmetry about the jet axis. The liquid core remains intact while fibers are formed on the liquid

surface. These fibers point in the direction of higher velocity. The diameter of the fibers is several

orders of magnitude smaller than the diameter of the liquid jet. The fibers break into drops

according to the Rayleigh mechanism. Disturbances develop in the liquid core which lead to

breakup with the formation of ligaments with lengths of the order of 5 times the initial jet diameter.

Fibers are formed on the ligaments which peel off and then break into drops. This process leads to

the formation of a train of clusters of drops. When the momentum of the liquid flow is high

enough, penetration by the large organized structures of the air flow is prevented.

For high speed annular air jets with a velocity of 183 m/s and liquid jets of 1 m/s, the liquid jet

is broken up into very fine droplets within the short distance of two jet diameters. As the liquid

velocity is increased, the liquid core is extended and the number density of drops increases because

of the increased mass flow rate of liquid. Bursts of liquid can be clearly identified. The size of the

bursts and the liquid concentration inside each burst increases as liquid flow rate is increased. The

high speed annular air jet confines the spray so that it remains symmetric about the jet axis. These

jets have been classified as super-pulsating. Organized structures from the air flow enter the liquid

flow but do not penetrate fully in the case of high liquid flow rate.

Intact length (L) is a measure of the axial distance from the nozzle exit to liquid breakup on the

liquid jet axis. Figure 10 shows the variation of L/d with Weber and Reynolds numbers based on

nozzle exit conditions. The reduction in intact length with increase in Weber number is ascribed to

the increase in relative velocity. The increase in intact length with increase in Reynolds number is

due to the increased liquid velocity and flow rate which, in this Reynolds number range, increases

the length of the liquid jet. The derived empirical equation is:

L/d = 0.5 We^^'^ReJ'^

The variation of intact length with Reynolds number was found to be the same for both flat

sheets and round jets. The subscripts R refer to relative velocity, i to liquid and a to air.

The modes of liquid jet disintegration are shown in Fig. 1 1 where each jet is plotted as a function of liquid Reynolds number and aerodynamic Weber number. Each jet is classified

according to its photographic appearance. For Re^ = KP, as Wca is progressively increased, it can

be seen that up to Wea = 8 disintegration is of the axisymmetric Rayleigh type. At Wea = 17, the

disintegration is of the non-axisymmetric Rayleigh type. By Wca = 33 the membrane type region

is reached and by Wca = 130 fiber type disintegration is found. The particular type of

disintegration can be found easily from Fig. 1 1 with knowledge of Rcj^ and Wca12
o Re=1456

Re=4370

o Re=9328

60

40

20 -

0 40 80 120 160 200 240 280

We

Fig. 10 Average jet liquid intact length in coaxial gas flow

(Re = 1456, 4370, 9328)

u«

mn

•s

«

a

1)3.

Axisymmctric Rayleigh type disintegration

Non-aiisymmetric Rayleigh type disintegration Membrane type disintegration Fiber type disintegration Super—pulsating sub—mode

O

Membrane

type region

Rayleigh type region

Aerodynamic Weber Number

Fig. 1 1 General classification of liquid jet breakup in coaxial air streams

New information on frequency, wave length and wave amplitude measured along the lingth of

liquid jets and sheets for a wide range of liquid and gas velocities is reported by Eroglu and

Chigier, 1991.

CONCLUSIONS

• Close similarities are found between the breakup of 1) liquid sheets 2) round jets and

3) large droplets under the influence of air streams.

• Cylindrical jets and flat sheets curl up to form "ladle" shapes.

• Ordered cell structures form in liquid sheets with thin membranes stretched between thick rims.

• Large drops measured in the spray are formed from thick rims while small drops are formed

from thin membranes in the sheet.

13
• Disturbances generated within the nozzle grow in amplitude along the liquid surface.

• Airflow causes small disturbances to grow into major disturbances that lead to disruption.

• When the thickness of the liquid sheet falls below a critical value, perforations start to appear in

the sheet.

• Disintegration of the liquid sheet occurs when the amplitude of waves exceeds a critical value

and half waves are torn off the disturbed sheet.

• Increasing the liquid flow rate causes decreases in the spray angle while the air flow is

maintained constant.

• Increasing the air velocity causes the spray angle to increase for a fixed liquid flow rate.

• The initial mean velocity distribution in the liquid jet determines the character of the surface

wave formation at the liquid/air interface.

• A new equation is derived for intact length as a function of Reynolds and Weber numbers.

• A new classification of modes of disintegration has been presented.

ACKNOWLEDGEMENTS

The principal ideas contained in this paper have evolved over a number of years of intensive

research at Carnegie Mellon University. I wish to particularly thank Hasan Eroglu, Zoltan Farago,

Adel Mansour and all other memters of my research team who have made important contributions.

I wish to acknowledge the financial support of NASA Marshall Space Flight Center and Parker

Hannifin Corporation.

REFERENCES

Bergwerk, W., Proc. Inst. Mech. Engrs. . 173, 655, 1959.

Buschulte, W., ISABE, 9th InL Symp. of Air Breathing Engines, Athens, 1989.

Cheremisinoff, N.P. (Ed.), "Gas Liquid Flows", Encyclopedia of Fluid Mechanics. Vol. 3. Gulf

Publ. Co., Houston, TX, 1986.

Chigier, N., Energy. Combustion and Environment. Mx^Graw Hill, 1981.

Chuech, S., Przekwas, A., Yang, H., and Gross, K., AIAA, 90-2060, 1990.

De Juhasz, K.J., ASME Trans., Vol. 53, OGP-53-5, 1931.

Dombrowski, N. and Eraser, R.P., Philos. Trans. R. Soc. London Ser. A., Math. Phys. Sci. . Vol. 247, No. 924, pp. 101-130, 1954.

Dombrowski, N., Hasson, D. and Ward, D. E., Chem. Eng. Sci. . Vol. 12, pp. 35, 1960.

Eroglu, H. and Chigier, N., ASME Journal of Fluids Engineering. 1991.

Eroglu, H. and Chigier, N., Atomization and Sprays. 1991.

Eroglu. H. and Chigier, N., "Liquid Sheet Instability in a Coflowing Air Stream", Proceedings of

ICLASS 91, NIST, Gaithersburg, MD, July 199L

Eroglu, H. and Chigier, N., "Wave Characteristics of Liquid Sheets in Cocurrent Air Flow",

Atomization and Spravs. Vol. 2, 1992.

14
Eroglu, H., Chigier, N. and Farago, Z., Physics of Fluids A . Vol. 3, No. 2, pp. 303-308, 1991.

Farago, Z. and Chigier, N., ASME Paper No. 90-GT-81, 1990.

Farago, Z. and Chigier, N., ILASS-Americas 4th Annual Conference, pp. 122-126, Hartford,

CT, May 1990; ILASS-Europe 6th Annual Conference, Pisa, Italy, July 1990.

Giffen, E. Muraszew, A., The Atomization of Liquid Fuels. John Wiley & Sons, New York,

1953.

Hagerty, W. and Shea, J. F., J. Appl. Mech.. Vol. 22, No. 4, pp. 509-514, 1955.

Kennedy, J.B., Roberts, J., ILASS-Americas, 4th Annual Conference, Hartford, CT, May 1990.

Krulle, G., Mayer, W., and Schley, C.A., AIAA 90-1959, 1990.

Lin, S.P. and Woods, D.R., "Induced Atomization", ICLASS-91.

Mansour, A. and Chigier, N., Physics of Fluids A. 1991.

Mansour, A. and Chigier, N., Physics of Fluids A. Vol. 2, No. 5, pp. 706-719, 1990.

Miesse, C.C., Ind. Eng. Chem., Vol. 49, No. 9, 1955.

Ohnesorge, W.V., A.V.D.I., Vol. 81, No. 16, 1937.

Ranz, W.E., Can. J. Chem. Eng.. 36, 175 1958.

Rayleigh, Lord, Proc. Roy. Soc., 1878.

Reitz, R.D. and Bracco, F.V., Phys. Fluids. Vol. 25, No. 2, pp. 1730-1741, 1982.

Reitz, R.D., "Atomization and Other Breakup Regimes of a Liquid Jet", Ph.D. Thesis, Princeton

University, 1978.

Rupe, T.H., JPL Tech Report n. 32, 207, 1962.

Sattelmayer, T. and Wittig, S., ASME 86-GT-150, 1986.

Scheubel, F.N., NACA Tech. Memo No. 644, 1931.

Schweitzer, P.H., J. Appl. Phvsics. Vol. 8, 1937.

Squire, H.B., Brit. J. Appl. Phvs. . pp. 167-169, 1953.

Weber, C, AZMIM, Vol. 11, 1931.

15

ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper B

Keynote Lecture

EXPERIMENTAL AND THEORETICAL STUDIES ON THE STRUCTURE

OF FUEL SPRAYS IN DIESEL ENGINES

H. Hiroyasu

Department of Mechanical Engineering

University of Hiroshima

Hiroshima, Japan

ABSTRACT

The combustion process in a diesel engine is very complex, and its detailed mechanisms are not well

understood. The study of diesel engine emissions and fuel economy requires a basic understanding of the processes

of mixture formation, because the diesel combustion is strongly controlled by a fuel spray injected into the

combustion chamber.

This paper reviews the experimental results and the spray formation modeling of the diesel fuel sprays. The

experimental results for break-up length, spray angle, spray tip penetration and drop size distribution of the diesel

sprays are introduced to discuss the internal structure of the spray. A review is also made of the potentials of

specific models for atomization of the spray-jet.

INTRODUCTION

The high thermal efficiency on the diesel engine, especially the direct injection diesel engine, is mainly

caused by the relatively high compression ratio, the lower pumping loss as a result of the absence of the

throttle valve and overall lean mixture required to achieve an efficient heterogeneous combustion process. However,

diesel combustion processes are very complex and their detailed mechanisms are not well understood.

The primary factor which controls the diesel combustion is the mixture formation as shown in Fig. 1. The

mixture formation is controlled by the characteristics of the injection system, the nature of air swirl and turbulence

in the cylinder, and spray characterization.

Combustion in a direct injection diesel engine is closely related to the transient injection of a fuel spray into

air at high pressure and temperature. The air-fuel mixing processes, including atomization and vaporization, at

the edges of the spray are followed by the ignition of these regions and progressive diffusive burning of the

remainder of the spray. Therefore, accurate descriptions of the behavior of atomizing and vaporizing fuel in the

spray are important for any kind of models to comprehend and describe the processes of air-fuel mixing and

subsequent combustion.

The fuel injected from the nozzle into the combustion chamber of the diesel engine disintegrates into numerous drops of different sizes and concentrations in the spray. Figure 2 shows the main parameters the express

the aspect of a diesel spray. The motion of the spray tip and break-up length give clues as to the understanding of

the disintegrating process of a fuel jet. Spray angle and drop sizes distribution are the results of this

disintegrating process. These four parameters of the spray are related to each other in the disintegrating process

and affect the mixture formation through aerodynamic and thermodynamic processes in the combustion chamber

of the diesel engine.

17
Spray Characteristics

Spray Angle Break-up

Length

Drop Olstrlbullon Spray Tip Penetration Spray Angle Vaporizing Characteristics

Characteristics

of Ignition Delay

Spray Tip

Penetration

Flame Propagation

Partially Pro-mixed Combustion

Partially Diffusion Comlxistlon

Heat Losses

Diffusion of Combustion Products

NOx

Particulate HC

Rate of Heat Release

Fig. 1 Block diagram of diesel combustion Fig. 2 Parameters of a spray

EXPERIMENTAL STUDIES

Break-up Length

The injected liquid does not break-up instantly after injection. There is some unbroken portion which is referred to as the liquid break-up length. Break-up lengths average between 10 to 30 mm and appear to be present

for injection velocities above 200 m/s.

In a high speed small direct injection diesel engine of a light duty vehicle, break-up length of the fuel spray

and the distance from the nozzle to the combustion chamber wall are almost the same. Thus, the break-up

length may have a great effect on the spatial distribution of the liquid fuel and the formation of the fuel-air

mixture in that small engine.

Extensive experimental studies on the break-up phenomena of low velocity jets have been carried out (^"^).

Figure 3 shows break-up length which describes the continuous portion of the liquid column or unbroken

length as a function of the jet velocity. In this paper, the main interest is in the spray region. Break-up length

of the spray region was measured by electric resistance between the nozzle and a fine wire net detector located

in a spray jet ^^'^^

.

Figure 4 shows the effect of injection velocity on break-up length at various ambient pressures. Increasing

ambient pressure from 0.1 MPa to 3.0 MPa, the break-up length decreases. From 20 to 60 m/s of injection

velocity, the injection liquid column is disintegrated by the micro-turbulence of the liquid surface and is called

wavy flow^\ In this region, break-up length increases with an increase in injection velocity. When injection

velocity is further increased, the wavy flow is not observed and break-up length decreases. This region is called a

spray. Further increasing injection velocity, the break-up length reaches an almost constant value.

There is an apparent difference between the spray of which break-up length has decreasing tendency with an

increase of the injection velocity and one of which break-up length is hardly changed by the further increase of

the velocity. Categorizing sprays by their break-up length and related atomization, they are divided into two

regions; incomplete and complete sprays(^) . In the incomplete spray region, the liquid jet was ejected from the

nozzle to form a spray through a relatively slow disintegrating process. In other words, the process of

transformation from liquid column to drops had to be developed along the liquid jet and fine drops were only made

at a distance far from the nozzle exit. As injection velocity increased further, the disintegrating process from a

18
Dripping

Smooth Jet Region

Transition Flow Region

Wavy Jet Region

Spray

Injection Velocity Vj

Fig. 3 Break-up behavior of a liquid jet

AOi 1

Fig. 4 Break-up length vs. injection velocity Fig. 5 The effect of nozzle diameter on break-up length

liquid column to a fine spray appeared as soon as the liquid was injected. This region is called the complete spray

that is corresponded to a diesel spray. Spray as the general meaning, covers both the region mentioned above.

The effect of nozzle diameter on break-up length in a complete spray region is shown in Fig.5. Break-up

length in this region increases with an increase in the diameter of the nozzle. Break-up length is also affected by

the dimensionless length of a nozzle, that is, the ratio of the length L to the diameter D of the nozzle hole, L/D.

The following expression was derived for break-up length from experimental data which covered a wide range of

conditions.

U = 7.0 D ( 1 + 0.4 -L ) ( _Pa_ )0.o5

( _L )0.>3 ( _PL )0,5 D p,V? D Pa

This expression is valid for any complete spray region.

Spray Angle

Many photographs of spray were taken to measure the spray angle. Figure 6 shows the effect of injection

pressure on the spray angle. The spray angle increases with an increase in injection velocity, takes the

maximum value and reaches an almost constant value in a complcic spray region. The expcrical equation for the

19
20 40

Injection Pressure

60 80

MPa

100

Fig. 6 The effect of injection pressure on the spray angle

20

<^ 15

CD

a, 10

c

<

2

to ^

No22le-1

f- L/D=A

Pa=3.0-

Sd==0.1 mm

=0.2 mm

iD==0.3 mm

Lb -

50

I

AO J?

30 f,

c

20 -Ja

10 I

S

0 m

10 10^ 10

Injection Velocity V; m/s

Fig. 7 The spray angle and break-up length

spray angle when this angle reaches the constant value is expressed by the following:

L.

D

e = 83.5 ( )'°-^^

( D ^o•l5 Do( pi

0.26

(2)

Figure 7 shows spray angle and the break-up length versus injection velocity for several different nozzle

hole diameters. This figure is very important to discuss the internal structure of the spray. When the liquid was injected with the speed less than 70 m/s, the break-up length was not shortened and only a narrow spray

angle appeared even if the ambient pressure was elevated up to 3 MPa. However, the liquid injected with the speed

of over 100 m/s, was disintegrated quickly to form a complete spray. As the result of it, the break-up length was

shortened and the wide spray angle was achieved.

The region where the injection velocity ranged from 70 to 100 m/s, corresponded to an incomplete spray. The

maximum spray angle appeared at the injection velocity where the feature of a spray shifted from an incomplete

spray type to a complete one. In Figure 6, the injection pressure where the maximum spray angle occurs

increases with an increase in the liquid kinematic viscosity; this indicates that the transition velocity also

increases with the viscosity. However, Figure 7 suggests that the transition velocity is almost independent of the

nozzle hole diameter.

Spray Tip Penetration

The speed and extent to which the fuel spray penetrates in the combustion chamber has an important influence on air utilization and fuel-air mixing rate. So, spray tip penetration of the intermittent diesel spray was measured

with the aid of photographic techniques at various ambient temperature, ambient pressures, injection pressures and

injection durations.

Figure 8 shows the effect of injection pressure on spray tip penetration under constant ambient pressure and

injection duration(^). The logarithmic penetration on the ordinate is correlated with the logarithmic time from

the injection start on the abscissa. This figure shows the linear relationship with two different slopes between

logarithmic expressions of penetration and time. At the early stage of the spray, the slope is 1, but after a short

period, the slope changes to 0.5. This result shows that the spray velocity at the initiation of the injection is

constant and then the spray develops into a steady jet. The spray tip penetration decreases with an increase in

ambient gas pressure. The time at the intersection of the two lines decreases with an increase in injection pressure.

20
300

.03 05.07. .3 .5 .7 1 3 5 7 10 Time t ms

Fig. 8 Spray tip penetration at various injection pressures

The following expressions for the spray tip penetration are derived from the data obtained during our

investigation^^^).

0 < t < tu

s = 0.39 ( f t Pi (3)

t>tv

S = 2.95 ( ^ ( D • t f

pa

(4)

= 28.65 21—^5

( Pa AP )-

Results calculated from these equations are also expressed by the solid lines in Figs. 8.

Drop Size Distribution

The various methods employed in drop size measurement may be grouped conveniently into two broad

categories; mechanical and optical. The second category includes optical systems that have been developed in recent years and are finding an increasing range of applications. In mechanical method, there are liquid

immersion sampling technique moltcn-wax method ^'^ ) and frozen-drop technique ). In optical method,

there are several techniques that have been applied to drop sizing in dicsel fuel sprays; photography and

holography imaging technique('''\ line-of-sight extinction technique Fraunhofcr diffraction technique \

and the phase Doppler technique The effects of ambient pressure and injection pressure were measured using a diesel nozzle with a diameter of 0.3 mm, as shown in Fig. 9 by using Fraunhofer diffraction method.

21
150 e

3.

5100

1

eo

(5co

3

o

20 AO 60

Injection Pressure Pj

9 /

" J./X1U *I.AX 10 m /s

3.7x10- 27.

1 Diesel Nozzle

D=0.3 mm

Pa=3.0 MPq

4xK)'

1

— r ^

.61.1x10

. £i

6

I A ^ £k 2 —a , o

Injection Pressure Pj MPa

Fig. 9 The effect of ambient pressure and injection Fig. 10

pressure on the Sauter mean diameter

The effect of viscosity and injection

pressure on the Sauter mean diameter

The effect of liquid viscosity on the Sauter mean diameter was measured using glycerin-water solutions of

various kinematic viscosity. Figure 10 shows the measured results of the relationships between the solution's

viscosity and the Sauter mean diameter Increasing injection pressure, the Sauter mean diameter decreases,

but there is still some difference in the Sauter mean diameter according to the viscosity of the fuel. When the

viscosity of the fuel is over a certain value, the Sauter mean diameter increases rapidly.

Dimensionless analysis was conducted based on the above measurement results, leading to the following

experimental equations for the Sauter mean diameter^^^).

X32 = MAX [

L S X 3 2

J D (6)

-AijJl = 4.12 Re°''- We-°'' ( ( D l^a pa (6 -a)

= 0.38 Re We -''^

• ( ( £L Y'-''' D M^a pa (6-b)

where MAX [ ] denote the lager value of the two.

Equation (6-a) is the Sauter mean diameter for an incomplete spray with low injection velocity, and Eq. (6-b)

for a complete spray.

Incomplete and Complete Sprays

Overall feature of the break-up length, spray angle and mean diameter are summarized into Fig. 11. The breakup mechanism of a complete spray is different from that of an incomplete spray. Then, the empirical equation of

the Sauter mean diameter for a complete spray is expressed by the different form from that for an incomplete

spray as mentioned before. The maximum spray angle was obtained at the transition velocity from incomplete

spray to complete one. Further observation of the internal flow in the nozzle shows the cavitation area expanded to the exit of the

nozzle when the liquid jet was injected at the transition velocity. The highly turbulent internal flow due to the

cavitation caused the wide spray angle and sometimes forced to shorten the break-up length as shown in Fig. 7.

22
atc

3

I O

m

c

<

ow

Q.

atE

gb

c

a

a> S

. Lb: Break-up Length

\ 1 \ 1

1

1

1

1

1 •

1

1 Incomplete Spray

0 ' Complete

1 1 Spray

i

'

1

1

1 Injection Velocity

Fig. 11 Behavior of diesel spray

NUMERICAL MODELS

Multidimensional models of diesel engine combustion account for the temporal and spatial variations of

the flow field, temperature, composition, pressure, and turbulence within the combustion chamber. Most of

the multidimensional models that have appeared in the literature are based on the following spray equation ^^^\

which depend on time and on the radius, velocity, and temperature of the drops and may account for thick spray

effects [ drop collisions, coalescence, and volumetric displacement of the gas phase ] and for drop breakup.

A statistical description of the spray may be given by the distribution function [or density function ]

fj ( r, X, v, t ) dr dx dv

(7)

which is the probable number of particles of chemical composition j in the radius range dr about r located in the

spatial range dx about x with velocities in the range dv about v at time t. The spray equation can be written as

at 'at

(Rjfj)-Vx(V-fj)-Vv(F,fj) +Qj + rj

(8)

where v, F, R, and T denote the velocity, force, rate of change of the radius, and temperature of the drops,

respectively. Q represents the rate of increase of f due to nucleation or liquid breakup, and r will represent the rate of change of f caused by collisions and coalescence.

Main approach which has been used to analyze the flow field in diesel engines is Lagrangian-Eulerian

formulations, where Lagrangian equations are employed for groups of drops and Eulerian equations are employed for the gas phase. Lagrangian-Eulerian formulations can also account for thick spray effects such as drop collision, coalescence, and breakup.

23
The conservation laws for the gas, mass, momentum, internal energy and the scalar variables that characterize

turbulence ( k and e ) are expressed in the form of time-dependent, partial differential equations. These

equations were solved subjected to appropriate initial and boundary conditions using an explicit numerical

method. For computational efficiency, stochastic parcel injection was used for the spray drops. In this technique

each computational parcel represents a group of drops with similar physical attributes and the attributes are assigned using a Monte Carlo sampling method.

The spray equation is coupled with statistically averaged gas phase equations of mass, momentum, and

energy. However, there are many difficulties in treatments of the spray initial and boundary conditions.

One of the main difficulties in spray modeling has been the specification of drop size at the nozzle, since the

atomization process is not resolved in the computations and it is not understood. Experiments are difficult

since the spray near the nozzle is optically dense and drop size cannot be measured, except near the spray

edge. For this reason, most measurements have been confined to locations far downstream of the nozzle.

Some model may account for thick spray effect, that is, drop collisions, coalescence and volumetric

displacement of the gas phase and far drop breakup, but some model did not account for thick spray effects.

Thin spray Models

In thin or dilute spray models volumetric displacement effects can be neglected and the interactions between

individual drops can be disregarded or highly simplified.

Nguyen et al. used a thin spray approximation, that is, volumetric displacement effects on the gas phase

equation were neglected, and also drop collisions and coalescence were neglected. The numerical calculations were compared with the experimental data of Hiroyasu and Kadota and good agreement was obtained. Figure

12 shows a comparison between the predicted and measured penetration of sprays injected into stagnant

combustion chambers at different pressures.

Thick Spray Models

Thick or dense spray models that account for volumetric displacement effects seem to have originated with

Dukowicz's work (23)^ where a Lagrangian-Eulcrian formulation was used to compute the behavior of atomized,

non-evaporating liquid sprays injected into a gaseous environment. He accounted for volumetric

displacement effects but did not consider drop collisions, coalescence and breakup, and obtained good

agreement with the experimental date of Hiroyasu and Kadota ^^^^ for a variety of ambient conditions. This

agreement seems to have been obtained by employing small drops at the nozzle exit. Cloutman et al.^^''^ and Butler et al.^^^^ used the stochastic model proposed by Dukowicz law of

the wall to account for the turbulent boundary layers and heat transfer to walls, and a mean-reaction-rate model

in DI diesel engines. The calculation were performed with the CONCHAS-SPRAY code^^^).

Fig. 12 Comparison between predicted (solid line) and measurement (symbols)

spray tip penetration at several chamber pressure^^'^

24
Amsden et al/^^) developed the KIVA code to analyze two- and three dimensional flows in DI engines. The

KIVA code uses an ALE (an arbitrary Lagrangian-Eulerian) formulation, a subgrid- scale model and a meanreaction-rate model.

O'Rourke and Amsden^^^^ analyzed the spray penetration, vaporization, mixing and combustion in the

UPS-292 stratified charge engine with the KIVA code and showed that combustion occurs in the wake of the

spark plug.

Takenaka etal.^^^^ performed three-dimensional calculation in a Dl diesel engine with modified KIVA code.

They calculated the gas flow, spray pattern, air-fuel mixing process for the heavy duty DI diesel engine. The

results of computation were compared with the experimental data of pressure, flame temperature and mass

change of chemical species in cylinder (3^) and good agreement was obtained.

In 1987, KIVA-II computer program was reported by Amsden eta\P^\ KIVA-II code improves the earlier

version in the accuracy and efficiency of the computational procedure, the accuracy of the physics submodels,

and in versatility and ease of use. This model also includes a two-equation ( k- e ) model of turbulence, a

stochastic spray model, an implicit-continuous Eulerian (ICE) method with conjugate residual iteration, a quasisecond-order-accurate convection scheme, and several other improvements over the original KIVA code.

Figure 13 shows a comparison between the predicted and measured pressure diagram in the UPS-292

stratified charge engine and indicate that the KIVA-II code accurately predicts the cylinder pressure if the correct compression ratio is used in the calculation.

5000

4000

oa

i JOOO

2000

1000

-30" -20 -10 0 10 20 30 40 50 €0

Cronh Angle - degrees

Fig. 13 Measured and predicted pressure diagrams

in stratified charge engine^^')

Small Drop Models

Bracco's group of Princeton University were used very small drops in initial drop distribution near the nozzle

exit. Kuo and Bracco(32) and O'Rourke and Bracco^") used the thick spray model to calculate transient

axisymmetric non-evaporating spray, asses the importance of drop collisions and coalescence, and compute the

spray tip penetration and drop size distribution in spray injected into high pressure chambers. O'Rourke and

Bracco (33) used the correlations proposed by Reitz and Bracco^^^) for the spray cone angle and initial mean drop size,

that is, initial drop size is average radius is about 3 |im. Figure 14 shows the computed and experimentally

measured drop volume distributions. The computed penetration versus time curves are plotted in Fig, 15, along

with the experimentally measured curve. These curves show the expected trend that as the drop size is increased,

spray tip penetration is enhanced because the larger drops exchange their momentum with the gas more slowly.

The penetration is somewhat insensitive to change in the initial drop size, however, and only the case with

initial Sauter mean drops of 21 \im diverges markedly from the experimental curve.

25
TUT —txramtm ( ti^wMl

-X- uwHcs oonmunoN CMCUUnON ltMI>V"t —o-comnto oowNsncM

Fig. 14 Measured and predicted drop volume

distributions^-'^^

cm

Fig. IS Effect of varying injected drop size on computed tip penetration*^^)

Blobs Models

The assumption of very small size in initial drop condition near the nozzle exit is unrealistic. Reitz and

Diwakar^^^' injected drops having a diameter equal the nozzle exit diameter, and they assumed that the

dynamics and breakup of these large drops are similar to those of a liquid jet as illustrated conceptually in Fig.

16^^^\ Several kinds of blobs models are proposed. Breakup of the blobs is modeled using different stability

analysis by different researchers. Reiz and Diwakar^^^^ incorporated into KIVA a model for both bag and

boundary layer stripping breakup. In their model, bag breakup occurs if the drop Weber number. We, was greater

than a critical value of 6. An unstable lifetime, tj is expressed following equation. Boundary layer stripping

(BLS) is due to flow over the drop inducing a thin laminar boundary on the drop surface. The boundary layer is assumed to be stripped from the periphery of the drop as shown in Fig. 17. Boundary layer stripping is proposed to occur when the Weber to jet Reynolds number ratio We / V Re is grater than 0.5. For stripping, an

unstable lifetime t2 is expressed by equation (12),

For Bag Breakup

We > 6.0 (9)

t, = ci [ .eiii ]i

2 a (10)

For Boundary Layer Stripping

WcZ/rT > 0.5 (11)

t2 = C2 ( -2^- )2 W p. (12)

These large ' blobs ' fragment dynamically according to the rate equation

d r = - ( r ' Tsuble ) dt ^ (13)

26
The term ' r^^ie ' is found by equating the applicable drop stability criteria to the Weber number. We, for bag

breakup or We/ VRe for stripping. The characteristic time, t, is also found tj (eqn. 4) or t2 (eqn. 5) depending on the appropriate breakup mechanism. Figure 18 shows the variation of drop size with axial distance from the

injector for the solid-cone spray.

Flow

lOO

Fig. 17 Velocity profile created by shear forces

during boundary layer stripping

Fig. 16 Conceptual picture of the spray jet atomization

E

en

-r

20 40 60 80 100 120 Distance from Injector (mm)

Fig. 18 Computed far-field solid-cone spray drop Sauicr mean radius variation

with distance from the nozzle. Three data points at 65mm measured

by Hiroyasu and Kadota^^^^

27
Recently, Reitz has taken a different approach to jet breakup by considering Kelvin-Helmholtz instability

growth on the jet surface and using the fastest growing wavelength as the length scale for drop formation of the

parcel "blob". Thus, the jet "blobs" can breakup immediately upon exit from the nozzle by shedding small

drops of diameter. The parcel "blobs" decrease linearly in size from their initial size at the nozzle diameter by

this mass loss due to drop stripping relation.

O'Rourke and Amsden's model is based on an analogy suggested by Taylor^'*^^ between an oscillating

and distorting drop and a spring-mass system. The restoring force of the spring is analogous to the surface

tension force. The external force on the mass is analogous to the gas aerodynamic force. They also consider

the damping forces due to liquid viscosity. They called this model the TAB (Taylor Analogy Breakup) model.

To predict the drop sizes after breakup, they used an equation motivated by an analysis based on energy

conservation. In this analysis, they equated the energy of the parent drop before breakup to the combined energies

of the product drops after breakup. Before breakup, the energy of the parent drop in its own frame of reference is

the sum of its minimum surface energy and the energy in oscillation and distortion.

University of Wisconsin group developed a theoretical model on the basis of applying the unstable wave

penetration mechanism for a Rayleigh-Taylor instability as shown in Fig. 19 and boundary layer stripping (BLS)

to a wide range of Weber number.

In this model, only the large drop breakup mechanism is considered initially. Drop mass loss due to KelvinHelmholz instability and the boundary layer stripping mechanisms have been neglected initially. This model is used to predict the transient drop breakup phases and obtain a correlation of the fragment size with time.

Typical results for single drop breakup are shown in Fig. 20 for a range of drop Weber numbers. The slight

decrease in diameter with time is due to the boundary layer stripping mechanism mass loss and the catastrophic

breakup at a given time is due to the Rayleigh-Taylor instability.

Spray Impingement on the Wall

Fuel spray may impinge on the piston bowl in DI diesel engines and may have a great influence on the

mixture formation and combustion processes. Naber and Reitz showed that their wall-jet model yields better

agreement with experimental data of Kuniyoshi et al. Their results indicate that the width and high of the

wall jet decrease as the gas density and the wall inclination angle are increased.

(b) < ^ P°

2

Fig. 19 Rayleigh-Taylor instability droplet breakup conceptual model

28
0.0 t.O Z.e 3.0 4.0 t.O (.0

Olmenslonless Time

Fig. 20 Liquid droplet transient fragment size

CONCLUSIONS

The combustion process in a diesel engine is very complex, and its detailed mechanisms are not well

understood. The study of diesel engine emissions and fuel economy requires a basic understanding of the

processes of mixture formation, because the diesel combustion is strongly controlled by a fuel spray injected into

the combustion chamber. More experimental data are required to validate the predictions of multidimensional

models. Especially needed is a better understanding of spray breakup, include drop size distributions and drop

velocities in engines, and they can be used to determine the effects of turbulence and drop collisions, coalescence,

and breakup on the engine flow field and combustion.

NOMENCLATURE

^1,2 constants V drop velocity (m/s)

D nozzle diameter (m) \ initial spray velocity (m/s)

Do sac chamber diameter of nozzle (m) W relative velocity (m/s)

F force (N) We Weber number (m/s)

L nozzle length (m) X position

Lb break-up length (m) ^^32 Sauter mean diameter (m)

P. ambient pressure (MPa) , LS ^32 Sauter mean diameter

AP (m) difference between injection pressure at incomplete spray

and ambient pressure (Pa) ^32 Sauter mean diameter

Q rate of increase of distribution function at complete spray (m)

R rate of change of radius viscosity of air (pa • s)

r round radius of nozzle (m) viscosity of fuel (pa • s)

, drop radius (m) p. density of air (kg/m^)

Re Reynolds number Pi density of fuel (kg/m^)

S Spray Tip Penetration (m) 6 spray angle (degree)

T temperature (°C) c surface tension (N/m)

t time (s) r rate of change of distribution function

h unstable lifetime for BLS T liquid breakup time (s)

unstable lifetime for stripping

break-up time (s)

29
REFERENCES

1. DeJuhasz, K. J., "Dispersion of Sprays in solid Injection Oil Engines", Trans, of ASME, 53, p.65, 1931.

2. Castleman, R. A., "Mechanism of Atomization Accompanying Solid Injection", NACA Kept., No.40, 1932.

3. Tanasawa, Y. and Toyoda, S., "A Study on the Atomization of High Speed Liquid Jets", Bull, of JSME 20-

92,1954.

4. Hiroyasu, H., Shimizu, M. and Arai, M., "The Break-up of High Speed Jet in a High Pressure Gaseous

Atmosphere", 2nd Int. Conf. on Liquid Atomization and Spray Systems, ICLASS-82, p.69, 1982.

5. Shimizu, M., Arai, M. and Hiroyasu, H., "The measurements of Break-up Length in High Speed Jet", Trans,

of JSME 49-446,1983.

6. Arai, M., Tabata, M., Hiroyasu, H. and Shimizu, M., " Disintegrating Process and Spray Characterization of

Fuel Jet Injected by a Diesel Nozzle", SAE Paper No.840275, 1984.

7. Tanasawa, Y., "High Speed Diesel Engines I, Fuel Spray Systems and Spray Combustion, Sankaido, p.87,

1956.

8. Hiroyasu, H. and Arai, M., "Structures of Fuel Sprays in Diesel Engines", SAE Paper, No.900475, 1990.

9. Hiroyasu, H., Kadota, T. and Tasaka, S., "Penetration of Diesel Sprays", Trans, of JSME, 44-385, p.3208,

1978.

10. Hiroyasu, H. and Arai, M., "Fuel Spray Penetration and Spray Angle on Diesel Engines", Trans, of JSAE,

21, p.5, 1980.

11. Hiroyasu, H., Kadota, T., "Fuel Droplet Size Distribution in Diesel Combustion Chamber", Bulletin of

JSME, 19-135, p.1064, 1976.

12. Tanasawa, Y. and Hiroyasu, H., "Measurement of Size Distribution of Sprayed Drops by Means of Molten

Wax", Trans, of JSME, 26-162, pp.224-232, 1960.

13. Karasawa, T., Kurabayashi, T. and Kaeriyama, H., "Freezing Method by Liquid Nitrogen for Measuring

Spray Droplet Sizes", Journal of Fuel Society of Japan, 62-679, pp.90-912, 1983.

14. Nishida, K., Murakami, N. and Hiroyuki, H., "A Pulsed Laser Holography Study of the Evaporating Diesel

Spray in a High Pressure Bomb", Trans, of JSME, 52-478, p.2457, 1986.

15. Kamimoto, T., Yokota, H. and Kobayashi, H., "A New Technique for the Measurement of Sauter Mean

Diameter of Droplets in Unsteady Dense Sprays", SAE Paper, No.890316, 1989.

16. Arai, M., Hiroyasu, H. and Kishi, T, "Measurement of Drop Size Distfibution of the Spray by Using a Laser Diffraction Method", Trans, of JSME, 29-252, p.l795, 1986.

17. Tabata, M., Arai, M., Hiroyasu, H., "Atomization of High Viscosity Liquid by a Diesel Nozzle", Bull, of

JSME, 29-252, p.l795, 1986.

18. Bhatia, J. C. and Durst, R, "A Comparative Study of Some Probability Distributions Applied to Liquid

Sprays", Proc. Int. Conf. Mech. of Two-Phase Flows, Taiwan, pp.279-284, 1989.

19. Hiroyasu, H. and Arai, M. and Tabata, M., "Empirical Equations for the Sauter Mean Diameter of a Diesel

Spray", SAE Paper, No.890464, p.97, 1989.

20. Williams, F. A., "Process in Spray-Combustion Analysis", Eighth Int. Symp. on Comb., Williams &

Wilkins Co., Baltimore, 50-69, 1962.

21. Nguyen, H. L., Schock, H. J., Carpenter, M. H., Ramos, J. I. and Stegeman, J. D., "Numerical

Simulation of the Flow Field and Fuel Sprays in an IC Engine", SAE Paper No.870599, 1987.

22. Hiroyasu, H., Kadota, T., "Fuel Droplet Size Distribution in Diesel Combustion Chamber", SAE Paper

No.740715, 1974.

23. Dukowicz, J. K., "A Particle-Fluid Numerical Model for Liquid Sprays", Journal of Computation Physics,

Vol.35, pp.229-253, 1980.

24. Cloutman, L. D., Dukowicz, J. K. and Ramshaw, J. D., "Numerical Simulation of Reactive Flow in

Internal Combustion Engines", Lecture Notes in Physics, Vol. 141, pp.1 19-124, 1981.

25. Butler, T. D., Cloutman, L. D., Dukowicz, J. K. and Ramshaw, J. D., "Multidimensional Numerical

Simulation of Reactive Flow in Internal Combustion Engine", Progress in Energy and Combustion Science,

VoL 7, pp.293-315, 1981.

30
26. Cloutman, L. D., Dukowicz, J. K. and Ramshaw, J. D. and Amsden, A. A., "CONCHAS-SPRAY : A

Computer Code for Reactive Flows with Fuel Sprays", Report No. LA-9294-MS, Los Alamos National

Laboratory, 1982.

27. Amsden, A. A., Butler, T. D., O'Rouke, R J. and Ramshaw, J. D., "KIVA-Comprehensive Model for 2-D

and 3-D Engine Simulations", SAE Paper No.850554, 1985.

28. O'Rouke, P. J., Amsden, A. A., "Three Dimensional Numerical Simulations of the UPS-292 Stratified

Charged Engine", SAE Paper No.870597, 1985.

29. Takenaka, Y., Aoyagi, Y, Tsuji, Y. and Joko, I., "3D Numerical Simulation of Fuel Injection and

and Combustion Phenomena in DI Diesel Engines, SAE Paper No.890668, 1989.

30. Aoyagi, Y., Kamimoto, T., Matsui, Y. and Matsuoka, S., "A Gas Sampling Study on the Formation Process

of Soot and NO in a DI Diesel Engine", SAE Paper No.800254, 1980.

31., Amsden, A. A., Butler, T. D. and O'Rouke, P. J., "The KIVA-Computer Program for Transient

Multidimensional Chemically Reactive Flows with Sprays", SAE Paper No.872072,1987.

32. Kuo, T. W. and Bracco, F. V., "Computation of Drop Sizes in Pulsating Sprays and of Liquid-Cone Length

in Vaporizing Sprays", SAE Paper No.820133, 1982.

33. O'Rouke, P. J. and Bracco, F. V., "Modeling of Drop Interactions in Thick Sprays and a Comparison with

Experiments, Stratified Charge Automotive Engines", pp.1054-1064, 1979.

34. Reitz, R. D. and Bracco, F. V., "Mechanism of Atomization of a Liquid Jet", Physic of Fluids, Vol.22,

No.6, pp.1054-1064, 1979.

35. Reitz, R. D. and Diwakar, R., "Effect of Drop Breakup on Fuel Sprays", SAE Paper No.860469, 1986.

36. Reitz, R. D. and Diwakar, R., "Structure of High-Pressure Fuel Sprays", SAE Paper No.870598, 1987.

37. Bower, G., Chang, S. K., Corradini, M. L., EI-Beshbeeshy, M., Martin, J. K. and Krueger, J., "Physical

Mechanisms for Atomization of a Jet Spray : A Comparison of Models and Experiments:, SAE Paper

No.872089, 1987.

38. Reitz, R. D, "Modeling Atomization Process in High-Pressure Vaporizing Sprays", Atomization and Spray

Technology 3, pp.309-337, 1987.

39. O'Rouke, P. J. and Amsden, A. A., "The TBA Method for Numerical Calculation of Spray Droplet

Breakup", SAE Paper No.83 1318, 1983.

40. Taylor, G. I., " The Shape and Acceleration of a Drop in a High Speed Air Stream", The Scientific Papers of

G. I. Taylor, ed. G. K. Batchelor, Vol.11, Uviv. Press, Cambridge, 1963.

41. Naber, J. D. and Reitz, R. D, "Modeling Engine Spray / Wall Impingement", SAE Paper No.880107, 1988.

42. Kuniyoshi, H., Yamamoto, H., Fujimoto, H. and Sato, G. T., "Investigation of the Characteristics of Diesel

Spray (Third Paper : Impinging upon a Flat Plate)", Journal of JSME, Vol.15, pp.57-64, 1980.

31

ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper C

Keynote Lecture

ATOMISATION OF LIQUID METALS FOR THE MANUFACTURE OF

METAL POWDER AND FOR SPRAY FORMING

K. Bauckhage

University of Bremen, Verfahrenstechnik

Bremen, Germany

ABSTRACT

Among the numerous atomization techniques for melt spraying the most important ones

will be described and discussed in regard to their design criteria and operating conditions. They

can be classified according to the forces acting on the melt stream via the high energy fluid

streams (gas or water jets), by ultrasound activated vibrations, by impact or by rotating elements.

The equations for overall drop size distributions will be discussed. While this information

is important for an optimized production of powders, there is a strong demand for a more differentiated description of local drop size and drop velocity distributions which can be used for

the balances of mass, momentum, and energy during spray forming or metal spray coating.

Using the measurements of local velocity and size distributions of the melt droplets in cor- respondence with the differential equations of the a.m. balances gives a basis for a numerical

simulation of the gas atomized spray forming, the results of which are in good agreement with

the experimental findings.

INTRODUCTION

Atomization is a process in which a liquid stream is disintegrated into a large number of

droplets of various sizes. Because all solid state materials like minerals, ceramics, glass and of

course metals can be converted into the liquid state (it is only a question of heat and temperature) there have been developed numerous atomization techniques for molten materials which

differ greatly in the modes in which energy is supplied to the disintegration process and which

after comminuting the melt streams alter widely in the tasks and hence in the process lines of

accompanied unit operations. In many cases where melting is a necessary operation in the line to generate for instance metals of high purity, the atomization process in comparison with for instance milling represents the most economic way to produce particulate materials, i.e. metal

powders.

Atomization techniques for powder production and also for melt spraying have been

reviewed several times in the last decade, [1] to [3]. This has mainly been done from the

standpoint of metallurgical engineers, because of the economic and technical importance of

metal powder production. But this is not the only field of interest for melt spraying techniques

which are used mainly to produce the majority of available metal powders with a special metallurgical, crystalline or mechanical behaviour and with properties controlled by the individual

conditions of consolidation, rapid quenching, segregation or crystallization processes of the par33
ticulates and specified by their overall size distribution, the mean particle diameter, the particle

shape and the capacity to spill.

It is also the field of interest of mechanical, chemical and process engineers

to produce metallic layers with specific mechanical characteristics on tools or machine

elements, with electrical or optical characteristics in combination with surfaces of special

substrate materials;

to manufacture spray deposited preforms of near-net shape as a single operation in a new and very successful process line which combines the creation of spray and the consolidation of particulate materials, thus contributing to micro-structural refinement in

the billets, sheets or tube preforms.

These techniques are not restricted to molten metals. Ceramic and glass powders also are produced by melt atomization, mostly with spherical shapes. Atomization is not necessarily

limited to the production of droplets or powders. The high viscosity of molten minerals, glasses or ceramics for instance allows the generation of ultrafine fibres (rock wools and high temperature resistant fibrous refractory). And spray deposit preforms already include ceramic-fiber

reinforced metal- matrix composites.

MAIN CRITERIA OF MELT ATOMIZATION

The main difference between melt atomization and conventional atomization (of aqueous

and organic liquids) consists in the preceding energy transport for heating and melting the

material as well as in the subsequent release of this energy in order to consolidate the dispersed

droplets (and to convert them into powder particles) and to cool the particles after impact (and

coalescence) on the substrate or on the preform layer.

In most of the existing experimental rigs or production plants the main unit operations - melting, atomization, cooling and consolidation - are arranged below each other or integrated

into one common vessel following the same sequence. (Taken all together these unit operations

define the so called process of "pulverization".) Following gravitational forces the melt leaves the

induction heater and poures from the crucible/tundish to the disintegration zone of the

atomizer. Very often the basic operations have to be performed under an inert gas atmosphere

(such as nitrogen or argon) in order to prevent chemical reactions of the metal melt with for ex- ample oxygen. It should be noted that increasing safety hazards go hand in hand with the

production of finer powders of for instance aluminum and magnesium. Here one needs closed

plants including sieving and bottling, all together operating under inert gas.

Because of great differences in the thermal expansion of the various materials of the

crucible/tundish and the atomizer elements (ring nozzles, gas tubes, etc.) and considering the

enormous temperature differences due to standstill and stationary operating, this gives reason to keep the melting design separated from the atomizer design. Typical pressure atomizers (plain

orifices, simplex nozzles, etc.) or internal mixing atomizers, well known from conventional

atomization, cannot be used.

Conventional atomization is characterized by the liquid state of the aqueous or organic

fluids after atomization and therefrom by the tendency of the liquid droplets to coalesce or to evaporate thus causing the loss of the initial drop shape and size (as a consequence of impact,

deformation or heat transfer). In comparison to this the melt atomization in general provides for

its own analytical instrument, i.e. the powder product, which is extremely helpful for analysis

of the disintegration process. This "by-product" can be obtained as a representative powder

sample directly from the powder production process during atomization, or can be taken as momentary powder samples from the succeeding manufacturing processes if for instance the

tasks of generating metallic layers or spray deposited preforms have to be slightly manipulated.

After this powder "by-product" has been screened it functions as a prerequisite for an assess- ment of the disintegration process and of product quality. These powder samples supply a basis

for documentation not only of atomization efficiency but also for the quality of the total spraying system (including cooling and consolidation). Such an "alibi" cannot be obtained from conventional atomization.

34
Last but not least, an important difference of melt atomization in comparison with conventional atomization results from the material properties of the melts of metal, ceramic and

glass, i.e., surface tension, density and dynamic or kinematic viscosity, which are given in

Table 1 for some representative materials. Compared with water, the values of surface tension

rise to a 20-fold level, those of density to an 11 -fold one and those of dynamic viscosity to a

4-fold level; but the values of kinematic viscosity go down to a factor of 0.2. The ranges of

these properties differ greatly from those of aqueous or organic fluids, which show for the surface tension for instance a range of only 10 to 74 dyn/cm.

Table 1 Materials properties with T = temperature; O - surfache tensions; = dynamic viscosity; Pl = density; i^l = kinematic viscosity

Melt

1

ffl0-5[N/m)

1

tIlJ0J[P1Pl'0^( kg/in^ ] Aluminum 930

; 570 (865 2(3) 2.25 1 (U)

Cupper U60 1.150 (UOO) i2 (3.4) 8J (7.8) OJ

Iron 1,870 UOO (1,820) 43 (6.0) 7.2 (6.9) 0.6 (0.87)

Lead 600 400 2 10.7 0.2

Magaesium 923 570 1.2 U4 0.8

Pb-Bi-Sn-Eutecric 348 450 \J5 9.4 0.16

Steel 1,730 UOO 5 7i 0.66 Tm 500 560 1.6 6.6 024

Zinc 690 785 2 (3.6) 7.1 (6J) 028 (0.55)

Water 293 74 1 1 1

Glas 900 160 100 6.27 159

With regard to the task of this contribution one has to confine to the most common

atomizer types and to limit the selection of interesting atomizers. Such a limited selection is given in Fig. 1, which illustrates various principles of atomization of molten metals and melt

spraying. This overview can give only a faint impression of this wide field and of different

technical solutions, in which mainly two types have reached a wide spread industrial standard of

economic application, i.e., the gas or water atomization and the centrifugal atomization.

Fig. 1 Rotating and gas atomizers

35
DESCRIPTION OF THE MOST COMMON ATOMIZERS

A jjresentation and a systematic description of common atomizers for melt spraying

seems best to start with a classification of governing disintegration forces in regard to the

transport of momentum, i.e. a) the momentum transport by pressure gas or liquid streams

b) the introduction of centrifugal forces using rotating elements or via angular momentum

of the gas stream

c) the use of vibrating forces

d) momentum induced by the melt droplets themselves via impact.

It can also be useful to take advantage of combinations of these alternatives, i.e. a) and c) or b)

and d).

As has been described elsewhere in detail, [4] and [5], liquid atomization in general is characterized by typical formations of thin jets and sheets of liquid, which are disrupted into

ligaments and droplets by rapidly growing disturbances in or on the surface of these liquid

fragments. Primary jets and sheets as well as ligaments and fragments are caused by internal

forces of the liquid itself and by external ones due to tangential and normal forces of atomizing

gas jets (or vibrating gas columns) or due to centrifugal forces induced by rotating elements. All

the various types of internal and external forces have to overcome the consolidating influences

of surface tension, which is permanently contracting the different fragments to spherical shapes,

thus minimizing the surface energy. Conversely the liquid viscosity tends to stabilize the

momentary shape of liquid fragments (thus being the main cause for fiberformation during the

solidification process of the melt jets and ligaments). This means that liquid viscosity is opposing the disruption process but on the other hand - exerted by gas jets - gas viscosity may support the disruption. Inertial forces due to liquid or gas density promote the disruption process.

"Breakup occurs, when the magnitude of the disruption forces" commonly "just exceeds

the consolidating force of surface tension. Most of the larger drops produced in the initial disintegration process are unstable and undergo further disruption into smaller drops. Thus the final

range of drop sizes produced in a spray depends not only on the drop sizes produced in primary

atomization but also on the extent to which these drops are further disintegrated during secon- dary atomization." Reducing the different phenomena and also the very complicated and complex details of disintegration in this quotation in a pregnant manner to the main object,

Lefebvre [4] focused his description onto the conventional atomization of aqueous and organic

liquids. Nevertheless his description also holds for liquid melts and thus for melt atomization.

But though we principally may expect the same laws describing the disintegration processes of

melts as of aqueous or organic liquids, we may not assume the dimensionless correlations leading

to the same constants without having proved this. Relating to the wide variety of materials

properties (Pl> ^i) of molten metals some of the extreme large values give raise to prove the

range of validity of earlier relevant literature.

Gas- or Liauid-Atomization

Gas- or liquid-atomization techniques cover the main field in the manufacture of metallic particulate and spray deposited preforms. In recent years increasing numbers of reports

describe these atomization and melt-spraying techniques, see for instance [6] to [8] (also for

ceramics and glasses). When the atomization plants and their typical process design are considered, it can be said, that in most cases melting and atomization are integrated into the same

system, that both processes take place under the same pressure (i.e. Ap = 0); the latter mainly

because of the insufficient strength of the refractory material of the heated crucible/tundish or because of sealing problems. This causes an effluent velocity of the melt stream leaving the bottom exit of the tundish governed only by the level h of the liquid (melt) above the exit:

(1)

36
where h very often varies with time: h = h(t), thus making necessary a manipulation of

the factor 0 (by means of for instance a plugging device) in order to compensate the influence

of h(t) in the mass flow rate

M L = ^( -^^-^) Pl = Ml h) (2)

As the materials properties show strong temperature dependencies, these properties - depending on the task of atomization and of the cooling process - have to be fixed as constants near the melting temperature Tl, often necessarily taking into account a small temperature difference At depending on a process immanent undercooling or overheating.

In order to prevent the tundish exit or the atomizer from "freezing in" by a solidifying

melt stream or by impacting and also solidifying melt droplets one has primarily to avoid ex- treme low mass flow rates with insufficient heat transfer or on the other hand streaming conditions of the effluent melt jet that belong to zones of spontaneous disruption of free jets.

The main task is to bring the continuous melt stream into the disintegration zone where

the momentum of pressure gas or water jets (i.e. their amount and their direction) prevents

backsplashing of melt droplets.

There exist several solutions to this problem - one answer consists of the following

design principles for atomizers:

making use of the gravitational force for the effluent of the melt out of the

crucible/tundish,

leading the melt flow as a continuous central stream downwards into the atomization

zone,

surrounding it with concentric orifices for the atomization pressure gas or liquid,

using co-current flows of melt and gas (downwards) but inclining the gas jets under

small angles in order to rise tangential and normal forces.

Fig. 2 Various types of gas atomizers

It is of prime importance to manage the flow conditions of the melt stream and of the

essential energy supply within the melt stream itself (by convection and conduction) as well as through its conducting tundish walls in order to keep the melt liquid at the exit of the tundish

as well as down to the atomization zone. On the other hand one has to provide for the necessary

insulation against the cooling effect of the gas tubes in the vicinity of the exit and against the

expanding and therefrom cooling gas streams near the melt stream in the atomizer zone. Another design element to solve the problem is to limit the liquid level h of the melt

above the bottom exit of the crucible/tundish by means of an overflow. This prevents the melt

jet from breaking up spontaneously or from generating single drops (caused by interaction with

the surrounding gas) before it has reached the disintegration zone. Keeping the liquid level constant or the velocity Vq provides a constant mass flow rate ICIl and hence a stationary melt flow

and an unchanged turbulent situation of the melt jet in the atomization zone. A third method is, not to allow the melt stream to accelerate under free fall conditions

and to prevent it from generating drops by using a short tube, which leads from the tundish

exit to the atomization zone. (This is the so called closed coupled atomizer.) Using such a short

37
tube means that additional drag deminishes the melt velocity in the disintegration zone, but

on the other side to allow higher mass flow rates. Further effective design principles to avoid "freezing in" result in leading the gas stream

into a spiral form onto the melt, thus forcing the atomized droplets immediately to twist away

from the atomizer, as already shown in Fig. 1 after [8]. Or they find their expression in a twostage atomization, in which the upper (first) gas stream only takes over the task of minimizing

the relative velocity of melt and gas, of raising the drag forces of eventually back-splashing

droplets and of transporting them downstream into the spray cone. Whereas here the upper gas

stream does not serve at all to disintegrate the melt stream, the upper water-stream in a combined two-stage water- and gas-atomizer mainly has only this function. After primary atomization has been executed in this zone the water stream also carries away the particles.

Within the atomization zone the pressure-gas streams (the liquefied-gas streams or the

water jets) bounce against the continuous melt stream accelerate it downwards and disintegrate

it into droplets by tangential and normal forces. (In the above mentioned case of spiral gas

supply, parts of the tangential forces are centrifugal ones.) The gas jets (or liquefied gas or water jets) originate from an arrangement of identical orifices or from a slit nozzle arranged

concentrically to the central melt stream. Concerning the direction of the melt stream one can

recognize gas jets with moderate inclinations, see Fig. 2, nearly parallel ones, but also some with

angles a of about 45°.

In order to provide sufficiently high momentum forces for a satisfactory disintegration

very high gas velocities and gas mass flow rates are employed. While the velocity of the melt

stream in the atomization zone rarely exceeds values of about 20 m/sec, the gas jets reach

velocities of more than the 10- fold value, limited only by the sound velocity.

Atomization bv Means of Rotating Elements

The application of centrifugal forces to the disintegration process is mainly done in ac- cordance with the same scheme the melt poures under free fall conditions from the

crucible/tundish with a constant velocity. It enters the rotary atomizer of diameter D2 which

rotates at a constant speed u. Here the melt wets the rotating surface which has the form of a

flat disc, a drum or a cup shape, see Fig. 3. Because of the adhesion of the wetting melt layer

and its viscosity the melt is accelerated by the rotating surface from which it is spread out at least fairly uniformly, depending on the flow rate. (As there exists a pronounced region of

filament formation this type of atomization is also used for fiber production of melts with high

viscosities near the solidification temperature.) To prevent the melt from creeping over the rim

of the rotary atomizer, auxiliary secondary gas streams blow from below and carry the droplets

slightly upwards in the radial and tangential direction.

The generation of droplets by means of rotating elements can be classified into three

main groups governed by the rotational speed u as well as by the flow rate of the melt V and

corresponding to the main phenomena observed during the variation of these parameters, (see

Fig. 3) i.e.,

centrifuging off droplets directly from the rim .. (a),

drops emerging from previously formed ligaments .. (b),

drops forming from breaking liquid sheets (films). .. (c).

Generally expressed: for a given melt (constant temperature and fixed material

properties) the regularity of the powders (i.e. the narrow size distribution of the powder particles) will improve if the rotational speed of the elements is raised or the melt stream is diminished.

Looking at Table 2 a considerable number of equations are seen describing the borders

of the above mentioned fields of droplet generation with some serious contradictions. The equations have been derived for conventional atomization of aqueous or organic liquids, where the

material properties cover a far smaller range than they do for molten metals, ceramic or glass

melts. In Fig. 3, after [9], there can be seen differing results for the conventional atomization

(i.e. for the border between (a), (b) and (c).

38
.*0.S 0.083 V Z

Fig. 3 Liquid atomization by rotating elements

Table 2 Equations describing the borders between droplets centrifuged off directly

(a), drops emerged from filaments (b)

and drops created by breaking liquid sheets (films), i.e. (c)

ZONE Equations Lit. cit No.

a/b:

drops

-

drops

Troni

ligaments

i'^e0.25zOA61 = 2.08 10-5 [10]

^'^tO.lSzO.m = 3.068 [11]

v'We 0.85Z0.45 = 0.113-0.195

2 [12]

f 10.667 a

^ = 13.40

1''-r J p^(l + 10 Z 0.167)

V [mVs]; D [m]; w [rad/s]; a [N/m]; pjkg/m^]

[13]

b/c:

ligaments

-

film v' We 0.6 , ZO.167 = 1.77

2 [14]

^'WejO.SSS zO.166 = 1.45 [10]

^' = 4.443 10-2 . D2-0-5 u-0.667(a/^)0.833

V [m3/s] and o±er terms as given above

[13]

with V2/>L „. pL^^i ^

V = =5 ; We - — ; Z - ^.^

The atomization of metal melts takes place in the region left from the border (a), after

[9]. This is why in the following only drop size equations will be discussed, which have their

origin in this parameter combination.

39
Utilizing Forces Induced bv Ultrasonic Vibrations

Also in the area of melt atomization by utilizing ultrasonic forces the above described

free fall of the melt stream from the crucible/tundish to the vibrator zone is used as sketched in

Fig. 4 after [15]. The melt wets the surface of the conical ultrasound vibrator before being disintegrated into very small droplets under the influence of capillary forces. The droplets are sprayed off in all directions.

The free falling melt jet is applied for leading the melt stream into the central one of

several pressure node§ of a horizontal standing ultrasonic wave generated by two sonotrodes

after [16], as sketched in Fig. 4b. This arrangement prevents wetting of the two vibrating surfaces of the sonotrodes by the melt and helps for instance to keep the molten metals extremely

pure. The efficiency of this disintegration process can be increased by raising the pressure of

the gas atmosphere in the atomization chamber, i.e. the density of the vibrating gas column between the sonotrodes.

Another type of ultrasonic atomizer uses vibration of (20 to 100) kHz of the gas jets

impinging onto the melt. The so called USGA-nozzles surrounding the melt stream each incorporate a Hartmann Shockwave tube in order to generate pulsed high velocity gas jets, Fig. 4c.

To change the flow characteristics one needs different atomizer design (altering inclination

angle, number of gas nozzles and insert length), because of the restriction in the mass flow

ratio, i.e. keeping constant the ultrasound parameters. The vibrating gas jets assist the production of fine disoersed Dowders.

(a) (c)

Fig. 4 Atomizers using ultrasonic vibrations: (a) a vibrating conical surface, (b) an ultrasound

standing wave, (c) USGA nozzles

Utilizing Drop Impaction for Further Disintegration

Utilizing controlled drop impaction for further disintegration of larger^ drops into significantly smaller ones can only be an additional process step after primary atomization.. Assuming that the drops leave the primary atomization with high velocities, this together with the high

density of several metals can be used for subsequent impact disintegration, because of the

remainder high momentum of the particles.

If for instance rotary atomizers are used for the primary disintegration this first step

together with the impact disintegration can be handled under vacuum conditions, which may be

important with respect to a postulated extreme purity of the materials. Using vacuum means also

that the drag resistance of the particles can be diminished significantly thus raising the impact

velocity necessary for further disintegration.

The limiting condition for impact disintegration is given by the declining line in the

diagram after Fig, 5, where the vertical and the horizontal axes are given by the number A and

the Weber-number We3 after [17]:

1. The adjectives larger and smaller have to be understood in this context only as a relative characterization, there are no

quantifying equations available!

40
A =

We3 =

(3)

(4)

If a drop of diameter d impinges perpendicularly on the flat (not wetted) surface of the

disintegration element with the impact velocity V3(d) further disintegration in several significantly smaller drops will appear only when the couple of A and We3 mark a point to the

right of the declining line. The diagram has been established for aqueous liquids and for mercury. There exists no equation that describes the diameters of the secondary significantly smaller

droplets.

Other Principles of Atomization

Leading the melt flow as a continuous central stream into the atomization zone, and sur- rounding it with a concentric gas stream has been explained as a typical design principle of gas

atomization. Utilizing this principle in a radical manner in order

to bring the ambient gas stream into a close contract with the metal stream

to intensify the heat transfer between both phases

to introduce shear forces from the gas stream to the melt stream and

to avoid backsplashing of droplets

leads to solutions sketched in Figrs. 6a and b. Using Laval principle (injection principle) for the

surrounding gas stream, thinning the walls between the melt and the gas phase (with completion

already in or above the acceleration zone of the gas stream) and raising the gas temperature

above or up to the melting temperature opens the possibility to accelerate both phases and

therewith to augment their inertial forces as well as to lift the viscous forces of the gas stream

due to higher temperatures. There have also been successful attempts to verify laminar boundary

layers because of the high gas viscosity and to produce extremely small droplets. Accelerating of

the melt jet means diminishing its diameter and creating an excellent basis for break-up of an

extremely thinned melt jet just below the smallest nozzle cross section. It has been reported that

this method allows a powder production of ultrafine particles, [18].

a

10

10

10

10

^ OUInt*gr*tlon

A H,0- Qlyearln XO

0 »9

A H,0

• • •

10 10 10 10

d tfp

A = ^

(a) (b)

Figrs. 5 (left): Diagram for further drop integration after impact on a plane (not wetted) surface

and 7 (right): Atomization by means of a rotating porous cylinder

A similar way of atomizing the melt received from the melting zones of a wire is shown

in Fig. 6b after [19] and [20] for a flame spray gun. The wire (identical with the central melt

stream) is supplied continuously from behind (left) into the melting zone of the burning gases.

Starting from the surface the material of the wire is molten and thrown away by the gas stream

41
onto the surface of the substrate. The melt droplets are kept liquid until reaching the impact

zone by radiation, conduction and convection within the gas flame of the gun thus forming a spray cone of only 10 to 30 cm in length.

The flame spray is used to generate special metallic or ceramic coatings on surfaces

(improving the resistance to wear, heat or corrosion). The droplet impact and the complete spray

deposition process is similar to that described below as the Osprey spray compaction process. In

this context there has to be mentioned the plasma spraying as an alternative spray deposition

process (widening the application of materials and properties of coatings by raising the plasma

gas temperatures up to some 10^ Kelvin. While the two phase flow in the spray and the deposition process are once again similar to the above mentioned process and to that of Osprey, the

melting process differs greatly by using powder particles which are molten within a very short

distance because of the extreme high temperatures in the plasma flame. The assumption is that

the droplet diameters are prescribed by the sizes of the powder particles, whether the droplets

underlay further disintegrations cannot be answered till now.

A combination of the principles of rotating and impacting has been verified by [21]

using a hollow cylinder of porous material of sintered monodispersed spheres of ceramics or

metals. This sintered material should not be wetted by the melt, thus allowing only very dedicated combinations of metal/metal or metal/ceramics. The principle is illustrated in Fig. 7a: the

melt is fed from above and centrifuged through the pores. Fig. 7b shows how the melt is pressed through the upper pores and disrupted into larger and smaller droplets (generating also

secondary smaller droplets by impact). The pores are forcing the smaller drops to coalesce

before leaving the last (outer) zone of the pores thus giving cause for a very narrow size distribution.

Fig. 6 a) Laval principle after [18] and b) flame spray gun after [19] and [20]

OVERALL SIZE DISTRIBUTION

The most important prerequisite for atomizer design, for process layout, process optimization, and process control as well as for the product quality is knowledge of the droplet

diameters. This holds not only for the conventional atomization of aqueous or organic fluids but

also for melt atomization, where for instance the size distribution of the pulverized metal

directly influences its acceptance and its price. The dependencies on the conditions of unit

operations integrated in the same process play an important role. And last but not least problems

arise from environmental and safety conditions especially for the treatment of fine and ultrafine

powder particles.

Great efforts have been made to describe the size distributions of droplets obtained by

atomization in different atomizer types depending on atomizer design and geometry and also on

42
the material properties of the process media. Some attempts have been successful. Especially for

conventional atomization, there have been derived numerous dropsize equations for the very

different types of atomizers and operating conditions. A very helpful and worth reading survey

of such equations is given in [4] including also results of his own research. Other equations also

valid for aqueous or organic liquids are mainly derived from ratios of similitude and lead to

dimensionless correlations. Some attempts failed completely in transforming them from conventional to melt atomization because of the not allowed extrapolation of parameters and because

up till now no theoretical description of the disintegration process in general is available. Especially with respect to melt atomization this may also depend on the deficits in awareness of the

necessary material properties of melts with regard to the great variety of alloys and composites.

As long as it is satisfactory for a well operating atomization process to keep constant the

overall-size distribution with respect to the total mass flow of the powder production, it may be

sufficient to take periodical samples of the powder and to screen these samples or to measure

the size distribution by means of (for instance) a diffraction instrument and to describe or to

control the size distribution by the results of these measurements.

In order to describe the overall size distributions of melt droplets generated by gas- or

liquid atomization one can recognize especially one equation (after Lubanska [22]) on which the

experimental reports are focused chiefly and which has been used successfully in a great number of experiments:

*50,3 = B 1 + Wei D1 (5)

u2 Di Pl

with Wei = (6)

and B = 50

Lubanska observed size distribution closely following a log-normal pattern

\2]

(7)

q3 (d) = dR

dJ

100

S 2 TT

|

(l0g d -^lOg d,n)^|

(8)

with dR being the cumulative weight oversize percentage per screen after sieving and

^50 = 3 as given by equation (5). He gave the standard deviations as depending on d5o^3 :

d5o,3 = 13 [ Mm ] • s2

where s is given as

(9)

s = log

^

(10)

"50,3

In this equation Dj is the diameter of the continuous melt jet before reaching the gas

jets (i.e. the zone of disintegration) and u is the velocity difference between both types of jets

at this zone. Because it is difficult to measure these terms one may use instead of Dj, i.e. the

diameter of the exit of the crucible/tundish, and one has to make assessments for the velocity

difference.

As mentioned above the melt jet rarely exceeds velocities at the disintegration zone of

more than 20 m/sec, whereas the gas streams may reach more than the 10-fold value, because

the gas jets normally are operated at velocities somewhat below the sound velocity. It's now

depending on the distance between the disintegration zone and the gas nozzle exits, Az, as well

43
as on the characteristics of gas- velocity governed by the nozzle cross sections and the gas pressure prior to the nozzle exits of how strong the decline of the gas velocity will be along the distance z, see for instance paper [23] of these proceedings.

For the utilization of rotating elements for melt atomization one has a quite simple

overall drop size equation:

d50,3 = CWe2"^ D2 (11)

with

We2= -^i^ -f (12)

As can be seen from [9] there are - derived from experiments and from reports in the

literature - factors C of 2.67 up to 6.55, the latter also given for molten metal, but aware of

material constants. Often no information can be found whether the melt mass flow rate had

been changed or not. (The angle j of the edge of the rotating plate - as can be seen from Fig. 3

- may influence the factor C.) As mentioned earlier, for melt atomization there will be of importance only the drop formation region directly from the edge of the rotating plate. This is why here are given only equations for this region.

Because of the similarity of the rotating elements compared with the rotary atomizers for

conventional atomization and because of the same forces governing the disintegration processes

it should at least in this context be allowed to cite one of the various equations which have been

derived for aqueous or organic liquids, for instance those which have been discussed by [9]. One

should expect that the deviations between metal melt atomization and conventional atomization

could be explained by the material properties.

As also derived by [9] the size distributions mainly can be described by the log-normal

type. While there is no information about the standard deviation, the d503 diameter, which

principally contains the described influences in the right manner, has been given as

^..J VL^^.r -0-07 ^-0.33

d5o,3 = 0.376 j We2 • V . Z , (13)

with V and Z given already in Table 2.

LOCAL DROP SIZE DISTRIBUTIONS AND VELOCITIY DISTRIBUTIONS IN METAL

SPRAYS

The principles of rapid quenching or of spray - or of plasma - deposition make it necessary to posses of detailed knowledge of the multi-phase-flow situation in the spray cone (of gas,

plasma, melt droplets, solidified particles) in order to describe and control the flow field generated by the atomizer. Research programmes are concerned with the laws which govern the particle transport onto a shaped collector (mold, substrate or preform), and with the phenomena

together with the mechanical processes of particle impact, particle flattening and welding

together, and last but not least with the overall heat transfer and solidification process of the

droplets.

For a stationary spray casting process (as sketched in Fig. 8) there are mainly six independent process parameters which govern the spray process and which must be optimized and

controlled: The melt temperature, its effluent velocity and its mass flow rate describe the

material delivery and its flow situation before comminution. The gas mass flow rate and its pressure govern the gas jets which leave the ring nozzle and bounce against the melt stream.

These parameters guide the disintegration process, which results in an overall droplet size distribution. This is given by eqs. (5), (8) and (9), if the relative velocity difference between the

melt stream and the gas jets at the disintegration zone and the material properties of gas and

44

I
melt are known. Now the droplets are accelerated by the carrier gas towards the substrate. Here

they impinge and consolidate completely to form a bulk-net or near-net shape. The stand-off

distance between the atomization point and the substrate is important for the flow situation of

the multi-phase flow in the spray and the heat transfer efficiency between the particles and the

carrier gas up till the moment of impact. This distance therefore governs the temperature and

the status (liquid or solidified) of the particles directly before impact.

melt temperattire Tj, ±

Fig. 8 Stationary spray casting process with six independent parameters;

impact situation on the right

In order to guarantee successful operation an uniform distribution of the fractions of

still liquid droplets and already solidified ones on the substrate should be attaint. This has to be

observed by the on-line measurements of velocity- and size distributions of droplets in the

spray as a part of a process model of the atomization and the multiphase flow situation (i.e. the

mass- and momentum-balances) which also give the basis for the corresponding heat balance. In

[24] have already been described the principles of a numerical simulation model especially for

the Osprey process including the model for the heat transfer in the spray and the different

solidification processes of larger and smaller droplets during their flight downwards. This

simulation model has been adapted to the results of temperature and phase-DoppIer (PDA)-

measurements. The local velocity vectors of gas or droplets in the spray chamber can be seen from Fig. 5, the gas- and droplet velocities versus flight distance from Fig. 6 of paper [23] of

these proceedings, the local mass fluxes of droplets, measured and calculated can be seen from

Figures 4 to 6 of paper [25], also in these proceedings.

CONCLUSIONS

Having focused the discussion first onto the gas atomization it can in general be said

that the necessity to introduce sufficiently high momentum forces in order to guarantee a satisfactory fine pulverization by means of the gas streams requires high mass flow rates and high

velocities of the gases and hence high Reynold-numbers, i.e. immense turbulence ratios. (This

turbulent disintegration process cannot yet be described theoretically nor can the equations be

deduced which describe the drop sizes or size distributions generated by these turbulence

forces.) After joining, the gas streams form a common jet stream and create a flow field that is

45
similar to those which are well known from free jets [25], loaded in our case with the dispersed

melt phase, mostly accompanied by high densities.

In comparison with the conventional atomization the problems of melt conduction seem

to simplify the situation by reducing the types of applicable atomizers. This also holds for attributing the complicated two phase flows to the free jet stream and in addition for having only

one important drop size correlation (eq. (5) to (10) of Lubanska [22]) till now which fairly

governs the gas atomization scenario and most metal powder size distributions.

On the other hand it has to be remembered that in general there exists no information

about the difference velocity between the gas jets and the melt jet near the disintegration zone

(needed in eq. (6)). Further the exact material properties (i.e. the density, the kinematic viscosity and mainly the surface tension) are very often unknown. If we rush to a conclusion this

would make it trivial to use the formula for the overall particle size. But keeping in mind the

extremely helpful special feature of melt atomization (which by its product generates an alibi

for an assessment of the process, i.e. the powder metal which can be taken as local or overall

samples, which can be screened and which provides the necessary size distribution) there is support for changing the operating conditions for a given atomizer in a very successful manner

and nevertheless to utilize therefore the equations (5) to (10)!

Now considering the rotary atomizers and their very similar design compared with the

conventional atomizers, one should be allowed to speak of similar or identical equations that

describe the drop generation phenomena as well as the overall drop size. But the attempts to describe melt atomization by the equations derived for the conventional atomization seem to be

successful only for the region of drop formation directly from the edge of the rotating plate. An

explanation can be found in what has been said in regard to the deficit knowledge about the

material properties. This is the cause why also in the case of rotary atomization (for a given

design) the equations can be successfully used only for relative alterations of the operating conditions after having fixed the type of melt stream.

These results can be generalized also for the ultrasonic atomizers and for the secondary

atomization of impacting drops. Only under very special preconditions, for instance where pure

or well known metal melts or glass melts have to be atomized, the material properties are well

established or given in tables versus melt temperature. (These are opportunities to check the

above mentioned equations for their validity.)

With regard to the more common case where the material properties have to be extrapolated or interpolated or where have to be made assumptions there can be received important

results from model experiments (from model melts and for fixed atomizer types and geometries).

In this case it would be very helpful, if results of model experiments with for instance water

(and air) were available. But here the problem arises that water droplets coalesce and tend to evaporate after atomization (i.e. losing their alibi); which in the past made it nearly impossible

to compare results from the same atomizer obtained from extremely different liquids.

Here the development of new instruments, for example of the PDA and of very short

lightning imagination techniques open an interesting field of future research work. The PDA

especially seems to be able to provide quasi on-line results from metal melt atomization as well

as from water atomization thus permitting the crossing of the bridge from conventional

atomization to metal melt atomization and melt spraying and to answer the questions for a generalized theory for all types of liquids under the task of atomization.

List of Symbols

A Constant factor in eq. (3) Vg Gas velocity

B Constant factor in eq. (5) and (7) Particle velocity

C Constant factor in eq. (12) Melt velocity in eq. (1)

d Drop diameter Impact velocity in eq. (4)

Drop diameter (50 %) of volume-dist. Volume flow rate of melt after Table 2

Do Diameter of bottom exit in tundish Dimensionless volume flow rate of melt

D2 Diameter of rotary atomizer z Dimensionless factor after Table 2

46
g Acceleration due to earth a Angle of inclination of gas jets versus melt

h Level of melt above exit of tundish Effluent factor in eq. (1)

Mq Gas mass flow rate r^L Dynamic viscosity of melt

ICIl Melt mass flow rate l^L Kinematic viscosity of melt

Ap Pressure difference 1/q Kinematic viscosity of gas

q Size distribution, vol.fraction in eq.(8) (7 Surface tension of melt/gas

R Cummulative weight oversize in eq. (8) Density of melt

T Temperature of melt w rotational speed of atomizer

Tl Melting temperature We^ Weber-number in eq. (5) and (6)

AT temperature difference We2 Weber-number after Table 2 and eq. (11)

u Velocity difference in eq. (6) We3 Weber-number in eq. (4)

References

1. Lawley A., Doherty R.D., "Atomization Models and Mechanisms in Advanced HighTemperature Alloys" in N. Grant Symp. 1985 Mass. Am. Soc. for Metals

2. Mehrotra P.S., "Mathematical Modelling of Gas Atomization Process for Metal Powder

Production", Part 1, Powd. Met. Int. Vol 13 No 2 (1981) 80-84, Part 2 ibidum, 3, 132-135

3. Savage S.J., Froes F.M., "Production of Rapidly Solidified Metals and Alloys" J. of. Metals

April (1984) 20-33

4. Lefebvre A.H., "Properties of Sprays" Particle & Part. Syst. Charact. 4 (1989) 176 - 186

5. Grassmann P., "Physikalische Grundlagen der Verfahrenstechnik" Salle & Sauerlander

Frankfurt/Main (1983)

6. Ruthardt R., "Basic Characteristics of LAB-Scale Argon Atomization Units"; Mod.

Developm. in Powder Metall. Vol. 15-17 (1985) 67-72

7. Lewis R.E., Fitzpatrick J.M., Crooks D.D., Singer A.R.E., Jenkins W.N., "A Novel Approach for Rapid Solidification Processing of High Strength PM Aluminum Alloys", SAE

Technical Paper Series 841514, Aerospace Congress & Exposition, L.B. Cal. Oct. 15-18

(1984) 27-33

8. Schellenberg R., "Zerstaubung von Metallen, insbesondere von Zink", Diss. T.H. Stuttgart

(1964)

9. Mehrhardt E., "Zerstaubung von Fliissigkeiten mit rotierenden Scheiben", Diss. TU Berlin

1978

10. Dombrowski N., Lloyd T.L., "Atomization of liquids by spinning cups", Chem. Engng. J. 8

(1974) 1, 63-81

11. Dunskii V.F., Nikitin N.V., "Monodisperse atomization of a liquid by a rotating disc", JH.

Engng. Phys. 23 (1972) 5, 1474-1475

12. Matsumoto S., Saito K., Takashima Y., "Phenomena Transition of Liquid Atomization

from Disk", J. Chem. Engng. Japan 7 (1974) 1, 13-19

13. Tanasawa Y., Miyasaka Y., Umehara M., "On the filamentation of liquid by means of

rotating discs". Trans. Soc. Mech. Eng. Japan 25 (1959) 879-905

14. Hinze J.O., Milborn H., "Atomization of liquids by means of a rotating cup", J. Appl.

Mech. 17 (1950) 145-153

15. Bendig L., "New Developments of Ultrasonic Atomizers", Proc. of the 4th Int. Conf. on

Liquid Atomization and Spray Systems, August 22-24, 1988, Sendai, Japan

16. Bauckhage K., Schreckenberg P., "Control of Powder-Metal Production; A New Application of Phase-Doppler-Anemometry"; Intern. Conf. on Mechanics of Two-Phase-Flows,

12.-15.6.1989, Taipei/Taiwan, ROC, Proc.

17. Walzel P., "Zerteilgrenze beim Tropfenprall", Chem.-Ing.-Tech. 52 (1980) 4, 338-339

18. Geropp D., "Die Berechnung reibungsbehafteter Strfimungen in Lavaldiisen mit Anwendungen" Manuskript eines Kolloq.-Vortrages am 1990 in Bremen sowie

Walz A., "Entwicklung eines Verfahrens zur Herstellung keramischer Fasern", BMFTBericht NT 321 Juli 1974

47
19. Meringolo V., "Stop Tube Corrosion and Wear in Coal-fired Power Boilers", Powermagazine, May 1981, Mc Graw-Hill, Inc.

20. Anonymus, "Advanced Plasma Gun", bulletin No. 283 of Metco, Percin Elmer Westbury,

Long Island N.Y.

21. Schmidt P., "Zerteilen von Flussigkeiten in gleich groBe Tropfen"; Chem.-Ing.-Tech. 39

(1967) 5/7, 375-379

22. Lubanska H., "Correlation of Spray Ring Data for Gas Atomization of Liquid Metals", J. of Metals (1970) 45-49

23. Fritsching U., Liu H.-M., Bauckhage K., "Numerical Modelling in the Metal Spray Compaction Process"; paper presented on ICLASS '91, included in these proceedings

24. H. Liu: "Berechnungsmodelle ftlr die Geschwindigkeiten und die AbkQhlung von Tropfen

im Spriihkegel einer Stahlzerstaubungsanlage"; Diss. Universitat Bremen, Febr. (1990)

25. Uhlenwinkel V., Fritsching U., Bauckhage K., "The Influence of Spray Parameters on Local Mass Fluxes and Deposit Growth Rates During Spray Compaction Process", paper

presented on ICLASS '91, included in these proceedings

48
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper D

Keynote Lecture

TWIN-FLUID ATOMIZATION: FACTORS INFLUENCING MEAN DROP SIZE

A.H. Lefebvre

School of Mechanical Engineering, Purdue University

West Lafayette, IN, U.S.A.

ABSTRACT

The principal factors governing the mean drop sizes produced by twin-fluid atomizers are examined. It is shown

that atomization may occur by one of two different mechanisms. For both "classical" and "prompt" modes of

atomization, the principal factors governing the fineness of atomization are air velocity, air/liquid ratio, and surface

tension. For the classical mechanism of breakup, involving flow instabilities and wave formation, atomization

occurs slowly and is strongly influenced by variations in Uquid viscosity, air density, and the initial diameter (or

thickness) of the liquid jet (or sheet). If atomization occurs very rapidly, these parameters have little effect on mean

drop size.

INTRODUCTION

The description "twin fluid atomizer" can be applied to any device in which the driving force for liquid breakup is either air, gas, vapor, or even another liquid. The following discussion will be confined to the atomization of liquids

by gases which, for convenience, will be referred to henceforth as air. Most twin-fluid atomizers fall into the

categories of "air-assist" or "airblast," the main difference between them being that the former use relatively small

quantities of air or steam flowing at very high velocities (usually sonic), whereas the latter employ large amounts of

air flowing at much lower velocities of around 100 m/s. Airblast nozzles are thus ideally suited for atomizing liquid

fuels in continuous-flow combustion systems, such as gas turbines, where air velocities of this magnitude are

usually readily available. The most common form of airblast atomizer is one in which the liquid is first spread into a thin conical sheet and then exposed to high-velocity airstreams on both sides of the sheet. The atomization

performance of this prefilming type of airblast nozzle is superior to that of the alternative plain-jet airblast nozzle,

in which the liquid is injected into the airstream in the form of one or more discrete jets [1]. In recent years, considerable interest has been aroused in an alternative method of twin-fluid atomization in which

a gas (usually air) is injected into the bulk liquid at some point upstream of the nozzle discharge orifice. The

injected gas forms bubbles which produce a two-phase flow at the injector orifice. The liquid flowing through the

injection orifice is "squeezed" by the gas bubbles into thin shreds and ligaments. When the bubbles emerge from the

nozzle they "explode", thereby shattering the surrounding liquid into small drops.

This paper examines the factors governing the mean drop sizes produced by twin-fluid atomizers. Attention is focused primarily on the effects of air and liquid properties on mean drop size. It is shown that at least two different

mechanisms may be involved in airblast atomization. Recognition of these different mechanisms helps to explain

some of the apparent anomalies that exist in the atomization literature. It will be shown that the effects of certain

air and liquid properties on mean drop size may be large or insignificant depending on the mode of atomization.

These considerations open up the possibility of designing airblast atomizers specifically to employ either one or the

other of these two different modes of atomization, depending on which is best suited for any given application.

CONVENTIONAL TWIN-FLUID ATOMIZATION

During the past 25 years airblast and air-assist atomizers have been subjected to considerable experimental study.

An early investigation on prefilming airblast atomization by Rizkalla and Lefebvre [2] led to the following

dimensionally-correct equation for the Sauter mean diameter (SMD).

SMD =3.33 X 10-3 + i/ALR) 13.0 x 10^-^ t^-^'^^^i + i/ALRp (1)

^aUa Wpl/

An interesting feature of this equation is that the SMD is directly related to the initial thickness of the liquid

sheet, t. For liquids of low viscosity, SMD OC lO S por liquids of high viscosity, the dependence of mean drop

size on t is slightly higher.

49
A drawback to Eq. (1) is that t is usually unknown and difficult to estimate. This problem is avoided in the

following "basic" equation for the mean drop sizes produced by prefilming airblast atomizers [1]

SMD _

[1 + 1/ALR] (2)

where A and B are constants whose values depend on the atomizer design and must be determined experimentally.

Equation (2) shows that the SMD is proportional to a characteristic dimension Lq, which represents the linear

scale of the atomizer. It also shows that mean drop size is inversely proportional to the square root of the prefilmer

diameter Dp. This is consistent with Eq. (1) because, if other parameters are kept constant, an increase in Dp will

reduce the liquid sheet thickness t, and hence the SMD. It will merely be mentioned here, and discussed more fully

later, that the strong dependence of SMD on t, as expressed explicitly in Eq. (1) and embodied in Eq. (2), is not

always observed in practice. For some atomizers, it is found that the influence of initial sheet thickness on SMD is small, which tends to suggest that the mechanism whereby a liquid sheet is converted into drops may differ from one

atomizer design to another.

Influence of Air Properties

Perhaps the least controversial aspect of airblast atomization is the role played by atomizing air velocity. A

unanimous conclusion from all experimental studies on airblast atomization is that air velocity has an important and

sometimes overiding effect on the drop size distributions in the spray. Generally, it is found that mean drop size is roughly inversely proportional to atomizing air velocity, i.e. SMD a U~^, although the literature contains values

of velocity exponents varying from -0.75 to -1.33- It is important to bear in mind that the velocity of most

relevance to airblast atomization is not the absolute velocity of the atomizing air but rather the relative velocity

between the atomizing air and the liquid jet or sheet. This point is well brought out in the results obtained by

Suyari and Lefebvre {3} who used an external-mixing, air-assist nozzle of the type shown in Fig. 1, which

comprises a simplex, pressure- swirl nozzle surrounded by a co-flowing stream of swirling air. Fig. 2 exhibits some interesting features which appear to be characteristic of this type of atomizer. First, it shows that an increase in APa, which corresponds to an increase in atomizing air velocity, has a beneficial effect on

atomization quality, especially at low liquid flow rates. Of more interest, however, is the observation that, for all values of APa, SMD increases with increase in liquid flow rate until a maximum value is reached, beyond which

any further increase in riiL causes SMD to decline. The unusual shape of this curve can be explained in terms of the

relative velocity between the conical sheet of liquid produced by the simplex atomizer and the surrounding coflowing

air. At low liquid flow rates, the device operates primarily as an airblast atomizer, and any increase in liquid flow

rate impairs atomization quality, despite the increase in liquid injection pressure, by reducing both the air/liquid ratio

and the relative velocity between the air and the liquid. At the peak of the curve, where the mean drop size attains

its maximum value, the relative velocity and hence the interaction between the liquid sheet and the surrounding air is minimal. Beyond this point the device functions mainly as a pressure atomizer, and any increase in liquid flow rate serves to improve atomization by increasing the liquid injection pressure and hence the relative velocity between the

liquid and the air. In most practical forms of air-assist atomizer, the air and the liquid issuing from the nozzle have both radial and

tangential as well as axial components of velocity. In this complex two-phase flow field the term "relative velocity"

is difficult to define. Nevertheless, for any given air velocity there will always be one particular value of liquid

velocity at which the level of interaction between the liquid and the air is a minimum. An increase in atomizing air

velocity will cause the condition of minimum relative velocity to occur at a higher level of liquid velocity. Thus

one would expect the value of niL at which the SMD attains its maximum value to increase with increase in

APa/Pa, and this is indicated by the dashed line drawn in Fig. 2. These results underline the importance of relative

velocity in the airblast atomization process.

Another important factor influencing atomizer performance is the air/liquid mass ratio. With conventional

airblast atomizers, atomization performance starts to decline when the ALR (air/liquid ratio) drops below around 3,

and deteriorates quite rapidly at ALR's below about 1.5. At ALR's higher than about 5, only marginal reductions in SMD are gained by the utilization of more air in atomization. This is because air which is physically far removed

from the liquid cannot play an effective role in the atomization process. For this reason, it is customary in equations

for mean drop size to express the dependence on air/liquid ratio as SMD OC [1 + 1/ALR]", where n has values

50
ranging from 0.4 to 1.0. This expression allows the influence of ALR on mean drop size to diminish with increase

in ALR, in accordance with experimental observations.

Influence of Initial Sheet Thickness

In their early studies on prefilming airblast atomization, Lefebvre and MiUer [4] stressed the importance of

spreading the liquid into the thinnest possible sheet before exposing it to the atomizing air, arguing that "any

increase in the thickness of the liquid sheet flowing over the prefilmer lip will tend to increase the thickness of the

ligaments which, upon disintegration, will then yield drops of a larger size."

Subsequently, Rizk and Lefebvre [5] conducted tests to examine the effect of initial liquid sheet thickness on the

drop sizes produced by prefilming airblast atomizers. They employed a specially-designed airblast atomizer in which

a flat liquid sheet was produced across the centerline of a two-dimensional air duct, with the liquid sheet exposed on

both sides to high velocity air (see Fig. 3a). This atomizer was designed with the capability of varying the initial liquid sheet thickness in order to examine its influence on atomization quality. The results obtained showed that

thicker liquid sheets resulted in coarser sprays. Measurements of mean drop size carried out over wide ranges of air

velocity, air/liquid ratio, and initial sheet thickness, indicated that the effect of sheet thickness on mean drop size

could be expressed as SMD a t^, where x has a mean value of 0.38. This result was considered to be consistent

with Eqs. (1) and (2) and generally supportive of the classical wavy-sheet mechanism for liquid sheet disintegration,

as postulated theoretically by Squire [6} and Fraser et.al. [7], which leads to the relationship SMD a t^-^. Some of

the results obtained by Rizk and Lefebvre [5] on the influence of t on SMD are shown in Fig. 4. Other workers have also observed an appreciable effect of initial liquid sheet thickness on mean drop size. For

example, Inamura and Nagai [8] in their performance comparison of various types of airblast atomizers noted that

"SMD is nearly proportional to the film thickness at the point of impingement of the liquid film and the air stream."

Zhao and Gan [9] examined the influence of initial sheet thickness on the mean drop sizes produced by a prefilming

airblast atomizer. Their results for kerosine are shown in Fig. 5. According to Zhao and Gan, these data indicate

that SMD a t^-^^, which is remarkably close to the findings of Rizk and Lefebvre, as noted above. However,

close inspection of the experimental data in Figs. 4 and 5 reveals that the dependence of mean drop size on initial sheet thickness diminishes with increase in atomizing air velocity. This trend was also observed by Beck et.al [10,11] who employed a two-dimensional flat-sheet atomizer, shown schematically in Fig. 3b, to study the

influence of sheet thickness on mean drop size. They also found that the effect of t on SMD diminishes with

increase in atomizing air velocity, as illustrated in Fig. 6. In all these studies, the initial liquid sheet thickness was found to influence the mean drop size in the spray.

The magnitude of this influence varied appreciably with atomizing air velocity, but in all cases an increase in sheet

thickness was accompanied by an increase in mean drop size. Contrasting results were obtained by Wittig and his

co-workers [12,13] who found no effect of sheet thickness on mean drop size. They attributed this result to the fact

that in their atomizer the liquid does not flow continuously over the atomizing edge but is stored in the wake of the

edge in the form of a bulge. This bulge increases in thickness until the aerodynamic forces exceed the surface forces,

at which point the bulge disintegrates into ligaments and then drops.

Consideration of these and other apparent anomalies in the atomization literature suggest that the disintegration

of the liquid sheet into drops may not always occur via the classical wavy-sheet mechanism. According to this

mechanism, waves are formed which protrude into the coflowing air stream. The most rapidly growing wave

becomes detached from the leading edge of the sheet to form a ligament which subsequently breaks down into drops.

The size of these drops is directly related to the diameters of the ligaments from which they are formed. These

ligament diameters, in turn, are dependent on sheet thickness, so there is a very direct relationship between the mean

drop size of the spray and the initial sheet thickness, as postulated theoretically by Dombrowski et al. [7,14] and

demonstrated experimentally by Rizk and Lefebvre [5] and Zhao and Gan [9]. An essential prerequisite for wavysheet formation is sufficient time for the waves to develop. In many cases ample time is available, because the air

and liquid are coflowing (see Fig. 3a, for example) and there is no significant component of air velocity in a

direction which could promote rapid disintegration of the liquid sheet. If, however, the atomizing air is arranged to impinge on the liquid sheet at an appreciable angle, as illustrated in Fig. 3b, it now has a sizeable transverse

component of velocity. In consequence, the liquid sheet emerging from the nozzle has no time to develop a wavy

structure but is instead rapidly torn into small fragments by the vigorous interaction created between the liquid and

the impinging air jets. The maximum critical size of the liquid fragments produced in this prompt atomization

process is governed primarily by the magnitude of the air velocity component normal to the liquid sheet, the

air/liquid mass flow ratio, and the surface tension. An essential feature of this mode of atomization is that the rapid

and violent disruption of the liquid sheet into drops ensures that the ensuing drop sizes are much less dependent on

the initial sheet thickness. This could serve to explain the different results obtained by different workers on the effect

of initial liquid sheet thickness on mean drop size.

51
Analysis

If atomization occurs very rapidly just downstream of the atomizing edge, the flow instabilities which normally

precede wave formation have no time to develop. Sheet breakup occurs so rapidly that the mean drop sizes produced

in atomization are governed solely by the ratio of the kinetic energy of the atomizing air to the energy required to overcome the consolidating surface tension forces. This hypothesis leads to the following expression for mean drop

size [15]

SMD = p

cpLui

t 4a(l + l/ALR)_

(3)

in which the constant C represents the ratio of the energy required for atomization to the kinetic energy of the

atomizing air. For a flat-sheet airblast atomizer, the data of Beck et.al. [10,11] indicate a value for C of 0.0070.

Thus Eq. (3) may be rewritten as SMD = p

0.00175 PLUi

or, more basically, as

.t a{l + 1/ALR)_

(4)

SMD ^

t 1 +0.00175 We

(1 + 1/ALR)

(5)

Where We = pLUA^t/<7

Equation (4) has some interesting features. For example, at the normal operating conditions of most practical airblast

atomizers, the second term in the denominator is always appreciably higher than the first. This explains why in

Beck et. als'. [10] experiments the SMD was found to be fairly insensitive to variations in liquid sheet thickness.

Equation (4) also predicts that the dependence of SMD on t should diminish with increase in Ua, as observed

experimentally by Zhao and Gan [9] and Beck et. al [10]

Another interesting feature of Eq. (4) is the absence of any term to represent liquid viscosity. Normally, with

coflowing airblast atomizers, an increase in liquid viscosity has two adverse effects on atomization quality. First, it tends to increase the thickness of the liquid sheet formed at the atomizing edge which, for this type of atomizer,

produces an increase in mean drop size. Second, an increase in viscosity hinders the development of any natural

instabilities in the liquid sheet. This delays the disintegration process so that when breakup finally occurs it does so under conditions that are less conducive to atomization. In extreme cases, the breakup time may be extended so much that little or no breakup occurs, even though the initial aerodynamic force was large enough to produce

breakup.

Measurements carried out by different workers using various types of coflowing airblast atomizers almost

invariably show a significant effect of liquid viscosity on mean drop size. The results of these investigations have

been summarized elsewhere [16]. However, in other types of atomizers in which the air and liquid are not coflowing,

but instead the impact of the atomizing air on the liquid is such as to promote prompt atomization, it is found that

mean drop size is much less dependent on viscosity. For example. Fig. 7 shows results obtained by Beck et. al. on

the influence of viscosity on mean drop size. They show that increasing the liquid viscosity by a factor of more than

thirteen produced only a 16 percent increase in SMD, i.e. SMD a ^il^-^^. A similar result was obtained by

Sattelmeyer and Wittig [12]. These workers studied prefilming airblast atomization under conditions where "the

disintegration of the liquid film always takes place close to the atomizing edge, regardless of the liquid's properties,"

i.e. under conditions which might reasonably be described as "prompt." They found that liquid viscosity had no

effect on SMD, which agrees both with the findings of Beck et.al. [10] and with predictions based on Eq. (4).

Another example of the insensitivity of mean drop size to variations in liquid viscosity is provided in the results

obtained by Kennedy [17] who carried out a comprehensive series of tests on twenty-five different hydrocarbon fuels,

using six different pressure-swirl nozzles of large flow number. In contrast to all previous work on pressure-swirl

nozzles, Kennedy found no effect of viscosity on mean drop size. The only essential difference between Kennedy's

investigation and those of previous workers is that his measurements of SMD were carried out on nozzles operating

at exceptionally high Weber numbers (>10) i.e. under conditions of very rapid sheet breakup. These results are fully

52
consistent with Eq. (5) which predicts that when the Weber number is high the mean drop size is independent of

liquid viscosity.

Another interesting feature of Eq. (5) is that for very high Weber numbers the mean drop size should be directly

proportional to surface tension. This is exactly the result obtained by Kennedy. Sattelmeyer and Wittig also

observed a pronounced effect of surface tension on SMD. They reported that atomization characteristics were

"drastically improved" by an almost three-fold reduction in surface tension, Beck et.al. also noted a strong effect of

surface tension on mean drop size, as illustrated in Fig. 8. Thus, ample results are available to support the validity

of Eqs. (4) and (5) for conditions of prompt atomization.

According to Eq. (4) air density has no influence on mean drop size. Most previous work on prefilming

airblast atomizers has shown a density dependence which is usually expressed as SMD OC with values of z

ranging from 0.5 to 0.7. However, over a range of air densities from 1.2 to 6.0 kg/m^, Beck et.al. [11] found no

effect of SMD on p^^ which confirms that when the airblast atomization process occurs very rapidly, air density has

little or no effect on mean drop size.

The results of the various studies outlined above show that when sheet breakup occurs too rapidly to permit the

development of waves and other flow instabilities that normally precede atomization, the mean drop size in the spray

is strongly dependent on surface tension, largely independent of liquid viscosity and ambient air density, and only

slightly dependent on the initial sheet thickness. These experimental observations are in complete agreement with

the predictions of Eq. (4) which was derived on the assumption that the liquid issuing from the atomizer is converted

into droplets without the help or intervention of flow instabilities and wave formation.

Plain-Jet Airblast Atomization

In many practical airblast atomizers the liquid is injected into the high-velocity air stream in the form of one or more discrete jets. Similar arguments to those employed above for flat-sheet atomizers yield the following equation

for the mean drop sizes produced under conditions of "prompt" atomization.

SMD ^ 1.5

do

1 + CWe

(1 + 1/ALR) (6)

where We = PLU^do/CJ and C is a constant which includes a term to represent the energy efficiency of the

atomization process and must be determined experimentally.

Equation (6) again illustrates that under conditions of high Weber number the mean drop size becomes

independent of air density and initial liquid jet diameter. It also serves to explain some of the anomalies that exist in

the literature on plain-jet airblast atomization. For example, Nukiyama and Tanasawn [18], using an atomizer in which the atomizing air impacted the liquid jet at an angle of almost 90°, thereby causing rapid jet breakup, found

no effect of Pa and do on SMD, in accordance with Eq. (6), whereas Rizk and Lefebvre's [5] results for a coflowing

air stream exhibited a strong influence on mean drop size of both Pa and do, (SMD a d o VpA ^) • The conclusion to be drawn from these various studies on both plain-jet and prefilming airblast atomizers is that

for both types the dependence of mean drop size on initial jet diameter (or sheet thickness) and ambient air density is governed by the rapidity of the atomization process. If the atomizing air impacts the liquid at an appreciable angle,

of if coflowing atomizing air encounters the Uquid at high velocity, breakup occurs very rapidly and the ensuing

atomization process is characterized by a lack of sensitivity of mean drop size to the initial diameter (or thickness) of

the liquid stream and to air density. If, however, the atomizing air is flowing at a relatively low velocity and in

generally the same direction as the liquid, the atomization process occurs more slowly, via a different mechanism,

with the result that drop sizes become much more dependent on the physical dimensions of the liquid stream and the

ambient air density.

MECHANISMS OF AIRBLAST ATOMIZATION

The experimental evidence cited above suggests that airblast atomization occurs by one of two different

mechanisms, depending on the time available. If ample time is available, the breakup process follows the classical

mechanism whereby instabilities are created within or on the surface of the liquid jet or sheet which promote the

formation of waves that subsequently break down into ligaments and then drops. Liquid viscosity delays the onset

of wave formation and opposes ligament breakup. Thus an increase in viscosity is always accompanied by an

53
increase in mean drop size. If, however, the bulk liquid is shattered almost instantaneously into droplets as soon as

it leaves the nozzle and becomes exposed to the atomizing air, then waves and ligaments have no time to form and

the delaying action normally exerted by viscosity has no opportunity to manifest itself. Under these conditions of

"prompt" atomization, viscosity and initial sheet thickness have little effect on mean drop size.

For both mechanisms, the mean drop size depends on atomizing air velocity, air/liquid ratio, and surface tension.

For the wavy-sheet mechanism, the mean drop size is also strongly dependent on the initial liquid dimensions (jet

diameter or sheet thickness), liquid viscosity, and air density. For the prompt mechanism, the mean drop size is much less dependent on the liquid jet diameter or sheet thickness, and completely independent of Uquid viscosity and

air density.

For any given atomizer, the dominant mode of atomization will be governed largely by its design features and

operating conditions. As discussed above, if the atomizing air and liquid are coflowing, then, unless the atomizing

air velocity is exceptionally high, the wavy-sheet mechanism will dictate the atomization process. On the other

hand, if the atomizing air impacts the liquid jet or sheet at an appreciable angle, or at a velocity which precludes

wave formation, then the prompt mode of atomization will dominate.

Under certain operating conditions it is possible that both mechanisms may play a role in the overall atomization

process. For example, atomizers which are normally characterized by the "prompt" mode of atomization will switch

to the "classical" mode of atomization if either the atomizing air velocity or the air/liquid ratio are reduced so far

below their normal operating values that sufficient time is available for wave formation to occur. At this point the

mean drop size will start to become dependent on initial jet diameter or sheet thickness, liquid viscosity, and air

density.

EFFERVESCENT ATOMIZATION

The atomizers discussed above have one important feature in common, namely, the bulk liquid to be atomized is

first transformed into a jet or sheet before being exposed to the atomizing air. In effervescent atomization, the

atomizing air is introduced directly into the bulk liquid at some point upstream of the nozzle discharge orifice. As

this air is not intended to impart kinetic energy to the liquid stream, it is injected at low velocity, so that the

pressure differential between the air and the liquid is very small, and only what is needed to p)ersuade die air to enter

the flowing liquid. The injected air forms bubbles which are conveyed by the liquid stream to the injection orifice,

downstream of which the bubbles 'explode' thereby shattering the surrounding liquid into small drops.

For this type of atomization, it is reasonable to assume that the mean drop sizes produced should be related to the

average distance between the bubbles in the plane of discharge from the injector orifice. If the average liquid

thickness between adjacent gas bubbles is denoted as tg, then, from simple geometric considerations, it can be

where dA is the average diameter of the air bubbles, ALR is the airAiquid ratio by mass, and pA and pL are the air

and liquid densities, respectively.

The value of C* in Eq. (7) depends on the manner in which the gas bubbles are distributed throughout the twophase medium. If the bubbles are spherical in shape, all the same size, and aligned in rows, the value of C* is unity. If the spherical bubbles are all the same size and so closely packed that the volume of liquid in the

intervening spaces is the minimum geometrically possible, then the value of C in Eq. (7) becomes 2.6. For this bubble packing geometry, the maximum airAiquid ratio for operation in the bubbly flow regime is given by

In practice, a wide range of bubble sizes may be generated so that small bubbles occupy the spaces between larger

bubbles. Analysis of Whitlow's [19] data for an effervescent atomizer operating over a range of pressures from 0.27

to 0.69 MPA yields a value for C* in Eq. (7) of 4.4. For this flow structure, the maximum airAiquid ratio for

operation in the bubbly flow regime is obtained as

shown that

(7)

ALRniax = 2.85 (pa/Pl) (8)

ALR,max = 4.80 (Pa/Pl) (9)

54
this relationship is illustrated in Fig. 9 for air bubbles in water at a temperature of 288K. This figure shows that,

when operating at low injection pressures, i.e. low gas densities, the bubbly flow regime is confined to relatively

low air^iquid mass ratios. Equation (9) indicates that replacing air with a gas of higher density would extend the

bubbly flow regime to a higher gas/Uquid ratio, and thereby improve atomization quality.

It is of interest to note in Eq. (7) that the average thickness of liquid between adjacent gas bubbles is independent

of the injector orifice diameter, which suggests that the mean drop size in the spray should also be independent of the

discharge orifice diameter. This is confirmed in the results obtained by Lefebvre, Wang, and Martin [20] and Roesler

[21]. Roesler's results for water at pressure levels of 0.069, 0.137, 0.275, and 0.551 MPA are shown in Fig. 10. They clearly demonstrate that mean drop size is fairly insensitive to changes in discharge orifice diameter.

Roesler also studied the effect of air injector geometry on mean drop size, using a series of porous tubes of

different porosity to inject atomizing air into flowing water. His results for aerator tubes having porosities of 20,

70, and 120 |j.m are shown in Fig 11. They illustrate that mean drop sizes are mainly dependent on operating

pressure and air/liquid ratio and are largely unaffected by the manner in which the atomizing air is introduced into the

flowing liquid stream

The fact that the atomizing performance of effervescent nozzles is largely independent of injector geometry is a most useful asset from a practical viewpoint As discussed above, the discharge orifice diameter has little effect on mean drop size, an increase in orifice diameter serving merely to increase the flow capacity of the nozzle for any

given operating pressure. Another important practical advantage of effervescent atomizers is their low discharge

coefficient. If the air and liquid flow through the discharge orifice at the same velocity, the discharge coefficient of

this orifice is given by

Cd = [1+ALR(pl/Pa)]-^ (10)

The variation of discharge coefficient with air/liquid ratio and operating pressure is illustrated in Fig. 12 for air- water mixtures. Typical values of discharge coefficient for conventional plain-orifice atomizers are between 0.62

and 0.8, depending on the length/diameter ratio of the discharge orifice. The values of discharge coefficient shown in

Fig. 12 are clearly much lower, especially for high air/liquid ratios, where Cd values fall to around 0.1.

Energy Considerations

The nature of the effervescent atomization process suggests that the fineness of atomization should be related to bubble energy. For a stream of air bubbles to convert a jet or sheet of liquid into a fine spray they must possess

enough energy to overcome the surface tension forces that hold the liquid together. Figure 13 shows SMD data

obtained by Roesler [21] over wide ranges of operating pressure and air/liquid ratio as a function of bubble energy

which is defined as

E = RT (ihA/riiL) In (Pa/Pl) (11)

At the high

air/liquid ratios and high pressures corresponding to the right side of Fig. 13, the bubbles have sufficient energy to break the liquid into small droplets through bubble explosions. However, at low pressures and/or low air/liquid

ratios, the effects of bubble expansion are greatly diminished. Under these conditions, the manner in which the

bubbles enhance atomization is mainly by squeezing the liquid into thinner shreds and ligaments as it leaves the

atomizer.

For effervescent atomizers operating at high pressures and/or high air/liquid ratios, the high bubble energy is manifested as a multiplicity of violent bubble explosions which rapidly disintegrate the liquid into a fine spray. For

these conditions of prompt atomization, Eq. (4), can be rewritten as SMD ^ L5_

do

^ ^ C PL Ta (APa/Pa)

CT(1 + 1/ALR)

(12)

A significant feature of this equation is that mean drop size is shown to be independent of liquid viscosity. In this

context it is of interest to note that Buckner et al. [22], who employed an aerated- liquid atomizer to examine the

effects of variations in liquid properties on mean drop size, found virtually no effect of viscosity on SMD for liquids

ranging in viscosity from 0.384 to 0.968 Ns/m. These results, shown in Fig. 14, provide further confirmation of

the fact that when atomization takes place very rapidly the effect of liquid viscosity on SMD is small.

55
CONCLUSIONS

1 . All the experimental evidence obtained on many different types of atomizers in which air is used as the

principal driving force for atomization points to air velocity as the principal factor governing mean drop

size. For low-viscosity liquids, the SMD is roughly inversely proportional to the atomizing air velocity.

Unless the liquid injection velocity is negligibly small in comparison with the atomizing air velocity, it is the relative velocity between the two streams that governs the mean drop size.

2. The other major factor governing the fineness of atomization is air/liquid ratio. Most correlations for

mean drop size embody the term [1 + 1/ALR]. This form of dependence on air/liquid ratio is consistent

with experimental observations in showing that the effect of ALR on SMD is large for small values of

ALR and diminishes with increasing ALR.

3. The influence of liquid jet diameter or sheet thickness on mean drop size depends on the mode of

atomization. For relatively slow mechanisms of atomization, involving flow instabilities and wave

formation, mean drop sizes are roughly proportional to the square root of the initial liquid dimension. If

atomization occurs very rapidly, the mean drop size is much less dependent on jet diameter or sheet

thickness.

4. If atomization proceeds fairly slowly, the effect of air density can be expressed as SMD a Pa , with

reported values of n varying between 0.3 and 0.6. For prompt atomization, Eqs. (4) and (5) predict no effect

of PA on SMD.

5. Mean drop sizes increase with increasing surface tension, regardless of the mode of atomization. For the

classical mechanism of jet or sheet breakup, SMD OL G^, where x usually lies between 0.25 and 0.5. For

prompt atomization, the dependence of mean drop size on surface tension is much stronger, especially at high values of Weber number where the mean drop size becomes direcdy proportional to the surface

tension.

6. The well-known adverse effect of an increase in liquid viscosity on atomization quality applies only to the

classical mechanisms of jet and sheet breakup. For prompt atomization, the mean drop size is largely

independent of viscosity.

REFERENCES

L A.H. Lefebvre, Airblast Atomization, Prog. Energv. Combust. Sci. vol. 6, pp. 233-261, (1980).

2. A.A. Rizkalla and A.H. Lefebvre, The Influence of Air and Liquid Properties on Airblast

Atomization, ASME J.Fluids Eng.. vol. 97, no. 3, pp. 316 - 320, (1975).

3. M. Surayi and A.H. Lefebvre, Drop-Size Measurements in Air-Assist Swirl Atomizer Sprays, Paper

CSS/CI 86-5D1, Spring Technical Meeting, Central States Section/ The Combustion Institute, Cleveland,

Ohio (May 1986).

4. A.H. Lefebvre and D. Miller, The Development of an Airblast Atomizer for Gas Turbine

Application, College of Aeronautics Report, Aero no. 193, Cranfield, England, (1966).

5. N.K. Rizk and A.H. Lefebvre, The Influence of Liquid Film Thickness on Airblast Atomization,

ASME J. Eng. Power . vol. 102, pp. 706 -710, (1980).

6. H.B. Squire, Investigation of the Instability of a Moving Liquid Film, Br. J. Appl. Phys.. vol.

4, pp. 167 -169, (1953).

7. R.P. Eraser, P. Eisenklam, N. Dombrowski, and D. Hasson, Drop Formation from Rapidly Moving

Sheets, J.AIChE. vol. 8, pp. 672-680, (1962).

8. T. Inamura and N. Nagai, The Relative Performance of Externally and Internally-Mixed Twin-Ruid

Atomizers, Proc. 3rd International Conference on Liquid Atomization and Sprays, pp. II C/2/1-11, London

(1985).

56
9. Q.S. Zhao and X.H. Gan, Effects of Geometrical Size and Configuration of Airblast Atomizers on

Atomizing Performance, ASME Paper 85-IGT-138, (1985).

10. J.E. Beck, A.H. Lefebvre, and T.R. Koblish, Airblast Atomization at Conditions of Low Air

Velocity, AIAA J. Propul. Power, in press.

11. J.E. Beck, A.H. Lefebvre, and T.R. Koblish, Liquid Sheet Disintegration in a Coflowing Air Stream,

Atomization and Sprays, vol. 1, no. 2, pp. 155-170, (1991).

12. T. Sattelmayer and S. Wittig, Internal Flow Effects in Prefilming Airblast Atomizers: Mechanisms of

Atomization and Droplet Spectra, ASME J. Eng. Gas Turbines and Power, vol. 108, pp. 465-472, (1986).

13. M. Aigner and S. Wittig, Swirl and Counterswirl Effects in Prefilming Airblast Atomizers, ASME Paper

87-GT-204, (1987).

14. N. Dombrowski and W.R. Johns, The Aerodynamic Instability and Disintegration of Viscous Liquid

Sheets, Chem. Eng. Sci. . vol. 18, pp. 203-214, (1963).

15. A.H. Lefebvre, Energy Considerations in Twin-Fluid Atomization, ASME J. Eng. Gas Turbines

and Power, in press.

16. A.H. Lefebvre, Atomization and Sprays. Hemisphere, Washington D.C., (1989).

17. J.B.Kennedy, High Weber Number SMD Correlations for Pressure Atomizers, ASME Paper 85- GT-37,

(1985).

18. S. Nukiyamo and Y. Tanasawa, Experiments on Atomization of Liquids in an Airstream. Trans.

Soc.Mech. Eng. Jpn. . vol. 5, pp.68-75, (1939).

19. J.D. Whitlow, An Experimental Investigation of Internal-Mixing, Twin-Fluid Atomization, M.S.M.E.

thesis, Purdue University, W.Lafayette, Indiana, (1990).

20. A.H. Lefebvre, X.F. Wang and C.A. Martin, Spray Characteristics of Aerated- Liquid Pressure

Atomizers, AIAA J. Propul. Power, vol. 4, 293-298, (1988).

21. T.C. Roesler, An Experimental Study of Aerated-Liquid Atomization, Ph.D. thesis, Purdue

University, W. Lafayette, Indiana, (1988).

22. H.N. Buckner, P.E. Sojka, and A.H. Lefebvre, Aerated Atomization of High Viscosity Newtonian

Liquids, Proc. 3rd Annual Conference on Liquid Atomization and Spray Systems. Irvine,

California, pp.56-61, (May 1989).

NOMENCLATURE

ALR air/liquid mass ratio

CD discharge coefficient

Dp prefilmer diameter, m

do initial jet diameter, m

dA air bubble diameter, m

E bubble energy, J/kg

LC characteristic dimension of atomizer, m

ih mass flow rate, kg/s

P pressure. Pa

Pb bubble pressure upstream of discharge orifice. Pa

AP pressure differential. Pa

R gas constant, J/kgK

SMD Sauter mean diameter,

,

T temperature, K

57
t initial sheet thickness, m

tg average distance between bubbles

U velocity, m/s

We Weber number

\i absolute viscosity, kg/ms

p density, kg/m^

a surface tension, kg/s-^

Subscripts

A air L liquid

LIQUID

Fig. 1 Schematic diagram of external-mixing, air-assist atomizer

E

to

50

25

AIR-ASSIST ATOMIZER

LIQUID - GASOLINE

P^ = O.IOI MPa (1 atmos.)

// ^^"'^

r \

^

\

AP^/P^. %

A 2

4

O 6

1 1 1

rhj^, kg/s

3 x|0-3

Fig. 2 Influence of liquid flow rate and air velocity on mean drop size

58
AIR

LIQUID

AIR

AIR

LIQUID

AIR

RIZK AND LEFEBVRE (1980)

(a)

BECK ET AL. (1989)

(b)

Fig. 3 Two different types of flat-sheet airblast atomizers

140

120

100

E

Q 80

CO

60

40

20

FLAT- SHEET AIRBLAST ATOMIZER

LIQUID - KEROSINE

U^ , m/s

50 100 150 200 250 300 350 400 450

INITIAL SHEET THICKNESS,

Fig. 4 Influence of initial sheet thickness on mean drop size. Data from Ref. 5

59
50

40

E

§ 30

20

Q

ATr>MI7C"D

LIQUID - KEROSINE

Uyi^, m/s

A 70

100

KJ 1 to

1 1 1 1

1 100 200

INITIAL SHEET THICKNESS, /xm

300

Fig. 5 Influence of initial sheet thickness on mean drop size. Data from Ref. 9

300

250

200

E

a

^ 150

100

50.

FLAT-SHEET AIRBLAST ATOMIZER

LIQUID - WATER

U^ , m/s

A 39.9

4ZI

O 51.5 ^ 595

100 200 300

INITIAL SHEET THICKNESS, /xm

400

Fig. 6 Influence of initial sheet thickness on mean drop size. Data from Ref. 10

60
250

E

^200

150

100,

Fig. 7

20

FLAT-SHEET AIRBLAST ATOMIZER

LIQUIDS - WATER AND GLYCERINE- WATER

VISCOSITY. mN /sm

13.60

O 1.04

30 40 50

RELATIVE VELOCITY, m/s

60

Influence of liquid viscosity on mean drop size. Data from Ref. 10

200

150

E

Q 100

50

0

FLAT-SHEET AIRBLAST ATOMIZER

LIQUIDS - KEROSINE AND WATER

SURFACE TENSION, N/m

0.0746

o 0.0285

25 30 35 40 45 50

RELATIVE VELOCITY, m/s

55 60

Fig. 8 Influence of surface tension on mean drop size. Data from Ref. 10

61
V)<

>-m

^ 0.03

a:

AIR BUBBLES

IN WATER

T = 288 K

1.0

Fig. 9

Q2 04 06 08"

PRESSURE, MPa

Maximum air/water ratio for effervescent atomization

E

a.

CO

Pq=QI70 MPa

LIQUID - WATER

ORIFICE DIA. . mm

O 1-5 2.0 ^ 2.5

P0=O238 MPa

Pg =0.376 MPa

Pg =0.652 MPa

-O. 0.005 0.025

Fig. 10

0.010 0.015 0.020

AIR /LIQUID RATIO

Influence of final discharge orifice diameter on mean drop size. Data from

Ref 21

62
360 -

320-

280-

240-

200-

4

280-

240-

200-

160-

K

A

(H) o

LIQUID - WATER

ORIFICE 0IA =l.5mm

AERATOR POROSITY. ;iin O 20

D 70 A 120 ^ Pg =0.170 MPa

9

o

200 O Pg = 0238 MPa

160 -

'

120-

80-

leo

120

80

40

3

Pg = 0 376 MPo

Or.

8

^0

dB

eg Pg=0652 MPO

_l_ 0005 0010 0015

AIR/LIOUID RATO

0020

Fig. 1 1 Influence of aerator tube porosity on mean drop size. Data from Ref. 21

0.6

y 0.5

^ 0.4

CJ

UJg 0.3

<

X

o

0.2

o

0.1

0,

LIQUID - WATER

\ ^'^'^--^^^ Pg ,MPa

^v^^ ^ 1.0

^0.5

1

0.2

1 1 1 0.005 0010 0.015

AIR/LIQUID RATIO

0020 0.025

Fig. 12 Influence of bubble pressure and airAiquid ratio on discharge coefficient

63
0.04 0.06

BUBBLE ENERGY

0.10

Fig. 13 Relationship between bubble energy and mean drop size

50

40

30

E

20

10

0

0

AERATED- LIQUID ATOMIZER

LIQUIDS- GLYCERINE -WATER MIXTURES

VISCOSITY, Ns/m

o 0384

0520

^ 0968 P^ = l.48 MPa

Ol 02

AIR/ LIQUID RATIO

03 04

Fig. 14 Influence of air/liquid ratio on mean drop size for liquids of high viscosity.

Data from Ref. 22

64
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 1

RELAXATION OF A POLYDISPERSE SPRAY UNDERGOING VAPORIZATION

AND COALESCENCE IN A FLOW FIELD BEHIND A NORMAL SHOCK WAVE:

A SECTIONAL APPROACH

S. Zahavi and Y. Tambour

Faculty of Aerospace Engineering

Technion - Israel Institute of Technology

Haifa, Israel

ABSTRACT

A gas flow field behind a normal shock wave inflicts a step function in velocity, temperature and

pressure upon a polydisperse spray suspended in the flow. As a result, there are large differences in

velocity and temperature between the carrying gas and the droplets of the spray immediately after the

shock. These differences gradually decrease in a relaxation zone which ends at a distance, where the so

called "far-field" begins, at which the droplets eventually travel at a velocity which is close to the

host fluid velocity. Thus, the investigation of the relaxation zone may be regarded as the analysis of

a "near-field" problem in contradistinction to "far-field" problems which have been analyzed by

Greenberg, Tambour and their coworkers employing a sectional approach. Treatment of a "near-field"

problem via a sectional approach requires derivation of new generalized sectional-equations which

extend the previous sectional-equations which were derived by Tambour for "far-field" problems. Such

new generalized sectional-equations are derived in the present paper. They rigorously account for the

heritage of momentum and enthalpy that droplets carry with them as they "move" from one size section to another due to evaporation and coalescence processes, e.g. a large droplet formed by coalescence (of

two small droplets) will differ from the originally large droplets since the properties of the newly

formed droplet depend on the properties of the small droplets which participate in the coalescence

process, and generally in the "near-field", small droplets differ in velocity and enthalpy from large ones. These effects, as well as the influence of the velocity lag between droplets and the host fluid,

between droplets of various sizes, and between droplets of the same size on the evolution in pointwise

droplet size distributions along the axial direction are presented here. The computed results also

include changes of velocities and temperatures of each droplet size-section in the relaxation zone.

INTRODUCTION

The flow of a polydisperse liquid spray suspended in a gas is of importance to a wide range of

technologies in the chemical industry, in jet engine main combustion chamber and afterburner design,

and in liquid fuel rocket propulsion. In general, when the spray is injected, the two phases (the

liquid spray and the host gas) are not initially in equilibrium. There are differences in velocity and

temperature, not only between the two phases, but also between droplets of various sizes. Equilibrium

tends to be achieved by relatively slow processes of drag, heat and mass transfer as the flow proceeds

through a relaxation zone or a "near-field" zone. Since there are velocity differences between

droplets, coalescence resulting from collisions between droplets, may occur. In the light of the above, our specific interest in the present paper is the analysis of the relaxation of a polydisperse spray

undergoing simultaneous vaporization and coalescence in a flow field in which the nonequilibrium state

is brought about by a normal shock wave.

Early work which dealt with the relaxation of solid particles in a gas flow behind a shock wave

was presented in the late fifties by Hoenig [1] and by Carrier [2]. Hoenig [1] discussed the variation

65
of the velocity of a single solid particle behind a shock wave. Carrier [2] formulated the general

problem for solid particles that are numerous enough to affect the gas flow behind the shock wave. The

case of solid particles was also studied in the sixties by Kriebel [3] who analyzed three particle

sizes, and by Rudinger [4,5] who analyzed various shock strengths, drag laws and specific heats.

Rudinger pointed out that the velocities and temperatures do not always change monotonically towards

their equilibrium values.

Mass transfer between the two phases was considered by Lu and Chiu [6] for the case of water

droplets in air. Panton & Oppenheim [7] considered the case of two droplet size groups and demonstrated

the interesting possibility of a situation where some of the droplets diminish by vaporization while

others grow by condensation. Nayfeh [8] studied a gas containing ablating particles and Evans and

Mackie [9] modified his calculations by allowing downstream variation of particle temperatures and by

the use of a diffusive evaporation rate. Later, in 1975, Narkis and Gal-Or [10] pointed out that for a given Mach number there is a certain particle flow rate for which a minimum equilibrium velocity is achieved.

More recently, Ben-Dor and Igra [11,12], analyzed the relaxation zone behind normal shock waves in a dusty reacting gas. An analysis of normal shock waves in a carbon particle-laden oxygen gas was

presented by Alperin, Igra and Ben-Dor [13]. Although not dealing with a flow behind a shock wave, a

pioneering theoretical study of what can be regarded as a "near-field" problem, was recently presented

by Bellan and Harstad [39]. They analyzed dense and dilute clusters of drops evaporating in large,

coherent vortices.

Detailed formulations of the governing equations for the two phases can be found in the

literature, e.g. see Carrier [2], Soo [14,15], Kliegel [16,17], Marble [18-19] and Pai [20]. However,

since our objective here is to analyze a polydisperse spray undergoing simultaneous evaporation and

coalescence processes, we find a "sectional" approach [21-33] to be useful for handling this problem.

In the sectional approach, the spray is divided into size groups (size-sections), where for each size

group a sectional mass balance (and in the present study also a momentum and energy) equation is derived. This enables one to deal with a fairly small number of conservation equations for the analysis

of a polydisperse spray of a wide range of sizes. Another advantage of the sectional approach lies in

the ability to rigorously accoimt for intra- and inter-sectional mass (and in the present study also

momentum and energy) transfer for simultaneous evaporation and coalescence processes.

Thus far, the sectional approach has been used [21-25,27-33] for the analysis of "far-field"

zones, where essentially most of the droplets travel at a velocity close to the host fluid velocity

field. Theoretical simulations have been carried out for (i) "far-field" downstream changes in integral

droplet size distributions of evaporating sprays in "cold" and "heated" environments [23] and for (ii)

"far-field" downstream evolution in pointwise droplet size distributions of polydisperse sprays

undergoing simultaneous evaporation and coalescence processes [24]. The good agreement that has been

observed comparing these theoretical simulations [23,24] with independently reported experimental data

[34,35] has demonstrated the simplicity and reliability of the sectional approach.

As to "near-field" problems, rapid developments in measurement techniques supply plenty of

experimental data on pointwise and integral droplet size distributions of polydisperse sprays in the

"near-field" [36-38]. Such data can serve as initial conditions for simulations and also for verifying

theoretical predictions.

In light of the above, the purpose of the present study is to extend the mass balance

sectional-equations which were previously derived for the analysis of "far-field" problems [21-24] and

present new generalized sectional equations which account for the intra- and inter-sectional transfer

of momentum, kinetic-energy and enthalpy. The derivation of these equations has been presented

elsewhere [40]. These generalized equations will be employed here in order to analyze a "near-field"

66
problem, that is, the relaxation zone of a polydisperse spray undergoing evaporation and coalescence

processes behind a normal shock wave.

RESULTS AND DISCUSSION

Computations were carried out for a polydisperse spray suspended homogeneously in a

one-dimensional gas flow at a Mach number of M=1.5, and under conditions of equilibrium of velocity and

temperature, between the two phases (the liquid spray and the host fluid). Due to a normal shock wave

at x=0, the host fluid undergoes a step change in velocity, temperature and pressure. We assume that

immediately after the shock wave, at x=0 , the volume concentration of the spray z^, and the size

distribution of the spray, and the mass fraction of vapors Xq» given. Based on these initial

conditions, we analyze the evolution in spray size distribution and the variation of velocity and

temperature of each size group (section) of the spray along the axial direction in the relaxation zone. These changes are not only coupled to the variation of velocity and temperature of the host fluid in

the relaxation zone due to the interaction between the two phases, but also intercoupled due to mass,

momentum and energy transfer between droplets of various sections.

The spray size distribution is divided into 10 sections (size groups). The smallest section is 0-8

monomers and the limits of the largest section are 5832-8000 monomers. If one denotes the diameter of a

single monomer by d^, then, section No. 1 includes droplets of a diameter up to 2d^. Hie "width" of all

other sections is also 2d^ (in terms of droplet diameters), so that droplets of a diameter which lies

within the range of 18d to 20d , are in section No. 10. Hence, the ratio of the largest droplet in

0 0

w r

section 10 to the largest droplet in section 1 is 10:1 in diameters and 1000:1 in volume (or mass).

We begin with the analysis of a trivial case in which there is neither coalescence nor

evaporation. The computed results of this case are shown in Figs. 1 and 2, and will serve as a comparison for more complex cases. The variation of velocity in the relaxation zone is shown in Fig. 1. The gas loses almost half of its original velocity as it passes through the shock. Then, its velocity

continues to decrease in the relaxation zone due to dissipation of kinetic-energy which results in an

increase of the internal energy of the gas, as indicated by the increase in gas temperature in the

first part of the relaxation zone (see Fig. 2). (Afterwards, the gas temperature decreases as the gas

heats the droplets.) As to the variation of the velocity of each size section of the spray, as expected, the small droplets (e.g. section No. 1) adjust to the gas velocity within a shorter distance

than the large droplets.

We now return to analyze the results presented in Fig. 2. Since immediately after the shock the

gas is hotter than the droplets, the droplets are heated along the axial direction till thermal

equilibrium is gained. The thermal relaxation zone is wider than the momentum relaxation zone (compare

Figs. 1 and 2) due to the difference in characteristic time scales between heat transfer and momentum

transfer. It is also interesting to note that large droplets (e.g. sections 2 to 10) increase their

temperature gradually, while the small droplets (sections 1 and 2) exhibit a temperature overshoot.

This behavior is due to the fact that small droplets are heated faster than large ones, so when the gas

is "hot", they reach high temperatures and then, the gas cools down at a faster rate. There is an

interesting region in which the temperature of these small droplets is locally higher than that of the

surrounding gas. However, in this region, the slope of temperature variation always points downwards.

Next, coalescence effects are analyzed and compared with the previously described case in which

neither coalescence nor evaporation were assumed. In a polydisperse spray such as considered here,

collisions between droplets are expected due to the velocity differences between droplets of various

sizes. Collision rates depend on the characteristic distance between droplets (which is a function of

67
number concentration of droplets per unit volume, or the volume concentration of the spray, z), the

relative velocities between droplets, the collision cross section, and the characteristic time of

droplets in the relaxation zone. We assume here that all collisions result in coalescence.

We focus here on three coalescence rates: (i) "Low", C_,_=0.0002, (ii) "Medium", C_,^=0.(X)5, and

(iii) "High", Cp^=0.02. The variation of velocity of each size-section of the spray and the velocity of

the host fluid, in the relaxation zone, are shown in Figs. 3a(i, ii and iii) for the above three

coalescence rates. Comparison of Fig. 1 (No coalescence) with Fig. 3a-i (Low coalescence) indicates

that the effects of "Low" coalescence rate, on the velocity profiles are hardly noticeable. However, at higher coalescence rates (see Figs. 3a-ii and 3a-iii), the velocity profiles of the various spray

sections become more crowded and the relaxation zone becomes shorter. This behavior is due to the fact

that coalescence serves as a "mixing" process for the momentum of droplets, in which the velocity of

each newly formed droplet is constructed on the basis of the momentums of the original droplets that

participate in the coalescence process.

The same arguments hold for explaining the results that show the effects of coalescence on the

temperature profiles of the various spray-sections. Coalescence also serves as a mixing process for

thermal energy and thus shortens the relaxation zone [compare Figs. 4 (i4i4ii) with Fig. 2].

The results described above are based on solutions of the "near-field" equations that have been

derived in the present paper. These equations rigorously account for the heritage of momentum and

enthalpy that droplets carry with them as they "move" from one size-section to another due to coalescence (and evaporation). To demonstrate the importance of this rigorous treatment, for comparison

we present in Figs. 3b (i, ii, iii) and 4b (i, ii, iii), velocity and temperature profiles obtained by

solving a set of more simplified sectional-equations in which intra- and intersectional mass transfer

due to coalescence is accoimted for, but bound to the simplifying assumption that inter-sectional

momentum and energy transfer are negligible (as in the "far-field" equations). Thus, newly formed

droplets adopt the momentum and energy of the new section which they join without altering its average

momentum and energy.

One can clearly see that the rigorous derivation of the "near-field" sectional-equations is necessary since otherwise solutions may show erroneous results such as presented in Figs. 3b and 4b.

For example, the second peak in temperature (see Figs. 4b, i and ii) is erroneous. It is explained as follows. Small droplets have a higher temperature than large ones, so when they coalesce the newly

formed large droplets have a higher temperature than the originally large droplets. Failing to consider

this effect by assuming that the newly formed large droplets immediately adopt the energy state of the

originally large droplets results in an energy difference which, due to energy conservation in the

system as a whole, is added erroneously to the host gas.

The variations of the liquid-phase mass fiow rate along the axial direction are shown in Fig. 5

for each of the spray size-sections. Variations of the spray volume fractions are presented in Fig. 6. Both figures are for the "medium" coalescence rate (Cp^=0.()05) and both show the increase in mass flow

rate and in volume concentration of droplets in the upper sections due to the creation of new droplets

by the coalescence process. It should be emphasized that due to the overall mass conservation, section

No. 10 actually also includes all droplets larger in size than the upper limit of section No. 10.

The integral volume fraction of the liquid phase, z, is shown in Fig. 7 for various coalescence

rates. All profiles asymptotically approach the same value, and the increase in spray concentration, as one marches downstream, is due to the decrease in droplet velocities see Figs. 3a.

The effects of heat transfer on the variation of temperature in the relaxation zone are shown in

68
Fig. 8. These results indicate that a shorter relaxation distance is obtained for droplets of a high

initial temperature, which is closer to their boiling point temperature (compare Fig. 8-i with Fig.

8-ii). If a more accurate heat-transfer model is used, e.g. with Ranz-Marshall convective correlation

(see Fig. 8-a), temperature profiles of the various spray sections become more crowded due to the

higher convective heat transfer rate.

The latter results are for an evaporating spray without coalescence. When simultaneous evaporation

and coalescence occur, the general behavior remains the same (see Fig. 9) since on the one hand

coalescence hinders vaporization (due to the decrease in the surface area of the droplets), but on the

other hand, coalescence serves as a "mixing" process and causes the profiles of the various spray

sections to become more crowded.

CONCLUSIONS

To summarize, new generalized "near-field" spray section-equations have been derived in which

intra- and inter-sectional transfer of mass, linear-momentum, kinetic-energy and enthalpy has been

rigorously accoimted for. These equations have been employed here to analyze the behavior of a

polydisperse spray undergoing evaporation and coalescence processes in a relaxation zone behind a normal shock wave.

REFERENCES

1. Hoenig, S.A. (1957), "Acceleration of Ehist Particles by Shock Waves", J. of Applied Physics, Vol.

28, No. 10, pp. 1218-1219.

2. Carrier, GP. (1959), "Shock Waves in a Dusty Gas", /. of Fluid Mechanics, Vol. 4, No. 4, pp.

376-382.

3. Kriebel, A.R. (1964), "Analysis of Normal Shock Waves in Particle Laden Gas", /. of Basic Eng.,

Vol. 86, No. 4, pp. 655-665.

4. Rudinger, G. (1964), "Some Properties of Shock Relaxation in Gas Flows Carrying Small Particles",

Physics of Fluids, Vol. 7, No. 4, pp. 658-663.

5. Rudinger, G. (1965), "Some Effects of Finite Particle Volume on the Dynamics of Gas-Particle

Mixtures", AIAA J., Vol. 3, No. 7, pp. 1217-1222.

6. Lu, H.Y. and Chiu, H.H. (1966), "Dynamics of Gases Containing Evaporable Liquids under a Normal

Shock", AIAA J., Vol. 4, No. 6, pp. 1008-1011.

7. Panton, R. and Oppenheim, A.K. (1968), "Shock Relaxation in a Particle-Gas Mixture with Mass

Transfer Between Phases", AIAA J., Vol. 6, No. 11, pp. 2071-2077.

8. Nayfeh, A.H. (1966), "Shock Wave Structure in a Gas Containing Ablating Particles", Physics of

Fluids, Vol. 9, No. 12, pp. 2351-2356.

9. Evans, PJ. and Mackie, J.K. (1971), "Shock Waves Interaction with an Evaporating Aerosol",

Physics of Fluids, Vol. 14, No. 3, pp. 539-541.

10. Narkis, Y. and Gal-Or, B. (1975), "Two - Phase Flow Through Normal Shock Waves", J. of Fluid

Engineering (ASME), Vol. 97, pp. 361-365.

11. Ben-Dor, G. and Igra, O. (1982), "The Relaxation Zone Behind Normal Shock Waves in Reacting Dusty

Gas. Part 1. Monatomic Gases", /. Plasma Phys., Vol. 27, pp. 377-395.

12. Igra, O. and Ben-Dor, G. (1984), "The Relaxation Zone Behind Normal Shock Waves in a Dusty

Reacting Gas. Part 2. Diatomic Gas", J. Plasma Phys., Vol. 31, pp. 115-140.

13. Elperin, I., Igra, O. and Ben-Dor, G. (1986), "Analysis of Normal Shock Waves in a Carbon

Particle-Laden Oxygen Gas", /. of Fluids Eng., Trans. ASME, Ser. D, 86, pp. 655-663.

14. Soo, SL. (1961), "Gas Dynamics Processes Involving Suspended Solids", AIChE J., Vol. 7, No. 3,

pp. 384.

15. Soo, SL. (1967), Fluid Dynamics of Multiphase Systems. Blaisdell Publishing Co., Waltham.

16. Kliegel, J.R. and Nickerson, G.R. (1962), "Flow of Gas-Particles Mixtures in Axially Symmetric

Nozzles", In Progress in Astronautics and Rocketry, Vol. 6, Academic Press, New York, pp. 173-194.

69
17. Kliegel, J.R. (1963), "Gas Particle Nozzle Flow", Ninth Symposium (International) on Combustion,

Academic Press, New York, pp. 811-826.

18. Marble, FE. (1963), "Dynamics of a Gas Containing Small Solid Particles", Fifth AGARD Combustion

and Propulsion Colloquium, Pergamon Press, London, pp. 175-215.

19. Marble, F.E. (1964), "Mechanism of Particles Collision in the One-Dimensional Dynamics of

Gas-Particle Mixtures", Physics of Fluids, Vol. 7, No. 8, pp. 1270-1282.

20. Pai, S.I. (1977), Two-Phase Flows, Oswatitisch.

21. Tambour, Y. (1980), "A Sectional Model for Evaporation and Combustion of Sprays of Liquid Fuels",

Israel J. Tech., Vol. 18, pp. 47-56.

22. Tambour, Y. (1984), "Vaporization of Polydisperse Fuel Sprays in a Laminar Boundary Layer", Comb.

Flame, Vol. 58, pp. 103-114.

23. Tambour, Y. (1985), "A Lagrangian Sectional Approach for Simulating Droplet Size Distribution of

Vaporizing Fuel Sprays in a Turbulent Jet", Comb. Flame, Vol. 60, pp. 15-28.

24. Tambour, Y. (1985), "Coalescence of Vaporizing Kerosene Fuel Sprays in a Turbulent Jet",

Atomization and Spray Technology, Vol. 1, pp. 125-146. See also: "Simulation of Coalescence and

Vaporization of Kerosene Fuel Sprays in Turbulence Jet: A Sectional Approach", AIAA/ASME/ASEE 21st

Joint Propulsion Conference, Monterey, 1985. AIAA Paper No. 85-1315.

25. Tambour, Y., Greenberg, J.b. and Albagli, D. (1985), "Vaporisation and Coalescence of Sprays in

3-D Axisymmetric Stagnation Flow", ICLASS-85, 3rd International Conference on Liquid Atomisation

and Spray Systems, England, 1985.

26. Avital, G., Greenberg, J.B., Tambour, Y. and Timnat, Y.M. (1989), "Spray Dynamics and Fuel Vapor

Distributions in a Spinning Spherical Combustion Chamber", 12th International Colloquium on Dynamics of Explosions and Reactive Systems, Ann Arbor, Michigan.

27. Greenberg, J.B. and Tambour, Y. (1986), "Far Field Coalescence Effects in Polydisperse Spray Jet Diffusion Flames", Twenty First Symposium (International) on Combustion, The Combustion Institute,

Philadelphia, pp. 655-663.

28. Greenberg, J.B., Albagli, D. and Tambour, Y. (1986), "An Opposed-Jet Quasi- Monodisperse Spray

Diffusion Flame", Comb. Sci. Tech., Vol. 50, Nos. 4-6, pp. 255-270.

29. Tambour, Y. and Greenberg, J.B. (1988), "Fuel-Injector Far Field Temperature Distributions

Resulting from Coalescence, Vaporization and Combustion of a Polydisperse Spray", ICLASS-88, 4th

International Conference on Liquid Atomisation and Spray Systems, Japan.

30. Greenberg, J.B. (1989), "The Burke-Schumann Diffusion Flame Revisited-With Fuel Spray Injection",

Comb. Flame, Vol. 77, pp. 229-240.

31. Silverman, I., Greenberg, J.B. and Tambour, Y. (1990), "Asymptotic Analysis of a Premixed

Polydisperse Spray Flame", SIAM J. on Appl. Mathematics, in press. 32. Silverman, I., Greenberg, J.B. and Tambour, Y. (1991), "Solutions of Polydisperse Spray

Sectional-Equations via a Multiple-Scale Approach: An Analysis of a Premixed Polydisperse Spray

Flame", submitted to ICLASS-91.

33. Tambour, Y., Silverman, I. and Greenberg, J.B. (1991), "Effects of Initial Droplet Size

Distribution on Polydisperse Jet Far-Field Di^sion Flames", submitted to ICLASS-91.

34. Yule, AJ., Seng, C.Ah., Felton, P.G., Ungut, A. and Chigier, N.A. (1982), "A Study of Vaporizing

Fuel Sprays by Laser Techniques", Combustion & Flame, Vol. 44, pp. 71-84.

35. Yule, AJ., Ereat, P.R. and Ungut, A. (1983), "Droplet Sizes and Velocities in Vaporizing Sprays",

Comb. Flame, Vol. 54, pp. 15-22.

36. Bachalo, W.D., Brena de la Rosa, A. and Rudoff, R.C. (1988), "Advances in Diagnostics for Complete

Spray Characterization", ICLASS-88, pp. 287-295.

37. Edwards, C.F. and Rudoff, R.C. (1990), "Structure of a Swirl-Stabilized Spray Flame by Imaging,

Laser Doppler Velocimetry and Phase Doppler Anemometry", 23rd Symp. Combust., Orleans, France.

38. Presser, C, Gupta, A.K., Avedisian, C.T. and Semerjian, H.G. (1990), "Fuel Property Effects on

the Structure of Spray Flames", 23rd Symp. on Combustion, Orleans, France.

39. Bellan, J. and Harstad, K. (1990), "The Dynamics of Dense and Dilute Clusters of I>rops Evaporating

in Large, Coherent Vortices", 23rd Symp. on Combustion, Orleans, France.

40. Zahavi, S. and Tambour, Y., "Relaxation of a Polydisperse Spray Behind a Normal Shock Wave: A

Sectional Approach", to be submitted to Atomization and Sprays.

70
Fig. 1: Variation of gas-phase velocity, and

velocity of each size-section of the spray in

the relaxation zone. No evaporation, No

coalescence.

Fig. 2: Variation of gas-phase temperature,

and temperature of each size-section of the

spray in the relaxation zone. No evaporation.

No coalescence.

; \^ (111)

r X

Fig. 4: Variation of gas-phase temperature,

and temperature of each size-section of the

spray in the relaxation zone, for three

coalescence rates: (i) "Low", Cp^=0.0002, (ii)

"Medium" Cp^=0.005, and (iii) "High",

Cp^=0.02.

(a) Solutions of rigorous "near-field"

sectional equations.

(b) Neglecting inter-sectional momentum and

energy transfer.

Fig. 3: Variation of gas-phase velocity, and

velocity of each size-section of the spray in

the relaxation zone, for three coalescence

rates: (i) "Low", C„„=0.0002, (ii) "Medium",

Cp^=0.005, and (iii) "High", Cp^=0.02.

(a) Solutions of rigorous "near-field"

sectional-equations

.

(b) Neglecting inter-sectional momentum and

energy transfer.

71
Fig. 5: Variations of the liquid-phase mass

flow rate along the axial direction for each

of the spray size-sections. "Medium"

coalescence rate Cp^=0.(X)5 (No evaporation).

Fig. 6: Variations of the liquid-phase volume

fractions along the axial direction for each

of the spray size-sections. "Medium"

coalescence rate Cp^=0.(X)5 (No evaporation).

TOTRL PRRTICLX:S VOLUME FRACTION

Fig. 7: Profiles of the integral volume fraction of the liquid phase (z) in the relaxation zone for

various coalescence rates.

Fig. 8: Effects of heat transfer on the

variation of gas-phase temperature, and

temperature of each size-section of the spray

along the axial direction (evaporation

without coalescence). Initial droplets

temperature is: (i) 0.77 of their boiling point

temperature, (ii) 0.91 of their boiling point

temperature.

(a) with Ranz-Marshall convective correlation.

(b) without Ranz-Marshall convective correlation.

scciiomL vaociiir nao scciiona ictvcfmiuRC nao

" " " X

" " " i-

-

Fig. 9. Effects of simultaneous evaporation and

coalescence rates on variation of (a) velocities

and (b) temperatures in the relaxation zone, for

two coalescence rates (i) C^„=0.(X)1 and (ii)

72
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 2

EFFECT OF INTERACTION ON THE RATES OF EVAPORATION OF A

GROUP OF DROPS AND PREDICTION OF VOID FRACTION

I. Taniguchi and K. Asano

Department of Chemical Engineering

Tokyo Institute of Technology

Tokyo, Japan

ABSTRACT

Measurements were made for the effect of adjacent solid spheres with

arbitrary arrangement on the rates of evaporation of a water and an octane

drop for Rep^=72-234,Sc=0.59-1.61, diameter ratios Dp/Dy^=0.57-0.89 and

dimensionless distance L/D^=1.12-9.44 ,W/Dy^=0-1 . 02 . New correlations for the

effect of solid spheres on the rates of evaporation were proposed. By use of

the correlations, prediction of void function was made and compared with the

experimental data for gas-solid fluidized bed.

INTRODUCTION

Prediction of void function, which relates heat and mass transfer rates

of a group of drops or of particles with that for a single drop or particle,

is important for design of spray columns or fluidized bed- or fixed-bed

reactors. Although in the past decades many theoretical or experimental

approaches (2) , (5) , (6) , (9) , (10) to motion ,heat and mass transfer of a group

of drops or particles have been made , few approach has ever been made to the

prediction of void function. In our previous papers, some experimental

(1),(13) and theoretical (11), (12) approaches to the effect of interaction on

the drag coefficients and diffusion fluxes of two coaxially arranged drops

were made under low- to high mass flux conditions and for intermediate

Reynolds number ranges. The purpose of the present work is to make an

experimental approach to the effect of a group of solid spheres on the rates

of evaporation of a drop and to develop a new method for prediction of void

function for low to intermediate Reynolds number ranges.

EXPERIMENTAL APPARATUS AND PROCEDURES

Experimental Apparatus

Figure 1 shows a schematic diagram of an experimental apparatus. The test

section was a brass-made rectangular duct of 70x70mm in cross section and

166mm in length with two large side view windows. In the middle of the test

section, a 0.5mm-outer diameter stainless steel tubes for liquid supply was

mounted by a three-dimensional fine adjustment traversing mechanism, through

which liquid to an evaporating pendant drop was supplied by a microfeeder73
operated hypodermic syringe. A

0 . 05mm-outer diameter Chrorael-Alumel

thermocouple was mounted by a twodimensional fine adjustment traversing mechanism for measurements of

drop temperature.

Measurement

Rates of evaporation of a pendant drop were taken equal to

equilibrium liquid flow rates

supplied by a microfeeder, of which

flow rates were so adjusted that the

diameter of the drop did not change

appreciably for a given time

interval

.

Distance between the center of

a drop and a solid sphere or distance between solid spheres were measured by a cathetometer with an

accuracy of 1/lOOmm. Diameters of a pendant drop were also measured by

the same cathetometer, from which

surface area of the drops were calculated by assuming that the shape of

the drop was symmetrical with

respect to vertical axis(3).

The drop temperatures were

measured by a 0.05mm-outer diameter

Chrorael-Alumel thermocouple mounted

on a two-dimensional fine adjustment

traversing mechanism, from which the

surface temperatures of the evaporating drop were estimated by assuming

that the temperature distribution

inside the drop is uniform (3).

Surface concentrations of the vapor

of an evaporating drop were estimated from saturated vapor pressures

of the pure liquid at the surface

temperatures

.

Ranges of variables were shown

in Table 1.

EXPERIMENTAL RESULTS

A '. Test section

B : Blower

C : Silica gel dehumidifier

01, 2: Heater

E : Divergent-convergent

nozzle with calming grid

F : Microfeeder

Fl : Orifice flow meter

G : Cathetometer

TCI, 2 '.Temperature controller

TI ; Thermocouple for

1:bull< flow

2: temperature profile

3: liquid drop

Fig. 1 Schematic diagram of an experimental apparatus.

Table 1 Ranges of experimental variables

System Water-Air Octane-Air

Dp [mm] 1.92-2.66 1.72-2.28

D^ [mm] 3.0 3.0

L [ram] 0.80-28.32 4.95-24.21

W [ram] 0-4.11 0-1.80

Too [K] 343-371 335-337

Uco [ra/s] 0.31-1.40 0.80-1.17

RepA [-] 73-234 138-232

Sc [-] 0.59 1.61

L/D^ [-] 1.12-9.44 1.65-8.07

Dp/D^[-] 0.64-0.89 0.57-0.76

W/D^ [-] 0-1.02 0-0.60

Effect of Two and Four Upstreara Spheres

Figure 2 shows the effect of two solid spheres of equal diameters

arranged symraetrically around center axis of the drop on the diffusion

fluxes of a water drop. The ordinate is the diffusion fluxes normalized by the

ones for a single drop (8) and the abscissa is the dimensionless distance

L/D^. The parameter ,W/D^, is the dimensionless distance between the center

axis and the solid sphere. The effect of L/D^ on the diffusion fluxes

becomes considerable as W/DAdecrease . By considering our previous numerical

74
2.0

1.0

0.8

0.6

0.5

O.A

0.3

0.2

Water - Air System

I

Sphere,

Sphere

2W—^DropL

Key W/Da • 1.0 2 0.9 3

0. 6 4

0. 5 5

0.6 0.8 1 2 3

L/D.

A 5 6

C-]

8 10

Fig. 2 Effect of two solid spheres

arranged symmetrically around

center axis of a drop.

VI 3

I

if)

3

I

if)

2.0

1.0

0.8

0.6

O.A

0.2

0.1

Water-Air System

I

I I I I Present model

-I—L. I I I 20V.

2 0'/.

Solid Sphere

RepAH 1A3

Dp/DaH 0.73

W/DaH 0.71

2 A 6 8 10 20 30

L/Da [-]

Fig. 3 Effect of four solid spheres

arranged symmetrically around

center axis of a drop.

correlation for the effect of interaction on the rates of mass transfer of

two coaxially arranged solid spheres (11), (12);

-0.26

f= l/{ l-H).77(Dp/D^)~°-^^(L/D^)"^-^^'^®PA } (1)

the present data were well correlated by the following equation;

Sh(l - a)g)/[Sh(l - =

(2)

_Q 2^

= l/{l-K).77(Dp/D^)"°-2^(L/D^)"^-^^*^^PA* (3)

where (1-W/D^)-^*^ represents the effect of eccentric arrangement of solid

spheres

.

Figure 3 shows the effect of four solid spheres of equal diameters

arranged symmetrically around the center axis of the drop on the diffusion

fluxes of a water drop. The solid line in the figure represents the

diffusion fluxes estimated by the following equation:

Sh(l - ojg)/[Sh(l - 0)^)]^ =

(4)

where f-j^ is estimated by use of Eq.(3). Good agreement is observed

between the data and the theory.

Figure 4 shows the effect of two solid spheres of arbitrary

arrangement on the rates of evaporation of a water drop, where the first

solid sphere(the one near to the drop) is arranged eccentrically to the

center axis and the second one is arranged coaxially with the drop.

The dot-dash line in the figure represents the effect of the first solid

sphere calculated by Eq.(3). Deviation from the dot-dash line may indicate

the effect of the interaction due to the presence of the second sphere. If

we assume that the effect of interaction with two solid spheres of

75
1.0

0.8

3*" 0.6

Solid

Sphere

Drop

Water-Air System

f1

g QA

I

02

Key RepA [-]

1.00 •

113

L/DaH 1.33 125

02 QA 0.6 0.8 1.0

Fig. 4 Effect of two solid spheres

with arbitrary arrangement.

3

I ^—

CO

3""

I

> Q8

W,/Da C-]

Fig. 5 Effect of second sphere in the

presence of the first solid

sphere.

arbitrary arrangement is represented by the following equation as was the

case with two- and f our-symmetrically arranged solid spheres:

Sh(l - 0) )/[Sh(l (5)

the effect of the second solid sphere may be obtained by use of Eq.{3) for f^^.

Figure 5 shows the results. The effect of the second sphere is well

correlated by the following equation:

—0

1«.77-(D^/D^)-°-21.(L/D^)-^-38-R-pI

•

,0.43

(1-W2/D^)1-^(W^/D^)(WV

0.43

(6)

where a parameter ^^i^^^^ ^ ^ i'^ the denominator represents the

effect of the first sphere on the interaction function f2 for the case of

arbitrary arrangement of two solid sphere.

Effect of Solid Sphere Layers

Figure 6 shows the effect of upstream solid sphere layers of equal

diameters in rectangular arrangement, where the drop is arranged on the

perpendicular bisector of the line connecting the centers of two spheres.

The solid lines in the figure represent the effect of number of solid sphere

layers on the rates of evaporation of a drop estimated by the following

equations

n=l; Sh(l - a)g)/[Sh(l - tOg)]Q = (2)

76
V 2D

9" 1.0 i Q8 ^ Q6 ^ 0.4

SI Q2

Plan View Side View

Solid Sphere

^rop

n.1

3

System Key Number of Layer

Water- Air

O n=1 • n=2

n=3

Octane-Air n-3

Present Model n-1

i

n=2

n=3

J L

RepA [-3

C^/D, [-]

W/Da [-]

167

0.76

0.60

2 A 6 8 10 20 30

l/Da r-]

Fig. 6 Effect of number of layers; a

drop is arranged on the

perpendicular bisector of the

line connecting the centers of

two spheres.

n=2:

n=3;

Number

of Layer

n = 1

n=2

n = 3

Flow

Fig. 7 Physical picture of the model

for calculation of void

function.

Sh(l - a)g)/[Sh(l - a)g)]Q = f^.f^

Sh(l - (Og)/[Sh(l - a)g)]Q = ^I'^l-^l

(7)

(8)

f3 = 1/(1+0. 77(D_/D.)-°-21(L/DJ"^-^^*^^Pa' (l-W^Dj^-^f \ (9) r A A J A n,K

where f. represents the effect of k-th layers on interaction function of n0.43

th layer 'ly

^n.k = l/'klif(VV''"W"^^» (10)

Good agreement is observed between the data and the theory.

PREDICTION OF VOID FUNCTION

Method of Calculation

Discussions in the previous sections may lead to estimation of void

function for heat and mass transfer of a group of drops or of particles. The

77
following simplifying assumptions were made for the present calculation.

1) Particle or drop size is uniform.

2) Particles or drops are arranged in rectangular arrangement on each layer

and the distance between each layer is equal to the distance between

particles or drops on the layer.

3) Relative position of an arbitrary particle or drop on each layer with

respect to the one on the other layers is independent.

4) Effect of downstream layers is negligibly small.

5) Effect of interaction of n-th upstream layers on mass transfer can be

estimated by the following equations:

n=l:

n>l:

Sh(l - a)g)/[Sh(l - ajg)]Q = fi'^-fs fn

—0 26

= 1/(1+0.77 (Dp/D^)'°-^^(L/D^)"^-^^'^^PA' (l-W./Dj^'^}

= l/{l+0.77(Dp/D^)"°*^^(L/D^)"^'^^*^^PA*

1' A'

1.5

n,k}

(11)

(3)

(12)

Figure 7 shows physical picture of the model. Calculations were made by

dividing one-quarter of a layer-element into 10x10 grid mesh. Void

function is calculated by the following equation.

L/2 L/2, , , -f^ f:-f^ f dXdY dXdY

0 0 12 n

F(e)

(13)

(L/4)

n

Void Function

Figure 8 shows results of the calculation. The ordinate is the void

Q55 0.6 0.7 0.8

e [-]

Fig. 8 Prediction of void function.

u

VI 3

I

CO

3

I

If)

1.0

0.8

0.6

O.A

02

Key Remarks

Present Model • Hsiung and Thodos

Wilkins and Thodos

;20'/o

OA 0.5 0.6 0.7

e [-]

0.8

Fig. 9 A comparison between experimental data and proposed

correlation.

78
function calculated by the present model and the abscissa is void

fraction of particle layers, where n is number of upstream layers of

rectangular arrangement. The figure clearly indicates that the void

function approaches to an asymptotic value as the number of layers increase.

The calculated void functions for n>5 are well correlated by the following

equations

:

F(e) = l/{ 1 + 1.29-(1 - e)°-'^^} (14)

Fig. 9 shows a comparison of the observed void function which were

obtained for mass transfer in gas-solid fluidized bed (4), (14) with the

prediction by the proposed correlation, where solid line in the figure

represents the correlation! Eq.(14)). Although some scattering of the data is

observed, fairly good agreement between the data and the theory is observed.

CONCLUSIONS

Measurements for the effect of adjacent solid spheres on the rates of

evaporation of a water- and an octane drop into dry air were made to show the

following conclusions.

1) The effect of adjacent solid spheres on the rates of mass transfer is well

predicted by Eqs. (3) and (12).

2) A new method for prediction of void function is proposed to show good

agreement with observed data.

ACKNOWLEDGMENTS

The authors are grateful to Mr. J. Narita of Tokyo Institute of

Technology for his assistance in measurements.

NOMENCLATURE

binary diffusion coefficient.m /s

diameter of front sphere,

m

diameter of liquid drop.ra

void function defined by Eq.(13)

interaction function defined by Eqs. (3) , (6) , (9)

interaction function defined by Eq.(lO)

distance between center of two spheres,

m

mass flux of component A,kg/(m • s)

number of layer

Reynolds number based on front sphere or drop(Dy^Uoo/ v ) Schmidt number (= v/Sf)

Sherwood number (=N^- Dp/g>g* Pg( - )

)

temperature ,K

velocity ,m/s

distance between center axis and center of solid sphere,

m

void fraction

kinematic viscosity of gas,m /s

density ,kg/m

mass fraction

0

Da

Dp

F(e)

^l'^2'^3"

^n,k

L

Na

nRepA

Sc

Sh

T

U

W,Wi,W-, =

eV

P

0)

79
Subscripts

00

A

B

G

0

s

= front sphere or drop

= rear sphere or drop

= gas phase

= surface of drop

= a single sphere without mass injection

= free stream

REFERENCE

1. Asano, K. , Taniguchi, I. and Kawahara, T., " Numerical and Experimental

Approaches to Simultaneous Evaporation of Two Adjacent Volatile Drops",

Proceedings of the 4th International Conference on Liquid Atomization and

Spray Systems , pp. 411-418, 1988.

2. Chen, S. J. and Tong, A. Y. , " Numerical Study on Hydrodynamics and Heat

Transfer of Droplet Arrays Using Grid Generation", AIChE S3Tnposium, Series

No. 257, Vol. 83, pp. 230-236,1987.

3. Chuchottaworn, P., Fujinami, A. and Asano, K.," Experimental Study of

Evaporation of A Volatile Pendant Drop Under High Mass Flux Conditions",

J. Chem. Eng. Japan, Vol. 17, pp. 7-13, 1984.

4. Hsiung, T. H. andThodos, G.,"Mass Transfer in Gas-Fluidized Beds:

Measurement of Actual Driving Forces", Chem. Eng. Sci., Vol. 32, pp. 581-

592, 1977.

5. Labowsky, M., " The Effect of Nearest Neighbor Interactions on the

Evaporation Rate of Cloud Particles ", Chem. Eng. Sci., Vol. 31, pp. 803-

813, 1976.

6. Marberry, M. , Ray, A. K. and Leung, K.," Effect of Multiple Particle

Interactions on Burning Droplets", Combust, and Flame, Vol. 57, pp. 237-

245, 1984.

7. Miura, K. , Miura, T. and Ohtani, S.," Effect of Surrounding Glass Beads on

Heat and Mass Transfer from A Water Droplet", Kagaku Kogaku Ronbunshu,

Vol. 1, pp. 241-246, 1975.

8. Ranz, W. E. and Marshall, W. R.," Evaporation from Drops; Part 1 and Part

2", Chem. Eng. Progr., Vol. 48, pp. 141-146, 173-180, 1952.

9. Ray, A. K. and Davis, E. J.," Heat and Mass Transfer with Multiple

Particle Interactions Part I. Droplet Evaporation" ,Chem. Eng. Commun.,

Vol. 6, pp. 61-79, 1980.

10. Shuen, J. S., " Effects of Droplet Interactions on Droplet Transport at

Intermediate Reynolds Numbers", AIAA 25th Aerospace Sciences Meeting , AIAA-87-0137, 1987.

11. Taniguchi, I. and Asano, K.," Numerical Analysis of Drag Coefficients

and Mass Transfer of Two Adjacent Spheres", J. Chem. Eng. Japan, Vol. 20,

pp. 287-294, 1987.

12. Taniguchi, I. and Asano, K., " Calculation of Drag Coefficients and Mass

Transfer of Two Adjacent Spheres of Various Diameter Ratios," J. Chem.

Eng. Japan, Vol. 22, pp. 701-703, 1989.

13. Taniguchi, I., Kawahara, T. and Asano, K.," Simultaneous Evaporation of

Two Adjacent Volatile Drops", J. Chem. Eng. Japan, Vol. 23, pp. 315-319,

1990.

14. Wilkins, G. S. and Thodos, G.," Mass Transfer Driving Forces in Packed and

Fluidized Beds", AIChE. J. Vol. 15, PP. 47-50, 1969.

80
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 3

INFLUENCE OF TURBULENCE ON THE MASS TRANSFER

FROM LIQUID FUEL DROPLETS

I. Gokalp, C. Chauveau, O. Simon and X. Chesneau

Centre National de la Recherche Scientifique

Laboratoire de Combustion et Systfemes R6actifs

Orl6ans, France

ABSTRACT

An experimental set-up and an experimental strategy have been developed to investigate the

influence of turbulence on the global mass treinsfer rates from fuel droplets. Heptane and decane

droplets of 1.5 mm initial diameter suspended in grid induced turbulent flows have been

investigated in the regime where the integral length scales of turbulence are, on average, 5 times

larger than the initial droplet diameter. The turbulence intensity has been increased up to 44%. A

new mass transfer parameter which distinguishes between the influences of the mean relative

velocity and that of turbulence structure has been introduced. Mass transfer from heptane droplets

has been found insensitive to turbulence. On the contrary, the same turbulence conditions exert a

significant influence on the mass transfer from decane droplets. It has been shown that the influence

of turbulence energy on the mass transport from decane droplets can be represented by a turbulence

Reynolds number. The Frossling coefficient has been found to increase with the turbulence intensity.

The differences in the sensitivity of heptane and decane droplets against the turbulence influence on

mass transfer are tentatively explained by introducing a "vaporization Damkohler number". The

present experimental results suggest that turbulence enhances the mass transport from liquid

droplets only for low values of this number.

INTRODUCTION

Spray combustion occurs in most cases in a turbulent medium. This enhances the

challenging nature of multiphase reacting flows, but also increases their complexity [1]. The studies

of the influence of a turbulent convective flow on the pure vaporization rates of single droplets are relevant for spray combustion, for one of the possible scenarios of the single droplet behaviour is

their total vaporization before reaching the flame zone. Furthermore, pure droplet vaporization is

also of interest for other applications such as spray drying or spray cooling [2].

One way to tackle the problem of the influence of turbulence on interphase transport

phenomena is to simplify it by investigating the interactions between a given turbulence structure

and a single vaporizing or burning droplet. This strategy could be justified if one is interested in

dilute dispersed flows, reacting or non reacting. But, even for a single droplet, current numerical

calculations try to avoid the problem of turbulence by assuming that a locally laminar region exists

81
near the droplet surface due to the much smaller size of the droplets compared to the typical large

eddy sizes [3].

In order to contribute to this area, we have initiated the investigation of the influence of

turbulent flows on the vaporization and burning characteristics of single fuel droplets [4, 5]. The

issues we are addressing are the modification of the heat and mass transfer characteristics of pure

and bicomponent fuel droplets and of the transition criteria between the envelope and wake flame

regimes, by varying the turbulence structure in terms of its intensity and length and time scales. In

the work described below, some preliminary results on the mass transfer (low temperature

vaporization) from n-heptane and n-decane droplets suspended in a grid induced turbulent flow are

reported.

The experimental set-up, the diagnostics used and the characteristics of the grid-generated

turbulence are briefly presented in section 2. Section 3 presents the experimental results on the

influence of turbulence on the global mass transfer parameters from suspended heptane and decane

droplets and proposes an interpretative framework for the observed results.

EXPERIMENTAL SET-UP AND FLOW CONDITIONS.

The experiments are performed in a facility specially designed for this purpose: The Forced

Flow Droplet Burning Facility (FF-DBF) which is fully described in Gokalp et al. [4, 5] and

Chauveau [6]. It is briefly presented here for completeness. The FF-DBF allows the investigation of

single and interacting droplet vaporization and/or burning in stagnant, laminar and turbulent

conditions. It consists of a closed loop flow channel in stainless steel (length: 3.20 m; height: 0.85 m)

which is fixed on a supporting structure. The flow channel is composed of an electrical axial fan, an

electrical heater, a set of laminarization ducts. The channel is connected to a stainless steel test

section of square cross section (side: 0.15 m; length: 0.50 m). A by-pass channel and a return channel

of circular cross-section (diameter: 20 cm) closed the loop to the electrical fan. Turbulence grids can be positioned at the inlet of the test section. Two circular glass windows are mounted vis a vis on the

parallel sides of the test section to allow optical access. The test section also supports the injection

and the ignition systems.

Droplets of approximately 1 mm to 1.5 mm initial diameter can be formed by using a syringe

located at the top of the test section. A quartz fibre (diameter 0.2 mm) is tilted to the syringe

extremity allowing the transfer of the droplet to the fibre. The lower extremity of the fibre has an

enlarged diameter of 0.5 mm. The fibre is then tilted back and positioned in front of the test section

windows. The sequence of operations is performed automatically by using step-motors controlled by

a micro-computer.

The air flow can be generated by the axial fan up to a maximum velocity of approximately 5

m/s in the test section. The characteristics of the turbulent flows are measured with laser Doppler

anemometry, and the flow monitoring system is calibrated against the mean longitudinal velocity at

the location of the droplet.

The principle diagnostic system used to investigate the droplet vaporization phenomena is

based on the visualization of the droplet. The imaging system is a high speed video camera, Kodak

Ektapro 1000, which can record up to 1000 full frames per second allows the detailed analysis of the

time variation of droplet dimensions. The Ektapro 1000 Motion Analyser's live, real time viewing

makes it possible to follow the investigated phenomena frame by frame. The digitized images from

the Ektapro system are transferred to a micro-computer where the image analyses are performed.

More details on the image analysis technique can be found in Chauveau [6].

In the calculation of the droplet surface regression rates, (or the global vaporization rate), we

evaluated a characteristic dimension for the droplet in terms of the diameter of a sphere of

equivalent projected surface area. The projected droplet surface area is determined by image

82
analysis. The droplet equivalent projected diameter is calculated as the diameter of a circle having

the same area as its projected image.

To investigate the mass transfer characteristics of fuel droplets in a turbulent medium,

turbulence parameters, such as the turbulence intensity and the length and time scales, have been

controlled by using square mesh grids. These parameters have been varied independently by varying

the mesh size, the mean flow velocity and the distance between the grid and the droplet. The mesh

sizes we used are, in mm, M = 1.4; 2.9; 5; 10 and 37. The grids with M = 37 mm and 1.4 mm or 5 mm

have also been used in combination. In the combined grid mode, another parameter we varied is the

distance between the two grids. For mean flow velocities equal to or higher than 1 m/s, the

turbulence structure in all cases is very close to that of isotropic turbulence [4, 6]. The relevant

turbulent length and time scales have been determined from spectral analysis of instantaneous

velocity signals and by using the theoretical relations for isotropic turbulence. The following flow

and turbulence characteristics correspond to those determined at the position of the droplet in the

test section.

The mean velocity has been varied from 1 m/s to 4.5 m/s with a step of 0.5 m/s. This means

that the Reynolds number based on the mesh size has been varied from 90 to approximately 6000. In

terms of the droplet Reynolds number, the variation is approximately between 100 and 450. The

turbulence intensity, defined as the ratio between the longitudinal r.m.s. velocity fluctuation and the

mean axial velocity, has been varied from Tu = 1.5% to 44%. The lowest turbulence intensity

corresponds to our quasi-laminar conditions as it was impossible to obtain a pure laminar flow in the

test section. The longitudinal integral length scale varies from 1 mm to 19.5 mm. This means that

the ratio between the initial droplet diameter and the integral length scale varies between 1.5 and

0.08. The maximum value obtained for the turbulence Reynolds number based on the integral length

scale is 350. For each mean velocity, 3 to 9 different turbulence conditions are obtained. In total, 42

different flow conditions have been explored for the vaporization of heptane droplets; for decane, this

number is 17.

However, it is important to mention that only 3 of these 59 experimental conditions

correspond to cases where the initial droplet diameter is lower than the integral length scale of the

turbulence. This means that in most of the cases, the vaporizing droplet interacts with small scale

and less energetic eddies.

RESULTS ON MASS TRANSFER RATES

The main objective of this study is to investigate the influence of a turbulent flow on the

global vaporization rates of suspended fuel droplets. More specifically, it is desirable to separate the

respective influences of the mean velocity from that of the turbulence. This is because in real spray

situations, small droplets may easily be transported "in the mean" by the carrier phase, thus with a

zero mean relative velocity. However, they may be subjected to the instantaneous fluctuations of the

flow field.

In this section, experimental results on the global vaporization rates of heptane and decane

droplets in grid induced turbulent flows are presented. The values of the vaporization rates that are presented below are the average values of 3 to 5 runs under the same flow conditions and for initial

droplet diameters equal approximately to 1.5 mm.

Vaporization constants in laminar and turbulent flows.

Fig. 1 presents the time variation of the droplet diameter for heptane droplets under quasilaminar conditions. The stagnant and two forced convective runs are shown. For these last cases, the

droplet Reynolds numbers are approximately 100 and 425. The droplet diameter is normalized by its

initial value; the time axis is normalized by the total vaporization time of the stagnant case. From the time variation of the instantaneous droplet diameter ds, we have determined the

variation of the instantaneous vaporization mass flux. The results corresponding to the previous

83
cases are presented on Fig. 2. The curves are normalized by the initial droplet mass and the total

vaporization time of the stagnant case. Except the initial transient phase, which is too short to be

observed in these low temperature vaporization experiments, these curves are very close to those

predicted by [7]. Similar results have also been obtained for the other mean convective velocities and

for decane droplets.

The experiments with heptane droplets have shown no influence of turbulence on the

vaporisation rate determined from the dMaw as K = - d(ds2)dt [5]. The same conclusion can also be

drawn for the influence of the droplet Reynolds number for increasing turbulence intensity. The

Frossling coefficients [8, 9] have been found slightly scattered around the value of 0.19. It is

noteworthy that the Frossling coefficient determined in this study is significantly lower than the

value of 0.276 obtained by Frossling himself and widely accepted today.

As shown on Figs. 3, 4 and 5, the results for decane droplets vaporizing under the same

conditions are significantly different. Indeed, for the same turbulence structures as for the heptane

case and for three values of the mean convective velocity (1, 2 and 3 m/s), an important increase of

K/Ko with the turbulence intensity is observed (Fig. 3). On Fig. 4, the influence of the turbulence

intensity on the time variation of the droplet diameter is shown for U = 2 m/s. On the other hand, as shown on Fig. 5, the Frossling coefficient for decane, which is equal to 0.19 for the quasi-laminar

case, increases gradually with turbulence intensity, and reaches the value of 0.274 for Tu = 20%.

The curves of Fig. 3 can be represented by the following linear relationship:

(K/Ko)t = Cl(uAJ) + (K/Ko)l (1)

where subscripts T and L refer respectively to turbulent and laminar cases. In other words, (K/Ko)l

represents the increase of the mass transfer constant due only to laminar forced convection. Eq. 1

can be written also as

Go = (Kt - Kl)/Ko = Cl(u/U) (2)

It is proposed here to consider the left hand side of this equation as representing the influence of

turbulence alone on the mass transfer rate, with respect to its stagnant case value. This parameter

is introduced in order to separate the effects of mean convective velocity from that of the turbulence.

In the following, we shall use this parameter in the discussion of the turbulence effect on the global

vaporization rates of droplets.

Influence of the turbulence Reynolds number

Fig. 6 shows the variation of Go with the turbulence intensity for three mean velocities. It is

clear that the influence of turbulence is enhanced with increasing mean velocity. The coefficient Cl

passes from 0.0053 for U = 1 m/s to 0.018 for U = 3 m/s.

Previous studies have shown that the transfer rates from particles or drops increase with

increasing turbulence intensity. They also show that this increase is stronger when the droplet

Reynolds number is high. This effect has been generally correlated with the parameter Re(u'/U).

If the Reynolds number is taken to be the one based on the droplet diameter, this parameter

can be written as (u/U)(Ud/v) or ud/v, which is a turbulent Re)molds number based on a turbulence

length scale equal to the droplet diameter, where m is the kinematic viscosity of the free stream. Fig.

7 shows the following correlation for three mean convective velocities:

Go = (Kt - Kl)/Ko = C2(ud/v) = C2ReTd (3)

The values for 02 are 0.0052, 0.0056 and 0.0060 for respectively U = 1, 2 and 3 m/s. Even if a slight

dependence against the mean velocity is still observed, this last correlation is significantly improved

compared to that given by Eq. 2. As the parameter Go represents only the influence of turbulence on

84
the mass transfer rate, the improvement of the correlation when the turbulence parameters are introduced instead of the mean velocity, seems encouraging.

Influence of the txirbulence length scale

Past experimental work on heat and mass transfer from spherical particles or drops, have

shown some influence of the ratio between the turbulence length scales and the particle diameter

[10, 11, 12]. A detailed inspection of our experimental results shows also some length scale influence

on the mass transfer from decane droplets.

For example, for U = 1 m/s, the parameter Go = (K-r - Kl)/Ko is larger for u/U = 8% and

Le/do = 2.7 than for uAJ = 10% and Le/do = 9. Similarly, for U = 3 m/s, this parameter is larger for

u/U = 11% and Le/do 4.7 than for u/U = 11.5% and Le/do = 11.8. These comparisons seem to indicate

that, for a given turbulence intensity, smaller scale turbulence structures are more efficient to increase the mass transfer rates. It should however be noted that this length scale influence does not

explain the differences that we observed between the sensitivities of mass transfer rates from

heptane and decane droplets against turbulence, as the Le/do ratios for both fuels are the same.

Influence of time scales and discussion

The main result of this investigation is indeed the differential sensitivities of heptane and

decane droplets for the same turbulence structure. We have just pointed out that this difference is

observed in spite of the similar turbulence length scale to droplet diameter ratios. In this part, we

shall propose a tentative explanation in terms of the time scale ratios of the phenomena involved.

Fig. 9 shows for decane droplets the variation of the parameter Go with the integral time

scale tg(j of velocity fluctuations. This time scale corresponds to turbulent eddies having integral

length scales equal or smaller than the instantaneous droplet diameter. This figure shows that, for a

fuel droplet of given initial diameter, the efficiency of turbulence to increase the mass transfer rates

decreases as the time scale of the eddies effectively interacting with the droplet increases.

In order to compare the heptane and decane cases on a time scale basis, we attempted to

define a time scale ratio based on tgd and on a time scale characterizing the thermodynamics of

vaporization, ty. This ratio could be considered as the equivalent of a "vaporization Damkohler

number", Da^. As ty is constant for a given fuel, the variation of Fig. 9 can also be interpreted as the

effect of Day. This variation indicates that the effect of turbulence on mass transfer from droplets is

more efficient for low values of Day. For the comparison of the heptane and decane time scale ratios,

we use a baseline case corresponding to the same turbulence structure for both fuels having an

initial droplet diameter of 0.15 cm. To estimate the time scale characterizing the thermodynamics of

vaporization, we propose to define it, as a first approximation, as the ratio between the film

thickness and the radial blowing velocity (or the Stefan velocity) V^. The radial velocity is calculated

by equating the expression of the instantaneous mass transfer rate as m = 4ji r^pgV^ to its

expression from the quasi-stationary theory and by assuming the Lewis number equal to unity. In

the estimation of the film thickness we used the calculation procedure recommended by Abramzon

and Sirignano [7]. The results of these calculations are summarized on the following table .

C7H16 C10H22

Vr (cm/s) 0.1 0.004

(8)m (microns) 266 191

tv (ms) 266 4775

Dav 0.03 0.0016

As expected, the radial velocity for heptane is much higher than that of decane. At the

present stage of this investigation, it is speculative to attribute any intrinsic significance to the

85
precise values of the vaporization Damkohler numbers. However, what is certainly significant is the

ratio between Dav(C7H16) and Dav(ClOH22) and which is equal to 18.8. This ratio is in fact very

close to that between the respective mass transfer numbers or the vaporization constants under

stagnant conditions Ko of the two fuels.

If a reference value of 1 ms is attributed to the time scale characterizing the thermodynamics

of vaporization for decane, Fig. 9 shows that for a vaporization Damkohler number higher than 15,

the turbulence effect on the mass transfer is negligible. As the characteristic vaporization time for

heptane is approximately 20 times shorter than that of decane, for the same turbulence conditions as on Fig. 9, the data point corresponding to heptane will be located at the high vaporization

Damkohler number end of the curve, where no turbulence effect on the mass transfer is expected.

This reasoning based on the comparisons of the characteristic times of turbulence and of

vaporization thermodynamics may then explain the sensitivity differences of heptane and decane

droplets against the same turbulence structure.

CONCLUSIONS

It has been shown above that the influence of turbulence on the mass transport from decane

droplets can be represented by the turbulence Reynolds number. The Frossling coefficient which

correlates the effect of mean relative velocity (or of the droplet Reynolds number) on the mass

transfer has been found to increase with the turbulence intensity.

The differences in the sensitivity of heptane and decane droplets against the turbulence

influence on mass transfer rates are tentatively explained by introducing a "vaporization Damkohler

number". The present experimental results suggest that turbulence enhances the mass transport

from liquid droplets only for low values of this number.

REFERENCES

1. Faeth G.M. (1989) in Turbulent Reacting Flows (Ed. R. Borghi and S.N.B. Murthy) New York:

Springer Verlag, pp. 764-814.

2. Lefebvre A.H. (1989) Atomization and sprays. New York: Hemisphere Publishing Corporation

3. Megaridis CM. and Sirignano W.A.(1990) Paper presented at the ASME Winter Annual

Meeting, Dallas, Texas

4. Gokalp I, Chauveau C, Richard J.R., Kramer M. and Leuckel W. (1988) Twenty-second

Symposium (International) on Combustion, The Combustion Institute, pp. 2027-2035.

5. Gokalp I., Chauveau C. and Monsallier G. (1989) in Drops and Bubbles. Third International

Colloquium on Drops and Bubbles, (Ed. T.G. Wang), AIP Conferenc Proceedings 197; pp. 362-

372.

6. Chauveau C. (1990) PhD Thesis, University of Orleans.

7. Abramzon B. and Sirignano W.A. (1988) AIAA Paper 88-0636. 14p.

8. Frossling N. (1938) Gerlands Beitr. Geophys., 52:170.

9. Ranz W.E. and Marshall W.R. Jr. (1952) Chemical Engineering Progress, 48:141 et 48:173.

10. Van Der Hegge Zijnen B.C. (1958) Appl. ScientJles., 7A:205.

11. Raithby G.D. and Eckert E.R.G. (1968) Int. J. Heat Mass Transfer, 11:1213.

12. Sandoval-Robles J.G., Delmas H. and Couderc J.P. (1981) AIChE Journal, 27:819-823.

86
5.5

5

a

1.2

1

0.8

0.6

0.4

0.2

HEPTANE Tu = 1.5 %

U = 0 m/s

0= 1 m/s

0= 4.4 m/s

0 0.25 0.5 0.75 1.25

t/ts

Fig. 1. Timewise variation of the droplet diameter. Heptane droplets under stagnant

and two laminar convectives cases.

7.5

6

on H 4.5

o

0 e 3

1.5

HEPTANE Tu = 1.5 %

\

\ - \ 0= 0 m/s

\ 0= 1 m/s

\ 0= 4.4 m/s - \

—

1 1 1 1 1 0.2 0.4 0.6

t/ts

O.B 1.2

Fig. 2. Timewise variation of the instantaneous vaporization rate. Same conditions as on Fig. 1.

a

3.5

2.5

DECANE

0 : 0 = 1 m/s

a : 0 = 2 m/s

. : 0 = 3 m/s

0 -<7 0____--« _ 0

: 1 1 1

OECANE U= 2 m/s

Tu' 5.3 t lu- 18.9 I

10 20 30

Tu ( % )

40 50

1.25

Fig. 4. Influence of turbulence intensity on the

timewise variation of the droplet

diameter. The case of decane.

Fig. 3 Variation of the vaporization rate with

turbulence intensity. The ca.se of decane.

87
y

y' - UcLAlNC

yy yy y

—

'X

Tu = 5% : K/KO = 0 245 Reg'^ Sc'" + 1

Tu = 10?.: K/KO = 0.263 Reg"'.Sc"3 + 1

X

1

Tu = 207*-

1

K/T<0 = 0.274 Regi'i.Sci^ + 1

1

8 12 16 20

Fig. 5. Frossling correlation for different

turbulence intensities. The case of

decane.

Fig. 6. Influence of the turbulence intensity on

the mass transfer rate. The case of

decane.'O^ 1 m/s^ 2 m/s,A3 m/s

Tu' 5.3 X Tu= 19.9 X Tu= a X

ReTd

Fig. 7. Influence of the turbulence Reynolds

number on the mass transfer rate. The

case of decane. Same symbols as Fig. 6.

1.2

Fig. 8. Influence of the turbulence intensity on

the instantaneous vaporization rate. The

case of decane. U = 2 m/s

Fig. 9. Influence of the turbulence time scale on the mass transfer rate. The case of decane.

88
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 4

SUPERCRITICAL DROPLET EVAPORATION

A. Umemura, X. Y. Chang and T. Fujiwara

Faculty of Engineering

Yamagata University

Yonezawa, Japan

ABSTRACT

Numerical calculation is performed to examine the evaporation process of a liquid fuel droplet which is immersed in an otherwise quiescent, inert gas at a various state beyond the critical point of the fuel. Either is realized at any

instant, of (i) the subcritical evaporation regime in which phase change takes

place discontinuously at the droplet surface or (ii) the supercritical

evaporation regime in which phase change occurs spatially continuously.

Satisfactorily to simulate the transition from (i) to (ii), it is essentially

important to take into account the fact that the coefficient of diffusion

between the fuel and the inert gas should vanish at the droplet surface when the

thermodynamic state of the droplet surface attains to the critical point of the

mixture system.

Conditions under which the droplet experiences the transition are

identified in terms of the initial system parameters. Consistent with past

expemental results no singular behavior is observed in the droplet temperature

change at the instant of transition. The core characterized by high fuel

concentration behaves like a liquid droplet even after the transition, and can

be distinguished from the other part by a sharp but continuous change at its

boundary. Once the core is about to disappear, the concentration as well as the

temperature changes abruptly to assume the gaseous state. The concept of "vapor

puff" which has been frequently used to model a supercritically evaporating

droplet seems only valid to describe the state after disappearance of the liquid

core.

INTRODUCTION

As well known a pure liquid under its supercritical pressure may undergo a spatially continuous phase change from a liquid to gaseous state. A similar

situation is expected for a hydrocarbon droplet evaporating in an air whose

temperature and pressure exceed the critical point of the fuel, as encountered

in Diesel engines and the like. The phenomenon is, however, not so simple

because the occurrence of the transition from subcritical evaporation regime,

in which phase change takes place discontinuously at the droplet surface, to

supercritical evaporation regime, in which phase changes in a spatially

continuous way, depends not only the system pressure but also on the ambient gas

89
temperature and is closely related with the droplet heat-up process itself.

A numerous attempt has been made to reveal the underlying physics of

supercritical droplet evaporation while there still remain ambiguities about the

transition f eature [

1 -1 0] . Among them the following are important to note. (1)

In most of the previous investigations the occurrence of the transition is

judged by whether the droplet surface temperature reaches the fuel critical

temperature or not. However, for a mixture system in which the initial droplet

and ambient gas comprise different chemical species, the mixture critical point

system should be considered instead of the fuel critical point. For a hydrocarbon/nitrogen system of interest the mixture critical temperature is

usually lower than the fuel critical temperature and the mixture critical

pressure exceeds the fuel critical pressure. This raises a question upon the

conventional criterion. (2) There seems to be anticipation of singular behaviors

by association of the critical behaviors of specific heat and thermal

conductivity observable for a pure substance. This is not the case for a mixture system because of different cri ticalities of he components. One of the

authors demonstrated in an asymptotic analysis[13] that, when the droplet

experiences the transition at a finite radius, the surface temperature

continues to increase regularly with a gap to the droplet center temperature

until it reaches the mixture critical temperature. An abrupt change in the

droplet center temperature, which is detected experimentally, may occur only

when the droplets gasifies completely. Considering a not large difference

between the fuel and mixture critical temperatures, there is a possibility that

this behavior is mistakenly identified as a singular behavior at the

transition. (3) Without the knowledge of critical material properties it is

impossible to simulate the transition satisfactorily. It was also shown [11-14]

that vanishing of the relevant binary diffusion coefficient at the critical

droplet surface is essential to realize the unsteady transition in the realm of

the current continuum theory which is based on the local equilibrium hypothesis

and the linear phenomenological relationships. Failures in past numerical

calculations are due to lack of this basic consideration.

In consideration of the above-mentioned status we perform numerical

calculation to reveal the characteristic feature of droplet evaporation at

elevated pressures. At the stage when knowledge about material properties for

mixtures is not complete it is, however, difficult to do in a quantitatively

satisfactorily way. Therefore, the calculation is based on an appropriate,

simplified physical model which allows for minimum items necessary to simulate

the phenomenon satisfactorily, and aims at a rather qualitative nature of the

problem. Nevertheless, the results will be found to be in good agreement with

the existing experimental data.

FORMULATION

Consider the spherically symmetric evaporation process of a liquid fuel

droplet which is immersed in an otherwise quiescent, inert gas at high

pressures above the fuel critical pressure. At initial time, t = 0, the droplet

of radius, ao , has a uniform temperature , T d. while the ambient gas,T,. The

droplet gasifies to produce the Stefan flow velocity, u, whose magnitude is so

small that the pressure, p, can be regarded as constant everywhere. The

phenomenon is thus governed by the following heat and mass transport equations.

Overall mass:

90
Energy: p[|h

^ ^|hj ^ _1^9 ^2;^lT

9t 8r' 8r ^dr (2)

Fuel: BY 8Y, 18 , BY

To close the equation system we need an equation of state, p= p(T,Y;P), and a caloric equation, h = h(T,Y;P), together with appropriate expressions for the

mixture thermal conductivity , A , and binary diffusion coefficient, D, as

functions of temperature, T, fuel mass fraction, Y, and pressure, P. All these

equations hold irrespective of liquid or gas phase. He shall use the subscript

* to distinguish liquid properties.

The initial and boundary conditions to be satisfied are

Initial: T = T , Y = 0 for r > ^0

T*= ' 1 for 0 < r < (4)

Boundary: T-^T^, Y-*-0 as r-*-co

(5)

Besides, in the subcritical evaporation regime the folloHing conservation

equations, (7)-(9), must be satisfied at the droplet surface, r = a(t). Since we

assume the hold of phase equilibrium at the droplet surface, the interfacial

fuel concentrations, Y,, and Y, , are determined from Eq. (6) as functions of

droplet surface temperature, T, , for a given system pressure, P.

Uj(T^,Y^;P) = M*i(Tg,Y^g;P) (6)

P*<u*- f|) = - fl^ = (7)

97 ~ *^r~ " ^ith L = h - (8)

P^l? - PAfF" = - ^^^^ A = Y^- Y (9)

All quantities appearing in the above equations are made dimensionless as

described below, so that the equations preserve the original forms with the

same symbols employed to denote the corresponding dimensionless quantities. The

temperature, T, and pressure, P, are reduced by the fuel critical temperature,

Tpc, and pressure, Ppc- The radial coordinate, r, and instantaneous droplet

radius, a, are divided by the initial droplet radius, ao (thereby ao = 1 in the

nondimensional form of the equations). The material properties, p , D, ht , and

A , are made dimensionls so that a Lewis number constructed from

nondimensionali zing reference quantities becomes unit. That is, we make

dimensionless the mixture density, p, by the fuel critical density, ppc, the

specific heat at constant pressure, h t. by the universal gas constant, R, and

91
the mixture thermal conductivity, A , by an appropriate reference value, A r

• Then, the diffusion coefficient, D, is nondimensionalized by the quqntity, p pcR/

Ar • Correspondingly, the time, t, is made dimensionless by ao^ /3 pcR/ A r

• Note that we do not assume unit Lewis number for the trasport field; the local

Lewis number changes significantly from point to point at any instant. From the

invariance of the governing equation form it follows immediately that the d^-

law holds between the initial droplet diameter and the evaporation time for the

case of subcritical evaporation regime dominating throughout the droplet

lifetime.

The material properties are estimated as follows. The Redlich-Kwong

equation of state with adequate mixing rules for binary mixtures[15] is adopted

to describe the local-equilibrium state of the fuel/air binary mixture. The

required thermodynamic variables are all derived as functions of pressure,

temperature and composition through the well-known procedures in thermodynamics.

Normal butane and nitrogen are chosen as the fuel and inert gas , respectively,

since the recommended values of the parameters appearing in the mixing rules

are readily available for this binary system. The mixture thermal conductivity

is estimated by applying the corresponding states principle to an empirical

relation for pure substance [ 1 6 ] . The binary diffusion coefficient is calculated

according to the proposed expression in Ref.[12] which satisfies the vanishing

condition at the mixture critical point. Identifying with the self -diffusion

coefficient of a hypothetical pure substance, the factor. Do, is estimated by

applying the corresponding states principle to self -diffusion coefficient for

pure substance [ 1 7] . In the present calcualtion the mixing heat, Hy , is assume

zero for simplicity.

RESULTS AND DISCUSSION

Numerical calculation was performed

for a particular value of initial droplet

temperature, Tj = 0.5, to examine the

variations of characteristic quantities

with ambient gas conditions, T, and P.

He start with the case, P < 1. In

this case the droplet has a liquid-gas

interface throughout its lifetime. This

can be seen from Fig. 1 which shows the

saturation curves calculated from Eq.(6).

Characteristic for P < 1 is that a pair of

saturated-liquid branch (solid line) and

saturated-vapor branch (dotted line) merge

at a point on the Y = 1 line, i.e., the

boiling point at the pressure under

consideration. Therefore, if we plot, on

the same figure, a curve Y(t,T) which is

obtained by eliminating r from the

instantaneous temperature and

concentration profiles, T(t,r) and Y(t,r),

the curve connecting the two states,

(To.Yo) at the droplet center and (T, ,0)

at infinity, must intersect the two

branches, indicating the existence of a liquid-gas interface. Since the latent

heat of vaporization, L, takes a non-zero

92
value, the state within the droplet becomes uniform as time proceeds and

eventually settles down at the wet-bulb condition that all heat transferred from

the gas phase is consumed for vaporization at the droplet surface.

Correspondingly the instantaneous evaporation constant, K, also becomes constant

and the d^-law holds exactly at any instant. The temporal changes in the

characteristic quantities in this subcritical pressure case are shown in Fig. 2.

For the supercritical pressure case, P > 1, there is a critical point in

Fig. 1 where the latent heat of vaporization vanishes. One can consider two

possibilities, (i) The curve, Y(t,T), intersects the saturation curve. In this

case the droplet has its surface and requires a non-zero heat for vaporization

similarly to the subcritical pressure case, (ii) The curve, Y(t,T), has no

intersection with the saturation curve. In this case, liquid phase changes to

gas phase spatially continuously so that there is no distinguishable liquid-gas

Figure 2 Temporal changes in characteristic quantities

(P = 0.9, T. = 1.2)

see nomenclature footnoted below Fig. 4

t

Figure 3 Temporal changes in characteristic quantities

(P = 2.5, T. = 1.1)

see nomemclature footnoted below Fig. U

93
1.2 r1.0

2.29

t = 1.50

(a) Temporal changes in temperature and fuel concentration profiles

(P = 2. T. = 1.25)

Transition

(b) Temporal changes in characteristic quantities

(P = 2.5, T. = 1.2)

Figure 4 Transition feature

Characteristic quantities:

Droplet center temperature, T ,o , and fuel mass fraction, Y,o

Droplet surface temperature, T , , and fuel mass fractions, Y,, and Y,

Instantaneous evaporation constant, K = - a(da/dt)

94
1.3 -

1.2 -

ho

1.1 -

0

Figure 5 Ambient gas condition

for transition to occur

interface. If the system pressure is not

high enough, the subcritical evaporation

regime (i) dominates until the droplet

gasfies completely. Figure 3 shows the

variations of the characteristic

quantities in such a case. The behaviors

are similar to those in Fig. 2. In

particular it should be noted that the

wet-bulb state is established toward the

end of the droplet lifetime. This is

because the relaxation time of the

transport phenomenon within the droplet

diminishes in proportion to the vanishing

size of the droplet.

As the pressure is increased for the

fixed ambient gas temperature, the wetbulb state shifts to the mixture critical

point and the droplet comes to reach the

critical state at the instant of complete

gasification of the droplet. Beyond this

critical value of pressure the transition

from sub- to-supercritical evaporation

regime occurs at a finite radius of the droplet. Results for this case are

shown in Fig. It. Figure 4(a) shows the temperature (solid line) and

concentration (dotted line) profiles at the time written alongside the curves

for the case of P = 2 and T, = 1.25- The value of r at which the concentration

profile is discontinuous gives the instantaneous droplet radius. In the figure

the state of the droplet surface attains to the mixture critical point just

before t = 1.8 when the droplet radius reduces to three quarters the initial

one. Correspondingly the discontinuities in the concentration and in the

temperature gradient at the droplet surface disappear and degenerate into

inflection points of the profiles. The subsequent change in the concentration

profile indicates that a portion of high fuel concentration, nearly Y = 1, still

remains qround the droplet center. This portion will be called "liquid core"

because it has a similar property to the liquid fuel which the droplet has

before the transition. The core radius, which may be defined by the radial

distance to the inflection point of the concentration profile, decreases faster

than the degression speed of the droplet surface. Once the core radius is about

to vanish, the fuel concentration at the origin rapidly decreases to zero

indicating total gasfication. At the same time the temperature at the origin

increases abruptly toward the asymptotic value of the ambient gas temperature.

These behaviors can be more clearly seen in Fig 4(b) in which the variation of

the characteristic quantities with time is plotted for a different ambient gas

condition. It is confirmed from this figure that there appears no singular

behavior in the droplet surface temperature toward the transition instant. The

droplet surface temperature continue to increase regularly with a finite gap to

the droplet center temperature. In particular the droplet center temperature

increases monotonically and shows no singular behavior at the transition

instant, consistent with the past experimental result. The rapid increase in the

temperature. To, about t = 2.3 is due to the vanishing of the liquid core at

that time.

In Fig. 5 the range of initial gas temperature is shown against pressure,

for which the droplet experiences the transition during its lifetime. Above the

solid curve the transition occurs. In the figure the variation of the mixture

95
critical temperature with pressure is also plotted by the dotted line for

reference. For P ~ 1 the ambient gas temperature must be very high to cause

the transition while the required ambient gas temperature decreases to the

mixture critical temperature asymptotically at high pressures.

CONCLUDING REMARKS

Satisfactorily to describe the transition from sub-to-supercritical

evaporation regime it is essentially important to take into account the fact

that the binary diffusion coefficient should vanish at the liquid-gas interface

in a mixture critical state. In the present study, adopting a model diffusion

coefficient which satisfies such property we have succeeded in revealing the

characteristic feature of droplet evaporation at elevated pressures and in

predicting the ambient condition under which the droplet experiences the

transition within its lifetime. As anticipated, the "droplet" preserves the

liquid-like property even after it lost the liquid-gas interface. Since

continuous phase change occurs in a thin layer, rough optical observations will

gives images as if ther still exists the droplet with surface. The concept of

vapor puff which has been frequently used to model the supercritical

evaporation regime is only valid for the state which follows the vanishing of

the liquid core. Since both temperature and concentration profiles have

inflection points at the boundary of the liquid core, it would be possible to

define apparent quantities pertaining the liquid-gas interface such as latent

heat of vaporization and concentration jump etc. even in the supercritical

evaporation regime so far as there exists the liquid core.

REFERENCES

1. Spalding, D. B. , ARS J., 29, 825 (1959)

2. Brzustowski.T. A. , Can. J. Chem. Eng., 43, 30 (1965)

3. Hieber,P.R., AIAA J.._l, 2764 (1963)

4. Rosner,D.E., AIAA J.,_5, 163 (196?)

5. Manrique, J. A. and Borman,G,L., Int. J. Heat and Mass Transf . , 12, 1081 (1969)

6. Faeth,G.M. et al.. Twelfth Symp. (Int'l) on Combustion, p. 9 , The Combustion

Institute, Pittsburgh, PA (1969)

7. Matlosz,R.L. et al.. Int. J. Heat and Mass Transf., ^5, 831 (1972)

8. Hiroyasu,H. et al.. Trans. Japan Soc. Mech. Engrs., 44, 3885 (1978)

9. Kadota,T. and Hiroyasu,H., Trans. Japan Soc. Mech. Engrs., 44, 3885 (1978)

10. Kadota.T. and Hiroyasu,H., Trans. Japan Soc. Mech. Engrs., 46, 1591 (1980)

1 1 . Umemura,A. , Trans. Japan Soc. Mech. Engrs., 478 , 82-94 (1986); also Heat

Transfer Jap. Re., 17 . 82 (1988)

12.lJmemura,A. , Twenty-first Symp. (Int'l) on Combustion, p. 463, The Combustion

Institute, Pittsburgh, PA (1986)

1 3. Umemura, A.

,

Lecture Notes in Physics 299: Mathematical Modeling in Combustion

Science (J.D. Buckmaster and T. Takeno, Eds.), p. 52, Springer-Verlag,

Heidelberg, Germany (1988)

14. Chang, X. et al.. Progress in Astronautics and Aeronautics Vol. 113, p. 168,

AIAA, Washington, DC (1988)

15. Chueh,P.L. and Prausnits, J. M. , I&EC Fund., 6^^ 492 (1967)

l6.Stiel,L.E. and Thodos.G., A.I.Ch.E. Jl, 10, 26 (1964)

17.Lee,H.L. and Thodos,G., Ind. Eng. Chem. Fund., 22, 17 (1983)

96
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 5

WATER DROPLET EVAPORATION IN FIRE PLUMES

G. Krishan*, R.P. Gakkhar^ and S. Prakash^

Fire Research Division, Central Building Research Institute, Roorkee, India

^Dept. of Mechanical Engineering, University of Roorkee, Roorkee, India

ABSTRACT

In the present paper, the vaporization of water droplet in a

sprinkler system encx>untering the ambient conditions vAiich exist along the

centreline of a fire plume, has been studied. The authors had developed an

unsteady droplet vaporization model for constant ambient conditions. In

the present work the unsteady droplet vaporization model has been modified to

accommodate the conditions of varying ambient teroperature and the free stream

velocity, for a water droplet passing throu^ a fire plume. The evaporation

of various sizes of water droplets ranging from 1 mm to 4 ram diameter falling

through the different intensities of fire plumes has been studied, and the

optimum size of water droplets vdiich should come out of the water sprinkler

heads has been proposed.

INTRODUCTION

Althou^ the automatic sprinkler systems have been using water as the

suppressant for many decades, the exact mechanism of direct extinction by

water as it is delivered from an automatic sprinkler system is not fully

understood [1]. The water in the form of spray comes out from sprinkler

heads. The water droplets of varying sizes during their fall, encounter

conditions of varying ambient teroperature and coirposition, and get heated up.

These droplets keep on vaporizing as they pass throu^ the hot convective

environment, before falling on the burning materials. The droplets coming

from sprinkler heads should be of a size, such that, even after vaporization

in a hot environment, they retain optimum size before falling on the seat of

fire, to quench it and also to avoid water damage.

As a water droplet enters a fire plume and continues to fall, it heats up

and vaporize. Moreover, it encounters a hot convective flow field with

increasing ambient tenperature and plume velocity. The relative velocity

between the droplet and the hot ambient gases decreases due to drag. These

varying ambient conditions along with the unsteady heating and shrinking of

the droplet make the problem of droplet vaporization in fire plume

essentially unsteady.

97
There have been several studies [1] on vaporization of water droplets in

steady state and with cxDnstant ambient conditions and therefore, they do not

adequately represent the droplet vaporization problem under cxDnditions

relevant to the actual fires. Therefore, it becxjmes inportant to study the

vaporization rate of water droplets vaporizing in a hot convective environment

with varying ambient conditions relevant to actual fires.

MODEL USED

The present study makes use of a model develcped by Prakash and Krishan

[2] . This model has been coiqjled with the model of Gakkhar and Prakash [3],

to accomraodate the varying ambient conditions encountered by the water

droplets as it falls through the plume. The plume is assumed to be stea<^.

However, the gas-phase flow conditions in the vicinity of the falling droplet

are assumed to be quasi-steady. The droplet is assumed to be falling in the

environment of the plume and the interaction between the neighbouring

droplets has been ignored. Thus, the model considers an isolated droplet

vaporizing in a convective flow field v*iich falls vertically downwards along

the centreline of fire plume. As the droplet enters a hot convective flow

field, a gas-phase viscous boundary layer develops around the droplet.

Energy is transferred from the ambience to the droplet, a part of vAiich goes

into heating up the droplet v^le the remaining part is utilized in vaporizing

the droplet. The heating up of the droplet is essentially unsteady. This

model was modified to accommodate the varying ambient conditions of

teirperature and the free stream velocity for a water droplet passing throu^

a fire plume. The quasi-steady, gas-phase momentum, energy and species

boundary layer equations were solved by an integral approach developed by

Prakash and Sirignano [4], v^ile the one-dimensional unsteac^ heat diffusion

equation was solved for the liquid-phase heating. Coupling at the gas-liquid

interface is in the form of conservation of heat flux and the water vapour

mass flux at the droplet interface. The coi:5)led problem was solved for

three plumes of different intensities, namely, weak, medium and strong

plumes [5] , for droplet diameters ranging from 1 to 4 ram. The hei^t of

sprinkler from the fire source is taken as 4.57 m.

Various correlations giving the teitperature and velocity along the

centreline of plumes were taken from reference [6] . The water droplet

penetration into the plume was corrputed by using the momentum balance in the

vertically downward direction.

dU

m = mg - Cr^Vf. S, . . . (1)

dt ^2

assuming the initial droplet velocity coming out of the sprinkler to be zero.

The drag coefficient, C^, variation with Reynolds number for vaporizing

droplet was coirputed, as suggested by Yuen and Chen [7], using the standard

drag curve and modifying the Reynolds number. For calculating the Reynolds

number, the free stream density was used v*iile the dynamic viscosity, jl4, was

evaluated at the reference temperature Tj., and the reference ambient

composition ^

given below:

Tj. = Tg + (T«, - Ts)/3 ...(2)

98
^w,r ^w,s • • • (3)

The mixture viscosity was then calcailated from the values of viscosity

for air and the water vapours at the reference tenperature [8]. The Cj^ vs

Reynolds number correlations for solid sphere was taken from reference [9].

PLUME CORRELATIONS USED

Hie ejq)erimental data of three plumes [5] , covering a wide range of plume

intensities (named, weak, medium and strong plumes for convenience) , were used

for the present stucty of water droplet vaporization in fire-plumes. Ihe

relevant data is given in Table 1.

Table 1

Fuel Pool

diameter

D(m)

Convective

heat flux

Qc(kW)

Flame Height

Zf(m)

Methanol 1.737 796 1.7

Hydrocarbon

Transformer Fluid

1.737 1546 3.7

Heptane 1.737 4367 5.5

The various correlations for the variation of centreline teroperature and

plume velocity were used following the recommendations of Beyler [6].

Accordingly, the correlations for the centreline tenperaure and the velocity

for the three regions viz. plumes above flame, intermittent and continuous

flame are described below:

Correlations for Pliimes Above Flame

Ihe centreline teirperaure and the velocity above the flame in a fire

plume is given as follows:

ATjj^ =22.0 Qc^/-' Z~^/^ ... (4)

Vj^ = 1.14 Q^^/^ Z"^/^ ... (5)

v^ere Tj^^ is the centreline terrperature rise (*»C) above the ambient

tenperature and Vj^^ is the centreline velocity (m/s) , and z is measured above

the virtual source (m) and not the height above the actual fuel source.

However, the virtual source location was estimated using the following

relationship [ 6 ]

.

(Zf - Zq) — = 0.12 for Zf/D < 1 ...(6)

Qc'

2/5

99
(Zf - Zq) — = 0.15 for Zf/D > 1 ...(7)

Qc

vi^iere Z^(ja) is the flame hei^t, and Z^im) the Icxiation of virtual source.

Correlations for Velocity and Temperature in Flame

The flame above a burning objec± can be divided into two regions for the

purpose of correlating centreline velocities and teroperatures, the

continuously flaming region and the intermittent region. Within the

continuously flaming region the centreline teirperture is approximately

constant and velocity rises as the square root of hei^t. The correlations

for continuously flaming region as proposed ty Beyler [6] were used and are

given below in a tabular form:

Table 2

^i^^ Vjn/Z^/'^

q2/^/D > 10.2 and 0.03 < Z/Q^^^ < 0.08 980 6.83

q2/5/d < 10.2 and 0.02 < Z/Q^/^ < 0.06 960 6.83

vdiere Q is total heat release rate (kW)

.

Correlations for the intermittent region as proposed by Beyler [6] are as

follows:

For the region Q^/^/d > 10.2 and 0.08 < Z/Q^/^ < 0.2

ATj(j =78.0 q2/^ Z~^ ... (8)

Vj^ = 1.85 Q^/^ ...(9)

For the region Q^/^/D < 10.2 and 0.06 < Z/Q^/^ < 0.12, the centreline

terrperatures are calculated from the correlations for plume above the flame,

vdiile the centreline velocity is calculated as given belcw:

Vjj^ = 1.65 Q^/^ ... (10)

The variation of the ambient terrperature and the plume velocity with

hei^t, vdiich are the iiput parameters for the water droplet vaporization, are

shown in Figs. 1 and 2. These curves have been made smoother in the matching

regions of the various plume zones by adjusting the various correlation

coefficients.

RESULTS AND DISCUSSION

In the present study, the time was non-dimensionalized with the

characteristic thermal diffusion time (in water) for a 2 mm water droplet (for

all the droplets) , so that the non-dimensional time is directly proportional

to the real time and the evaporation rates of various size droplets can be

compared.

100
HEIGHT (m) HEIGHT (m)

FIG. 1. PLUME TEMPERATURE VS HEIGHT FIG.2 . VELOCITY OF PLUME VS HEIGHT

Figures 3 to 5 show the teroporal variations of the falling water droplet

velocity with respect to ambient gas for weak, medium and strong plumes.

Initially the drcplet starts falling from zero velocil^ in a region vdiere

plume velocity is also small but in the opposite direction. As the droplet

continues to fall, its falling velocity and also the plume velocity increase.

Consequently the relative gas droplet velocity increases. This trend

continues iQ5to the start of continuous flaming region. As the droplet enters

the continuous flaming region, the plume velocity decreases v^ich results in

a lower gas-droplet velocity. In the region between the continuous flaming

region and the ground, the relative gas-droplet velocity goes on decreasing.

Thus the portion of the relative velocity curves after the maxima represents

the travel of droplets in a continuous flaming region of plumes. It may be

noticed from the figures that 2 mm drop can penetrate only weak plume. In the

case of medium plume after falling for some time, it is carried away by the

rising plume. In strong plumes it is not able to enter at all. However,

larger size drops are able to penetrate all the three plumes and reach the

seat of fire.

Figure 6 shows the total mass vaporized for various drops v\^en they fall

in various plumes from a hei<^t of 4.57 m. The 2 ram diameter drop is able to

enter only the week plume and vaporizes xipto 33% in the weak plume. The 3 ram drop is able to enter all the three types of plumes and vaporises upto 10, 15

and 35% in the weak, medium and strong plumes respectively.

101
(oas/ui) AilD013A

}~ I— I—

111

CO (NJ (Das/ui) AilD0T3A

COO

U.

5 5

r» to

va- cn

CO in t-»

II II II u u u O o O

o <

in

E ^ e

<

O

CO

I-Ul

_J

a

o

Q

I/)

f- a. <

go

a. a

< <

> >

CO 2

UJ ^2

o ^

u. o oCO o

NOIiVZiaOdVA SSVW »/o

o /

vt e-

\ V

-L. O 00 vj

(oas/ui) AiID0T3A

102
CONCLUSIONS

From the above stucJy it seems that the 2 mm droplet may not be the

optimum size droplet. This is so because -a 2 ram droplet can penetrate only the

weak plume v^iile in other plumes, it will be carried away by the hot gases.

Moreover, it evaporates i:pto 33% even in the weak plume, thereby the net mass

of water falling on the seat of fire is reduced considerably.

It is found that the droplets larger than 2 mm diameter can penetrate all

the three types of plumes. For a 3 ram droplet the reduction in mass is of the

order of 35% in the strong plumes.

In the present study the spray effects and the interaction between the

droplets have not been considered. In a spray, the vaporization of droplets

is modified because of lower tenperatures in sprays and also because of the

presence of water vapors in the ambient environment.

NOMENCLATURE

Cju coefficient of drag

D pool diameter (m)

g acceleration due to gravity

m mass of the falling water droplet

Q total heat release rate (kW)

convective heat flux release rate (kW)

Rq initial droplet radius

S frontal surface area of drop

t time

ATqq centreline terrperature rise above ambient (°C)

U velocity of the falling water droplet

Vj^ centreline mean upward velocity of the fire plume (m/sec)

Z height, measured above virtual source (m)

Zq location of virtual source (ra)

Zf flame hei^t (m)

Greek Symbol

ttj thennal diffusivity of the liquid (water)

"C t OL-^/'R^ , dimensionless time

103
f density

H c^namic viscosity of free stream

Subscript

g gas phase

1 liquid phase

0 initial value

00 ambient

REFERENCES

1. Harpole, G.M. , "Droplet Evaporation in High Ternperature Environment",

Jour, of Heat Transfer, Transaction of the ASME, Vol.103, p. 86-91 (1981).

2. Prakash, S. and Krishan, G. , "Convective Droplet Vaporization with

Transient Non-convective Liquid-Phase Heating", Twentieth International

Symposium on Combustion. The Combustion Institute, Pittsburg, PA,

p. 1735-1742 (1984).

3. Gakkhar, R.P. and Prakash, S., "Ifristeady Vaporization of Fuel Droplet

in a Convective Environment with Varying Ambient Conditions", Int.

Jour, of Heat and Mass Transfer, Vol. 33, p. 1003-1012 (1990).

4. Prakash, S. and Sirignano, W.A. , "Theory of Convective Droplet

Vaporization with Unsteady Heat Transfer in the Circulating LiquidPhase", Int. Jour, of Heat and Mass Transfer, Vol.23, p. 253-268 (1980).

5. Kung, H. and Stavrianidis, P. , "Buoyant Plumes of Large Scale Pool

Fires", Nineteenth International Symposium on Combustion . The Combustion

Institute, Pittsburg, PA., p. 905-912 (1982).

6. Beyler, C.L. , "Fire Plumes and Ceiling Jets", Fire Safety Journal,

Vol.11, p. 53-75 (1986).

7. Yuen, M.C. and Chen, L.W. , "On Drag of Evaporating Liquid Droplets",

Combustion Science & Technology, Vol.14, p. 147-154 (1976).

8. Reid, R.C. , Prausnitz, J.M. and Sherwood, T.K. , The Properties of Gases

and Liquids . McGraw Hill (1977)

.

9. Morsi, S.A. and Alexander, A.J. , "An Investigation of Particle

Trajectories in Two-Phase Flow System", Jour, of Fluid Mech. , Vol.55,

p. 193-208 (1972).

104
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 6

NEW RESEARCH APPROACH TO AIR SPRAY PAINTING

K.-C. Kwok* and B.Y.H. Liu^

*Graco, Inc., Minneapolis, MN, U.S.A.

^Particle Technology Laboratory, University of Minnesota, Minneapolis, MN, U.S.A.

ABSTRACT

Although spray painting is a widely used technique to apply decorative or protective coatings on a substrate,

very little scientific researches have been done to enhance the understanding of its deposition mechanism. This

paper first describes some parameters which are essential in controlling transfer efficiency (or deposition) and finish

quality. Then, the approach to study the particle size distribution and the deposition mechanism of air spray

painting is given. To illustrate the study approach, some experimental and theoretical results are also presented.

INTRODUCTION

Spray painting is an effective method for applying coatings on surfaces. In the past, the finish quality of a paint job was the primary concern. In recent years, environmental regulation has made it necessary to reduce solid

waste and volatile organic compounds (VOCs) emissions. This has led to the development of high-solids and

waterbome paints. The high cost of the new paints and the environmental issues have made it imperative to have

high transfer efficiencies in painting operation. In southern California, for instance, no spray painting equipment is allowed to be used if its transfer efficiency is less than 65%.

Throughout the spray painting industry, a great deal of practical experience has been accumulated over many

years, but very little systematic research has been done to understand the atomization process, and the transport and

deposition mechanisms of spray painting. Only when these process and mechanisms are fully understood there is the hope of making significant improvement in the transfer efficiency and finish quality.

IMPORTANT PARAMETERS

Through experience, it is known that particle size distribution (PSD) is closely related to finish quality and

transfer efficiency. Researchers in other fields also know that the PSD of airborne particles affects the deposition of

particles on surfaces, as for example, the deposition of inhaled particles in the human lung [1].

During spray painting, paint particles are driven towards the target and impact on it. For those particles that

fail to impact on the target, they become airborne. These are known as the overspray. The airborne particles will

eventually be deposited on the booth walls or be collected by the air filter in the exhaust. While the transfer

efficiency is routinely measured to gauge the performance of spray painting equipment, the amount of paint

deposited on the booth walls and collected by the exhaust are not measured. The last two pieces of information will

help determine the frequency of cleaning and filter change the spray booth may need.

For measuring the particle sizes of a spray, the conventional method is to collect samples on a microscopic

slide (or a substrate). Then the spot sizes of the paint particles are measured under a microscope. Unfortunately,

this sampling method only collects those panicles that can impact on the sUde, and the measured size distribution

represents only a part of the total PSD of the spray. If this result is used, erroneous conclusions can be reached.

Even though the total PSD can be estimated using instruments like the Malvern particle sizer, there is no

information on what kind of panicles will and can deposit on the target, the booth walls, or be collected by the

exhaust. The PSDs of these three particle fractions are important for understanding the aerodynamic motion of

particles in the spray booth. In addition, if the PSD in each of the three fractions is known, the total (or true) PSD

105
of the spray can be obtained by combining the three PSDs. With this information, it may be possible to find ways

to improve the transfer efficiency, reduce the paint deposit on the booth walls, improve the filtering efficiency of

the exhaust system, and eliminate the harmful respirable panicles.

AIR SPRAY PAINTING

Among all existing spray painting methods, air spray painting is the most widely used and versatile. The

advantage of air spray is its ability to change the degree of atomization without changing the paint flow rate, thus

allowing the operator to control both the finish quality and the coating speed. Another advantage of the air spray is that it can spray metallic paints with uniform and desirable shades of color. Hence, in view of the benefits listed

above, air spray is potentially the method to produce high finish quality and transfer efficiency. This is the

motivation for conducting a fundamental research on air spray painting [2]. The air flow field formed by the air cap of a spray gun is the main source of energy to drive the paint

particles forward and deposit them on the target. It is important to know the mean velocity structure of the air flow

field. For convenience, we shall refer to this air flow field as the "air-cap jet". Figure 1 illustrates the impingement

of an air-cap jet onto a target. The mean velocity profiles shown represent approximately the true profiles of the

air-cap jet. At a short distance from the air cap, the jet expands rapidly in the X-Z plane. The cross-section of the

jet is approximately in the form of an ellipse. As the jet travels forward, the ellipse is further elongated.

The velocity profiles in the plane of the minor-axis resemble the shape of a bell. The peaks occur at the

centerline of the jet. However, the peak decreases and the profile spreads out and becomes wider as the jet travels

downstream. In the major-axis plane, the velocity profiles look like the cross-section of a saddle-back. The peak

velocities of each profile are not at the centerline but near the two edges of the profile. As the jet travels farther

downstream, the length of the profile increases. Also, the difference between the peak and the centerline velocities

decreases. After impinging on the target, as shown in Figures 1(a) and (b), two walls jet are produced. The wall jet

produced from the ends of the jet is called the radial wall jet while that produced from the long side of the ellipse is called the plane wall jet

NEW EXPERIMENTAL APPROACH

To measure the important parameters affecting spray painting, a new experimental approach has been

developed [2]. The main idea is to construct a miniature spray booth which is designed to measure simultaneously

the paint deposit on the target, on the walls, and in the exhaust. Furthermore, the measurement of the PSDs on the

booth walls and in the exhaust can be done in the same experiment.

Figure 2 shows the basic construction of the miniature spray painting booth^. The booth was designed to be

a rectangular box with 0.61 m x 0.91 m (24" x 36") floor, and an interior height of 0.61 m (24"). For the sake of

clarity, the vertical walls are not shown. Two large panel filters were put on the top of the booth to filter the

incoming air. In the middle of the boodi, a target panel was hung. The gun-to-target distance could be adjusted. On

the floor, an 0.203 m x 0.086 m (8" x 3.375") rectangular hole was cut, which was the exhaust outlet of the

booth. Connected to the exhaust oudet was a 5-stage Hi-Flow Cascade Impactor (HFCI) which draws 0.0188 mVs

(40 cfm) of air. Cascade impactor is a device designed to classify airborne particles according to their aerodynamic

particle size. The HFCI was designed to classify particles with diameter below 20 |xm. The 5 stages have stage cut- points of 16.30, 9.87, 5.58, 3.10, and 1.62 |im. A filter was used to collect particles less than 1.62 \un.

To start an experiment using the miniature spray booth, the spray gun was mounted on a stand in front of

the booth. Only the front end of the gun was allowed to slick into the booth. Before an experiment, the target

panel was weighed and the HFCI was prepared. Then, the blower for the HFCI was turned on and the spray gun was triggered for about two seconds. After the gun stopped spraying, the blower was allowed to run for about 1.5 minutes. This ensured all airborne panicles were evacuated. To calculate the transfer efficiency, the net weight of

paint on the target, the percentage of paint solid, and the total paint output from the gun were recorded. The transfer

efficiency is calculated by dividing the net dry paint on the target by the total dry paint output from the gun. The

net weight gain by each stage of the HFCI was used to construct the mass distribution of the airborne paint

particles. The summation of the weights of all stages plus the filter gives the total weight exhausted by the booth.

The amount of deposit on the booth walls can be calculated by subtracting the sum of the weights of the target and

the exhaust from the total weight output from the gun.

^ The miniature spray booth is covered by an University of Minnesota patent (pending) licensed to the MSP

Corporation. The equipment is available commercially (MSP Corporation, 1313 Fifth Street S. E., Minneapolis,

MN 55414).

106
Illllllil

u \

X

Air Cap

Air "Horn"-

Atomization

Region

Calculation Domain

Radial WaU

Jet

llllllllll

W

Stagnation

Point

(a) Front View (b) Side View

Figure 1. Schematic IMagram of Impinging Air-Cap Jet

Panel Air

Filter

Air Spray

Gun

Pump On/Off

Switch

Pressure Gauge

Flow -\

Control \

Box

Pump

for HFCI

Target Panel

HFCI

Figure 2. Miniature Booth for Spray Painting Tests

107
PARTICLE SIZE DISTRIBUTIONS (PSDs)

As explained in the last section, the mass distribution of the airborne particles was measured by the HFCI.

To measure the PSD on the booth walls, microscopic slides were placed on the walls to collect paint particles for

future analysis. To measure the PSD on the target, the particle sampling shutter was used to collect particle

samples. The detailed construction and operational principle of the shutter has been given by Kwok [2]. To facilitate the measurement of the sizes of particles collected on the slides, an automatic image analyzer was used. This analyzer consisted of a microscope fitted with a motorized table. The slide was mounted on the table

and the particle images were viewed by a video camera. With the images displayed on a video monitor, the analyzer

could measure 500 images per second. This fast rate of image measurement allowed a large number of particles to be analyzed in a reasonable length of time. This improved the accuracy of the statistical analysis of the particles.

Furthermore, the motorized table and a computer program allowed many measurements to be done automatically.

The measured information was stored on a diskette for future statistical analysis.

The paint particles collected on the slides were flattened due to inertial impact Although the diameter of the

flattened particles is believed to be better related to the finish quality, for analysis of the aerodynamic particle

motion, the spherical diameter is needed. A new method has been developed to correlate the flattened diameter and

the spherical diameter of the paint particles [2]. The main idea of the method is to collect flattened particles on the

edges of the microscopic slide. The side profiles of the particles collected in this manner can be analyzed using the

automatic image analyzer. If the height and the flattened diameter of a particle are known, the volume of the

flattened particle can be calculated. Then the spherical diameter can be caJculated by equating the volume of the

flattened particle to the same volume of a spherical particle. The correlation of the flattened diameter (df) and die

spherical diameter (dp) is shown in Figure 3. MATHEMATICAL MODEL

The air-cap jet is an elongated ellipse which is a three-dimensional jet To calculate the three-dimension flow

field, even using a supercomputer, die computational effort and the data storage requirements are quite substantial.

Since the ratio of the major-axis length to the minor-axis width was about 8, the air-cap jet was modeled as a twodimensional plane jet The velocity profiles in the minor-axis plane (see Figure 1(b)) were assumed for the entire

jet. Furthermore, only the region near the wall is critical for panicle impaction so the calculation domain was

selected as shown in Figure 1(b).

The air-cap jet is quite turbulent and many turbulence models have been proposed to calculate turbulent flow

fields [3, 4]. Among these models, the k-£ turbulence model is the most popular and extensively tested. Therefore,

it was selected for calculating the air flow field of the impinging air-cap jet

Once the air flow field was known, the trajectories of a range of particle sizes could be calculated [5, 6]. The

calculation procedure was quite simple. First, a particle of known size was placed at an arbitrary position along the

starting line of the flow field and the trajectory was calculated. If the particle impacted on the target, it was moved

to a position farther away from the centerline for the next calculation. If the particle did not impact on the target,

the particle position was moved closer to the centerline. This process was repeated until a position, called the

critical position, was found such that the particle just impacted on the edge of the target. The percentage of flow

rate at that critical position was the collection efficiency for that particle size. After the collection efficiencies for a range of sizes were obtained, a curve of collection efficiency versus particle diameter was constructed. It should be

pointed out that the aerodynamic motion of a particle is better represented by the aerodynamic diameter rather than

die geometrical particle diameter. The relationship between the two diameters is d, = -/p^dp where dt = aerodynamic

diameter, pp = particle density, and dp = geometrical particle diameter. Using the information of the measured total PSD of the spray and the collection efficiency curve, the PSD on the target was calculated.

SAMPLE RESULTS

Figure 4 shows a plot of the paint deposit versus target distance for the target, the walls, and the exhaust.

As shown, the percentage of paint deposit on the target (i.e. transfer efficiency) decreased linearly with increasing

target distance. The slope of die line is -1.35% per cm (-3.44% per inch). This means that the transfer efficiency

decreased 1.35% for every centimeter the gun was moved away from die target. In normal spray painting, the gunto-target distance can vary between 20.3 cm to 30.48 cm (8" to 12"). As a result, die transfer efficiency can vary by

as much as ± 7% by diis factor alone.

The particles collected in the exhaust can be considered as airborne and an undesirable air pollutant, if emitted to the ambient atmosphere. They arise because they cannot impact on the target or settle on the walls. The

bottom line in Figure 4 shows only slight increment in weight wiUi a large increase in target distance. This means

108
B

ci. o

80

70

60

3 50

Q

13uI

CO

40

30

20

10

0

0

0 .—

«

dp = 2.5949 + (0.80341)df - (0.004157)d? + (1.43142 x 10-5)d?

... 1 ... 1 ... J ... 1 ... 1 .. . 1 . . . 1 . . . 20 40 60 80 100 120 140

Flattened Diameter (df), Htn

Figure 3. Correction between Spherical and Flattened Particle Diameter

160

8.

c

a.

100

90

80

70

60

50

40

30

20

10

0

Air to Paint Mass Ratio = 1.42 Paint Viscosity = 57 x 10"* m^/s

Paint Row Rate =330 gm/min.

a Target

Exhaust

o Walls

-L. 0 10 15 20 25 30 35 40

Target Distance, cm

Figure 4. Effect of Target Distance on Paint Distribution on Target, Walls, and Exhaust

60 80 100 120

Aerodynamic Diameter (da), [im

Figure 5. Mass Size Distributions on Target, Walls, and Exhaust

109
100

80

E

I

60

13>O 40

20 h

0

0

A

A

A

Air Pressure = 262 kPa

Target Distance =25.4 cm

Calculated

A Measured

J__l l__L. I I I t I I

10 15 20 25 30

X, cm

Figure 6. Comparison between Calculated and Measured Centerline Velocities

A

1^

A

A

6

A

Air Pressure = 262 kPa

Target Distance =25.4 cm

Calculated

Measured

A

± 9^ .

0 20 40 60 80 100

Aerodynamic Diameter (da), jim

120 140

Figure 8. Comparison of Measured and Calculated MSD on Target

110
that increasing target distance did not cause a substantial increase in airborne particles produced. Instead, most of

the particles that could not impact on the target end up on the walls. This is evident from the steeper positive slope

of the middle line in Figure 4. Figure 5 shows the mass size distributions of a spray at 25.4 cm target distance. The dotted line is the total distribution which was constructed by summing all the other distributions. Also, it should be pointed out that the area under each curve represents die total mass for that curve. As can be seen, the particles deposited on the target

had sizes ranging from a few micron to about 120 |im. For all practical purpose, the significant particle sizes are between 15 to 80 fim, and they are the most important particles affecting the finish quality of a coating. The peak

of the curve is located at an aerodynamic diameter of approximately 50 |jjn. Coincidentally, the mass median

aerodynamic diameter (MMAD) was also found to be 50 jim. The MMAD is defined as the aerodynamic diameter

for which 50% of the mass is less than the stated aerodynamic diameter.

From the mass distribution curve for the walls, it can be seen that the largest particle size was about 70 |im.

This means that any particles bigger than 70 (xm were all collected by the target. The peak of the mass distribution curve was about 25 jim and the MMAD was about 30 fxm. According to the settling velocity calculation, a 30 |im

particle will settle (due to gravitational force) at 2.72 cm/s. This is quite a fast settling rate, indicating that most

particles on die floor was probably deposited by settling. However, the smallest particle found on the microscopic

slides was about 2.5 pjn. The settling velocity for this size is quite small, only 0.021 cm/s. It is unlikely that the

2.5 [un particles collected were due to settiing. One possible deposition mechanism is turbulent deposition.

The mass distribution curve for the exhaust shows that the majority of the airborne particles was about

8 Jim and the MMAD was about 12 \m. Further, the curve also shows that particles above 35 |im could not

remain airborne. For a 12 |im particle, the settling velocity is 0.49 cm/s.

The total distribution curve shows that the mass distribution is bimodal. One peak is located at an aerodynamic diameter of 25 \im and the other peak at an aerodynamic diameter of about 50 \im. The comparison

between the total mass distribution and the mass distribution on the target immediately tells about the sizes and the

amount of particles that cannot impact on the target. This information can be used to guide equipment design and

improvement. For example, if the of an equipment is so designed that the inertial energy of those particles between

10 and 60 |im are increased, the transfer efficiency will be improved substantially.

To show the validity of the matiiematical model described earlier, some of the measured and the calculated

results are compared. The measurement method and data for the mean air velocities have been given in Kwok's

thesis [2]. Figiu^e 6 compares the measured and calculated centerline velocities as a fimction of axial distance for the

262 kPa air pressure. It can be seen that the agreement between the measured and the calculated data points is excellent. Also, the air-cap jet showed rapid decay of centerline velocity in the first 10 cm from the air cap. Then

the decay rate is reduced. At about 5 cm from the target (X=25.4 cm is the target location), the centerline velocity

decelerated quickly to zero. The mean velocity profiles in the major-axis and the minor-axis planes have been shown and explained in

Figure 1. So, they will not be repeated here.

Figure 7 shows three calculated and measured velocity profiles of the plane wall jet for three Y distances.

The two results show fair agreement. In fact, for Y=17 cm, the calculated velocity profile fit is in excellent

agreement with the measured velocity profile. However, for the other two Y distances, the "widths" of the

calculated profiles are slighUy larger than the measured ones. The calculated and measured mass distributions on the target for the 262 kPa air pressure and 25.4 cm (10")

target distance are shown in Figure 8. In general, the agreement between the calculated and the measured values is excellent for aerodynamic diameters larger than approximately 30 \im. Below 30 \xsr\, the calculated value is always

less than the measured value. The smaller the aerodynamic diameter, the bigger is the deviation in Amass/Ada. The

explanation for the deviation is that the particle number concentration was very high for particles with aerodynamic

diameter less Uian 30 |im. In fact, these particles constituted 98% of all particles by number. So, the flow situation was probably a dense two-phase flow which was not included in the mathematical model.

CONCLUSIONS

Air spray painting, though an old painting method, can be improved to give high transfer efficiency and

good finish quality through fundamental understanding of the atomization, transport and deposition process. The

independent control of the air flow and pressure makes air spray the preferred and versatile atomization method for

obtaining fine finish coatings.

New research approach has been suggested and explained. This approach allows the measurement of paint

deposition on the target, the walls, and Uie exhaust. Furthermore, the PSD in each of these three places can be

measured and the total PSD constructed. This infonnation is necessary in order to better understand the aerodynamic

motion and the deposition mechanism of the paint particles. Even though the aerodynamic motion of the particles

111
is difficult to study experimentally, their behavior can be studied using modem mathematical and computational

tools.

REFERENCES

1. Biirkholz, A., Widder, J., Miiller, W. and Dislich, M., "Particle Size Distribution of Overspray Generated

During Spray Painting", Polymers, Paint and Colour Journal, Vol 167, pp.285-290 (1977).

2. Kwok, K. C, "A Fundamental Study of Air Spray Painting", Ph. D. thesis. Particle Technology

Laboratory, University of Minnesota, Minneapolis, MN (1991).

3. Rodi, W., Turbulence Models and Their Application in Hydraulics. International Association for

HydrauUc Research (lAHR), Delft, The Netherlands(1980).

4. Launder, B. E. and Spalding, D. B., Mathematical Models of Turbulence. Academic Press, London

(1972).

5. Marple, V. A., "A Fundamental Study of Inertial Impactors", Ph.D. thesis. University of Minnesota,

Minneapolis, MN, Particle Technology Laboratory Publ. 144 (1970).

6. Rader, D. J. and Marple, V. A., "Effect of Ultra-Stokesian Drag and Particle Interception on Impaction

Characteristics", Aerosol Science and Technology. Vol 4, pp.141-156 (1985).

112
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 7

COATEVGS APPLIED WITH THE PLASMA-SPRAY PROCESS

USING NICKEL-ALUMINUM POWDERS

D.J. Varacalle, Jr.', J.R. Knibloe*, V.L. Smith-Wackerle*,

J.A. Walter^ and G. Irons*

*Idaho National Engineering Laboratory, EG&G Idaho, Inc., Idaho Falls, ID, U.S.A.

^Miller Thermal Inc., Appleton, WI, U.S.A.

*TAFA, Inc., Concord, NH, U.S.A.

ABSTRACT

As part of an investigation of the dynamics that occur in the plasmaspray process, an experimental study of the plasma spraying of nickel -aluminum

powder was conducted. The coating experiments used typical process parameters

that were varied in a Taguchi fractional factorial design parametric study.

Operating parameters were varied around the typical spray parameters in a

systematic design of experiments to display the range of plasma-processing

conditions and their effect on the resultant coating. Parameters varied

include the working gas mixture, power, powder injection location, spray

distance, traverse rate, and powder feed rate. The coatings were

characterized by hardness tests, optical metallography, and x-ray diffraction.

Coating qualities are discussed with respect to thickness, hardness, porosity,

and phase content.

INTRODUCTION

The plasma- spray process is one of the most versatile and rapid methods

of applying coatings, and its applications continue to increase [1]. The

plasma- spray guns typically use a nontransferred dc plasma torch configuration

with powers up to 100 kW. The basic design of the plasma gun consists of a

cone-shaped cathode inside a cylindrical anode nozzle. An inert gas (usually

argon, argon/hydrogen, or argon/helium) flows through the space between the

electrodes where it is ionized to form a plasma. Powders injected into the

plasma are then accelerated and melted by its high temperature. The molten

droplets (i.e., splats) are propelled onto the substrate where they solidify

and accumulate to form a thick, well-bonded coating. For quality coatings,

the particles must absorb enough heat from the flame to melt thoroughly, but

not so much that they overheat, oxidize, and/or vaporize. At the same time,

the particles must travel fast enough to have good particle adherence once

they strike the substrate.

Plasma- sprayed coatings are used for wear resistance, heat and oxidation

resistance, corrosion resistance, electrical or thermal conductivity or

resistivity, restoration of dimension, and clearance control. Many

operational parameters control the quality of the coating, making it difficult

to find the optimum processing conditions for any specific application [2,3].

As coatings become more sophisticated, a better knowledge of the

underlying principles is necessary to improve process control and coating

quality. Plasma spraying was developed to a large extent by empirical means,

while the scientific research lagged behind. This work attempts to further

113
the scientific understanding of the physical mechanisms involved in

nickel -aluminum coating formation by determining which process parameters

affect the structure and properties of the coatings.

Plasma- sprayed nickel -aluminum coatings are widely used in the

automotive, transportation, aerospace, and aircraft industries. The

nickel -aluminum powder system is used for restoration of dimension, wear, and

oxidation and as bond coats for thermal -sprayed ceramic coatings to enhance

the adhesion strength of the coating to the substrate.

EXPERIMENTAL PROCEDURE

A Taguchi [4] statistical -style fractional -factorial design experiment

was employed to evaluate the effect of seven plasma-processing variables on

three quantitatively measured responses. The parameters varied were amperage,

primary gas flow, secondary gas flow, powder feedrate, spray distance, gun

traverse rate, and powder injection angle. The resulting responses evaluated

were thickness and porosity with optical microscopy and hardness with a

Rockwell 15T test. Phase content was determined with x-ray diffraction (XRD)

.

Experiments NA15 through NA22 in Table 1 represent the eight runs

evaluated with the Taguchi approach. Two levels of each variable were

selected to band around the nominal settings (i.e., experiment NAOl) . The

quantitative Taguchi evaluation of the plasma- spray process is ideal because

it displays the range of measured coating characteristics attainable, and it

statistically delineates the impact of each factor on the measured coating

characteristics across all combinations of other factors. The Taguchi

analysis used PC-based software [5] on the measured responses.

A Miller Thermal Plasmadyne SG-100 plasma- spray system was used the

experiments in Table 1 . The primary gas was argon and the secondary gas was

helium. A four hole gas ring induced vortex flow. The powder carrier gas was

argon at 0.000102 m^s (13 scfh) for all experiments. The powder injection

was internal to the torch and was directed either perpendicular to the flow or

at a 20 -degree angle forward into the flow. An x-y manipulator ensured the

spray distance and repeatability in the experiments. Traverse rates varied

from 43.2 to 54.8 cm (17 to 23 in.)/s with a y-step of 0.3175 cm (0.125 in.)

/step. Ten passes were used to fabricate each of the coatings.

Table 1 Nickel -Aluminum Thermal Spray Experiments

Experiment Number

Current

(amps)

Power

(kW)

, Prima r> (m /s) (sc

Gas

fh)

Flow Secondary

(m^/s) (scfh)

Powder

Feed

(m^/s-lb/h)

Spray

Distance

( cm/ in.)

Gun Traverse

(cm/s-in/s)

Inject

(p or f)"

NAOl 800 30.9 0.000787 100 0.000370 47 0.00088 7 7.62 3.0 50.8 20 f NA15 700 23.8 0.000590 75 0.000236 30 0.00063 5 6.35 2.5 43.2 17 P NA16 900 39.2 0.000983 125 0.000472 60 0.00114 9 6.35 2.5 43.2 17 P NA17 700 27.4 0.000590 75 0.000472 60 0.00114 9 8.89 3.5 58.4 23 P NA18 900 35.5 0.000983 125 0.000236 30 0.00063 5 8.89 3.5 58.4 23 P NA19 700 27.7 0.000983 125 0.000236 30 0.00114 9 6.35 2.5 58.4 23 f NA20 900 33.0 0.000590 75 0.000720 60 0.00063 5 6.35 2.5 58.4 23 f NA21 700 30.5 0.000983 125 0.000720 60 0.00063 5 8.89 3.5 43.2 17 f NA22 900 29.3 0.000590 75 0.000236 30 0.00114 9 8.89 3.5 43.2 17 f

a. p = perpendicular to flow, f = 20 degrees forward into flow.

114
The study used a commercially available powder (Alloys International

AI-1037), consisting of a nickel core with an aluminum skin. The powder

chemistry and sieve analysis, as supplied by the powder manufacturer, are

shown in Table 2. The powder was plasma- sprayed onto 304 stainless steel

plates (51 X 63 X 3 mm) cooled by air jets on the back side. One side of each

steel coupon was grit-blasted with #30 alumina grit before spraying.

Table 2 Nickel -Aluminum Powder Chemistry and Sieve Analysis

Chemistry

Ni

Al

Sieve

% (pm) %

95.13 >106 0.1

4.87 > 90 1.3

> 45 99.8

MATERIALS CHARACTERIZATION RESULTS

Conventional air plasma spraying (APS) yields rapid solidification and

is considered a continuous, metastable materials forming technique. However,

APS is known to produce rapidly solidified structures that show defects, such

as unmelted particles, porosity, and oxide inclusions, which modify the

coating thermophysical properties and affect bond strength and toughness. The

microstructure of plasma- sprayed deposits is based on the rapid solidification

of individual splats. The splat morphology depends on the processing

parameters that dictate the particle velocity, temperature, and size at the

standoff distance.

Average coating thicknesses for each experiment, taken from 12

measurements using optical metallography, are listed in Table 3. The deposits

ranged from 187 to 603 fim for an average of 19 to 60 |im per pass, reflecting

influences of spraying parameters. Controlled deposition is important to the

substrate heat input that influences the residual stress buildup within the

coating. Also, deposition considerations substantially impact the economics

of the plasma- spray process.

Table 3 Coating Characterization Results

Experiment Ni-Al thickness Porosity

Number (um) Hardness''

NAOl 421 15T76.5 2.6

NA15 309 15T80.0 5.9

NA16 601 15T83.6 1.8

NA17 435 15T80.4 4.1

NA18 187 15T80.9 3.5

NA19 476 15T75.6 1.1

NA20 280 15T84.0 4.2

NA21 310 15T81.1 3.8

NA22 603 15T78.0 3.0

a. Rockwell 15T hardness

115
Variances in porosity were evident by optical microscopy. Average

porosities from four measurements per coating for each experiment are listed

in Table 3. The average porosity of the nickel -aluminum coatings ranged from

1.1 to 5.9%. Figs. 1 and 2 illustrate the microstructure of nickel -aluminum

coatings from experiment NA15 (highest porosity coating) and experiment NA19

(lowest porosity coating). The NA15 coating shows a significant amount of

oxide stringers and a somewhat higher NiO content (measured by XRD) than the

NA19 coating. Unmelted particles are more evident in the NA19 coating.

Typically, low porosity is desirable in nickel -aluminum coatings.

Hardness measurements were taken on the coatings normal to the deposit

using the superficial Rockwell 15T method. Twelve measurements were taken and

averaged. The hardness of the coatings varied from 75.6 to 84.0, as listed in

Table 3. Higher superficial hardness of a coating may be an advantage for

wear applications, but is not necessary when used as a bond coat.

X-ray diffraction measurements were taken to qualitatively assess the

phases in the coatings from the nine plasma- spray experiments. The feedstock

powder and all as -sprayed coatings exhibited XRD spectra indicating that the

samples primarily contained elemental Ni. This is not surprising since the

source powder had less than 5% Al . Results from the XRD studies are presented

in Table 4. The coatings

contained minor amounts of various

nickel -aluminides (Ni3Al, Al3Ni2,

NiAl) and dispersed A1203 and NiO.

A previous study [6] also reported

primarily nickel coatings with

some oxides for a similar source

powder, but no aluminides were

detected.

DISCUSSION OF TAGUCHI FRACTIONAL

FACTORIAL EXPERIMENT DESIGN

Taguchi-type fractional

factorial testing is an efficient

means of determining broad-based

factor effects on measured

attributes. This methodology

statistically delineates the

impact of each variable on the

measured coating characteristics

across all combinations of other

factors. It does not consider

second or third order

interactions, but it does

determine which factors have the

most influence on the measured

attributes. The factor

significance lies only between the

levels of test parameters selected

and used. Also, the test methods

used to evaluate the attributes

must be consistent and accurate.

116
,4

J

If

V

r<;

Fig.

NA19

2 Microstructure (400X) for coating

The spray tests were conducted

and evaluated once, and all data

points were considered in the

analysis of variance (ANOVA)

calculations. The rho percent

calculation indicates the

influence of a factor or parameter

on the measured response, with a

larger number indicating more

influence. The error calculation

is the percentage of the variation

observed in the test due to the

data scatter. The confidence

interval is determined by

comparing the rho percent with the

error percentage of the response

test. Thus, a parameter with a

low rho percent that is

significant at a 95% confidence

level indicates that the effect is

probably real, but relatively

insignificant when compared to

other parameters . The ANOVA

calculations guide further

experimentation by indicating

which parameters are the most

influential on coating attributes.

This information is extremely

useful in developing new coating

specifications and spray

equipment

.

Table 4 Coating X-Ray Diffraction Results

Experiment

Number Ni Ni3Al Ni2A13 NiAl NiO A1203

Powder sb s s t^

NAOl P t s t t

NA15 P t t t t t

NA16 P t t s t

NA17 P t s t

NA18 P t t s t

NA19 P t s t

NA20 P t t t t

NA21 P t s t t

NA22 P t t t t t

a. p = predominant phase

b. s = secondary phase

c. t = possible trace of phase

117
The analysis evaluated the effect of seven plasma-processing variables

on three quantitatively measured responses . The variation in the coating

characteristics displayed in the experiments indicates that minor variations

in plasma-processing conditions can affect coating properties. Table 5

illustrates the results of the Taguchi analysis.

Table 5 Results of the Taguchi Analysis

Factor

Rho Confidence Level Level Average

Percent i%} High Low

Superficial Hardness (Rockwell 15T) (12 Data Points)

Helium Flow

Amperes

Feedrate

Injection Angle

Error

37.2

15.0

11.6

5.8

29.4

99

99

99

95

82.3

81.6

79.4

79.7

78.6

79.2

81.5

81.2

Porosity (%) (4 Data Points)

Feedrate

Argon Flow

Injection Angle

Error

32

26,

5,

35

99

99

95

2.6

2.5

3.0

4.3

4.4

3.9

Thickness (nm) (12 Data Points)

Feedrate

Gun Traverse Rate

Amperage

Injection Angle

Spray Distance

Error

75.1

13.9

1,

1,

1,

7,

99

99

95

95

95

528

345

419

417

384

272

455

384

384

417

The Taguchi evaluation indicated that secondary helium flow was the most

significant contributor to hardness at a 37.2 rho percent with higher helium

flow resulting in greater hardness. Other contributors at the 99% confidence

level were amperage at a 15.0 rho percent with higher amperage resulting in

higher hardness values, and powder feedrate at 16.7 rho percent with lower

feed rates resulting in higher hardness values. At the 95% confidence level,

the injection angle had a small, yet probably real, contribution of

6 rho percent with rearward injection leading to greater hardness. The error

within the test accounted for 29.4% of the variation observed in the test.

At the 99% confidence level, the most significant contributors to

porosity were powder feedrate and primary flow with rho percentages of 29 and

24%, respectively. Higher powder feedrate and higher argon flow contributed

to lower porosity. A small but real contribution to porosity was made by an

injection angle with a rho percent of 5.1% at a 95% confidence level. The

error in this response test accounted for 35.5% of the variation observed in

the porosity data.

118
The Taguchi analysis of this study indicated that the key to maximizing

the economics of the coating deposition was controlling powder feedrate and

gun traverse rate. Their contribution to the total variation in thickness

observed (i.e., rho percent) was 75.1 and 13.9%, respectively, at a 99%

confidence level. Increasing the powder feedrate increased the deposition

thickness, as would be expected. Lower gun traverse rates lead to increased

deposition thickness. The error percentage for the thickness response test

was low at 7%, which made the small contributions of amperage, injection

angle, and spray distance significant at a 95% confidence level.

This Taguchi evaluation is significant because it directs further

experimentation on the most important process or coating attributes and the

process parameters that affect these attributes. The most important

attributes may differ for the same material in different applications, and the

baseline data generated in this study can be used to develop specific

confirmation runs that approach the desired application attributes.

SUMMARY AND CONCLUSIONS

This paper presents an experimental study of the plasma- spraying of

nickel -aluminum powder coatings. Experiments were accomplished using a

Taguchi fractional -factorial approach varying typical process parameters. The

coatings were then characterized for thickness, porosity, hardness, and phase

structures using optical microscopy and x-ray diffraction. The coating

thicknesses, reflecting influences of spraying parameters, ranged from 187 to

603 |im. Porosity of the coatings ranged from 1.1 to 5.9%. Hardness

measurements varied from 75.6 to 84.0 on the Rockwell 15T scale. Results from

the XRD coating studies indicated a predominantly nickel matrix with various

nickel -aluminides (Ni3Al, A13Ni2, NiAl) and dispersed oxides (A1203 and NiO)

.

The Taguchi analysis indicated that powder feedrate and traverse rate

were the significant factors affecting the deposition thickness. The

significant factors for hardness were secondary flow and amperage. Powder

feedrate and primary flow significantly affected porosity.

The objective of this and future work is to optimize the quality of

plasma- sprayed coatings. By generating baseline data on factors that

influence coating characteristics, operational parameters can be selected and

optimized for future nickel -aluminum plasma-spray processing experiments and

applications. Future work is needed to determine the effect of powder

fabrication methods on the coating morphology.

ACKNOWLEDGEMENTS

The technical input of G. C. Wilson and A. W. Erickson is gratefully

acknowledged. The work described in this paper was supported by the U.S.

Department of Energy under DOE Contract No. DE-AC07 - 76ID01570

.

119
REFERENCES

1. Kubel, E, D., Advanced Materials and Processes, (132), 6, 69-80, (1987).

2. Pfender, E., "Fundamental Studies Associated with the Plasma Spray

Process," Proceedings of the National Thermal Spray Conference, September

14-17, 1987, Orlando, Florida.

3. Fauchais, P., "State of the Art for the Understanding of the Physical

Phenomena Involved in Plasma Spraying at Atmospheric Pressure,"

Proceedings of the National Thermal Spray Conference, September 14-17,

1987, Orlando, Florida.

4. Taguchi, G., and Konishi, S., Taguchi Methods: Orthogonal Arrays and

Linear Graphs, ASI Press, (1987).

5. Gulp, R. F. , Sadie, Speedy Analysis and Design of Industrial Experiments,

Copyright 1990, R. F. Gulp.

6. Sampath, S., Herman, H. , and Rangaswamy, S., "Ni-Al Re - evaluated,

Proceedings of the National Thermal Spray Conference, September 14-17,

1987, Orlando, Florida.

120
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper

ATOMIZATION CHARACTERISTICS OF A

HIGH-SPEED ROTARY-BELL PAINT APPLICATOR

P.L. Corbeels, D.W. Senser and A.H. Lefebvre

School of Mechanical Engineering, Purdue University

West Lafayette, IN, U.S.A.

ABSTRACT

An experimental study was conducted on the influence of fluid properties and operational parameters on

the atomization characteristics of a high-speed rotary-bell paint applicator. Drop size measurements near the

bell edge were complimented with high speed photography to investigate the filming characteristics,

atomization modes, and hgament disintegration mechanisms of this type of applicator. Surface tension was

varied by a factor of 2.5 and viscosity by a factor of 100. Rotational speeds ranged from 4,000 rpm to 40,000 rpm, while the liquid mass flow rate was varied from 0.67 to 6.67 g/s. It was found that high

viscosity fluids film the bell very evenly and produce long regular ligaments, whereas low viscosity fluids

film incompletely and produce very irregular ligaments that disintegrate near the bell edge. Mean drop size was found to be fairly insensitive to large changes in flow rate and viscosity at bell speeds higher than

20,000 rpm.

INTRODUCTION

Rotary-bell atomizers are widely used in spray coating systems and generally provide excellent finishes

and high transfer efficiencies. Bells are generally preferred to cups because they provide a larger fluid

filming area for a given diameter. Consequently, a smaller diameter bell can be used without loss of

atomization quality, resulting in a smaller spray pattern that is more easily controlled. A typical rotary-bell

atomizer is shown in Figure 1. The coating material is delivered to the inside surface of the rotating bell

by an array of fluid feed ports. Under the influence of centrifugal and Coriolis forces the fluid films toward

the bell edge, generally decreasing in thickness as it move radially outward. Near the bell edge, radial

serrations promote the formation of discrete ligaments which subsequently disintegrate into droplets. An

axial flow of shaping air from a ring of small orifices positioned behind the bell imparts axial momentum

to the droplets, directing them towards the workpiece. In addition, a high voltage is applied to the bell

directly, or to a series of electrodes in the vicinity of the bell, to impart an electric charge to the liquid. The

resulting charged droplets are attracted to the grounded workpiece, thereby increasing transfer efficiency. For

paint spray applications, rotational speeds in excess of 20,000 rpm, and small bell diameters of less than 10

cm, are typically used. This combination of high rotational speed and small bell diameter provides good

atomization and a compact spray pattern.

While rotary atomization in general has been the subject of considerable research, as evidenced in the

reviews by Matsumoto et al. [1], Masters [2], Dombrowsici and Munday [3], and Marshall [4], few studies

have addressed high-speed, rotary-bell paint applicators (HSRBPAs). Hinze and Milbom [5] established that rotary atomizers produce drops by one of three principal mechanisms; atomization by direct drop formation,

by ligament formation, and by sheet formation. Small serrations near the bell edge allow HSRBPAs to operate almost exclusively in the ligament formation mode. Under most conditions the number of

ligaments is constant and equal to the number of serrations on the bell edge. The Sauter mean diameters

(SMDs) produced by these applicators range from 15 to 50 ^m under typical operating conditions. These

values are quite small in comparison with those obtained by Dombrowski and Lloyd [6] who performed an experimental study on the factors that control the spray characteristics of spinning cup atomizers, ranging in diameter from 50 to 150 mm. Most of their work dealt with atomization by film formation but a limited

number of measurements were made in the ligament mode. Within this mode, the measured SMDs ranged

from approximately 30 to 300 ^m. However, the maximum rotational speed of 10,200 rpm is slow in comparison to typical operating conditions for HSRBPAs. Contrary to general experience, they observed a

decrease in drop size as the liquid viscosity was increased from 7.3 to 101 mNs/m^. They attributed this

121
anomaly to more viscous ligaments stretching farther and becoming smaller in diameter before finally

breaking up into droplets.

Kayano and Kamiya [7] developed an analytical expression for the mean size of drops produced by a

rotating disk when operating in the ligament mode. Their calculations were based on a analysis which

predicted the number of ligaments produced at the disk edge and on the drop sizes produced by the

disintegration of these ligaments into drops. They assumed the ligament diameter to remain constant with

distance from the bell edge. Their model predicted drop size to increase with increasing surface tension and

with increasing viscosity. Comparisons of predicted and measured drop sizes in the range from 100 to 300

|im showed good agreement at all test conditions for disks of 3 to 7 cm in diameter and rotational speeds

from 2,500 to 6,000 rpm.

Frost [8] defined conditions for the occurrence of drop formation by ligament formation and derived an

expression for the prediction of mean drop size in the ligament mode. Comparison with measured values of

mean drop size for disks of 40, 60, 80, 100 and 120 mm in diameter showed good agreement over the range

of drop sizes investigated (circa 70 to 300 |im). Again, the maximum rotational speed of 9,500 rpm is low

in comparison to typical HSRBPA conditions.

In summary, few previous studies have addressed high-speed rotary-bell atomization. The authors are unaware of any published reports on drop size measurements for commercial HSRBPAs, despite their

widespread use. While some studies have dealt with rotary atomization in the ligament mode, the rotational

speeds employed were typically much less than 20,000 rpm. Consequently, the observed drop sizes were

large and the magnitudes of the centrifugal and Coriolis forces much smaller than in HSRBPA systems.

Finally, the drop-sizing techniques used in previous studies are less accurate than the laser-based diffraction

technique (Malvern particle sizer) used in this study.

The goal of the present research is to quantify the influence of the operational parameters (rotational

speed and liquid mass flow rate) and the key liquid properties (surface tension and viscosity) on the drop-size

distributions produced by a commercial HSRBPA. In order to isolate the hydrodynamic effects and develop

a baseline from which to assess their influence, neither shaping air nor applied voltage are used.

EXPERIMENTAL

The Turbobell III rotary paint applicator manufactured by Ransburg-Gema is depicted in Fig. 1. It uses an air turbine to spin a 70 mm diameter bell at speeds of up to 40,00 rpm. A digital controller is used to maintain any given preset bell speed. Fluid is delivered to the bell through a number of discrete circular

feed ports by a positive displacement pump. Once it emerges onto the the bell surface the fluid encounters

a series of bell steps which allow it to build up before flowing radially outwards across the smooth bell

surface. Along the edge of the bell are small triangular wedges called serrations. These serrations, spaced

approximately 0.3 mm apart, tend to force the fluid into discrete ligaments and extend the range over which

the atomizer can operate in the ligament mode.

This study combines qualitative visualization techniques with quantitative drop size measurements. A

0.5 \xs flash is used to "freeze" the motion of the bell. Both front and back-lit photographs are taken to visualize the flow on the bell surface as well as the ligament disintegration.

A Malvern 2600 particle sizer (MPS) with a 300 mm lens is used for quantitative measurements of

drop size distributions. The principles of operation and limitations of the MPS are well established; see for

example, Swithenbank et al. [9], and Dodge and Cerwin [10]. The drop size distributions are characterized

by the Sauter mean diameter (SMD) using the two parameter Rosin-Rammler [11] model. The RosinRammler model is defined as

where Q is the fraction of the total volume contained in drops of diameter less than D, and X and q are constants for a given drop size distribution. The exponent q in Eq. (1) provides a measure of the spread in

the drop sizes. The higher the value of q the narrower the size distribution.

As shown in Fig. 2, the axis of the atomizer lies in the horizontal plane. The 9 mm diameter laser

beam of the MPS passes through the spray at an angle of 30 degrees. This is done to obtain measurements

near the bell edge while providing sufficient space for the receiving optics. A light tube is used to prevent

laser light scattering from any other region except near the bell edge. The SMD was determined to be

invariant within plus or minus 10 mm about this 35 mm location. Also, at a distance of 35 mm the time

of droplet flight is so short that evaporation effects are negligibly small. For these reasons, all drop size

measurements are made 35 mm radially from the bell edge. Two thousand measurements of drop size

(1)

122
distributions are made at each test condition. As obscuration values are always less than 30% no

corrections for multiple scattering are required. The Rosin-Rammler [11] reduction procedure indicates log

differences between 3.0 and 4.5 which lies well below the stipulated maximum value of 5.0 for a good data

fit.

RESULTS

The effects of operational parameters and liquid properties were investigated over wide ranges, as shown

in Table 1. The bell speed and liquid mass flow rate (referred to henceforth as flow rate) were varied within

the allowable limits of the current installation. Surface tension and viscosity, the liquid properties having

the greatest influence on drop size, were varied by formulating mixtures of com syrup and water as well as mixtures of hydrocarbon oils. The resulting fluids represent a two order of magnitude change in viscosity

and a 250 percent change in surface tension. All liquids exhibited Newtonian behavior.

Order of magnitude changes in bell speed and flow rate produced all of the atomization modes identified

by Hinze and Milbom [5]. However, atomization by direct drop and sheet formation produced large drops

that are unsuitable for painting purposes. Hence, all the drop size measurements were made with the

atomizer operating in the ligament mode of droplet formation.

The effects of bell speed and liquid flow rate on Sauter mean diameter are illustrated in Fig. 3 for four

of the fluids listed in Table 1. The results show that, regardless of liquid properties, an increase in bell

speed at a given flow rate always tends to decrease the SMD towards a limiting size of between 15 and 20

um. At low bell speeds, the expected result of higher flow rates producing larger drop sizes is observed, but

at higher bell speeds the SMD becomes much less dependent on flow rate. The drop size data corresponding to a liquid flow rate of 3.3 g/s have been replotted in Figs. 4 and 5 to

illustrate the main effects of liquid properties on SMD. Qualitatively similar results were obtained at the

other flow rates examined. The influences of bell speed and surface tension on SMD for low-viscosity

liquids (circa 1 mNs/m^) are presented in Fig. 4. This figure shows that the higher surface tension fluid

produces larger drops over the entire range of bell speeds. The effects of viscosity on mean drop size for

fluids of low surface tension (circa 30 mN/m) are shown in Fig. 5. At low bell speeds, the expected result

of an increase in viscosity producing larger drops is obtained. Of special interest however, is that the

influence of viscosity on SMD diminishes with increase in bell speed and becomes negligibly small at the

highest bells speeds for which SMD data were taken.

The effects of bell speed and flow rate on the Rosin-Rammler distribution parameter, q, are shown in

Fig. 6. For water, the effect of increasing bell speed at constant flow rate produces no significant change in

q. For other liquids, having either significantly higher viscosity or much lower surface tension than water,

it is observed that increasing the bell speed diminishes q, thus indicating a broader range of drop sizes in the spray. For all four liquids, Fig. 6 shows that a reduction in flow rate produces a higher value of q, i.e. a more narrow drop size distribution.

It is often claimed that the sprays generated by rotary bell atomizers are characterized by an

exceptionally narrow drop size distribution. In this context it is of interest to note in Fig. 6 that when the

bell atomizer is operating at conditions of most practical interest [bell speeds of 19,000 rpm and higher],

the drop size distributions are just as broad as those produced by air or airless paint applicators [12,13].

DISCUSSION

Rotary-bell atomization in the ligament mode consists of four sequential processes: (1) filming on the

bell surface, (2) ligament formation at the bell edge, (3) attenuation of stable ligaments, and (4) ligament

disintegration into drops. Filming on the bell surface is strongly influenced by liquid properties and bell

speed. Water flowing at around 7 g/s films the bell as a complete wavy film at all bell speeds. At lower

flow rates, and all bell speeds, water flows over the bell surface in discrete wavy streams. An example of

these discrete streams is shown in Fig. 7a. The width of these streams is unaffected by changes in flow

rate between 1.17 and 3.3 g/s and bell speeds between 8,000 and 16,000 rpm. However, at bell speeds of

24,000 rpm and higher, stream widening is observed over a range of flow rates from 1.17 to 3.3 g/s.

Stream widening occurs because water, having low viscosity, is unable to resist the Coriolis force. The

other low viscosity fluid, fluid IV, exhibits slightly different filming characteristics. For bell speeds less

than 8,000 rpm this fluid traverses the bell as a complete wavy film. As the bell speed is increased from

8,000 to 40,000 rpm, the increased centrifugal force causes the fluid film to divide into discrete wavy

streams at all flow rates below about 7 g/s. At high rotational speeds, an increase in flow rate causes the

123
discrete streams to widen until, at flow rates higher than 7 g/s, the fluid flows over the bell in the form of a complete wavy film.

Separate streams are also observed near the fluid feed ports for high viscosity fluids. Figure 7b shows

how these streams merge into a smooth, uniform film shortly after flowing over the bell steps.

As a result of these filming characteristics, different types of ligament formation can occur, depending on the fluid viscosity. Low viscosity fluids reach the bell edge in the form of discrete streams, each of

which is converted into a cluster of ligaments. Fluids of low surface tension tend to produce wider streams

which result in a larger number of ligaments in each cluster. One such cluster for water is shown in Fig.

8a. The spacing of the ligaments in this high magnification photograph is approximately 0.3 mm. Water

forms very short, irregular ligaments that disintegrate into drops close to the bell edge. The ligaments

produced by fluid IV, which has a lower surface tension than water, tend to be similar in structure to those

of water except that the ligaments are generally longer. Figure 8a shows that the breakup mechanism for

these ligaments is either Rayleigh or first wind-induced [14], since the resulting drops are either equal or

larger in diameter to the ligaments from which they are formed.

High viscosity fluids film the bell completely. Two types of ligament formation are observed,

depending on the operational parameters. At bell speeds lower than 8,000 rpm, the fluid film is so thick at the bell edge that the serrations are incapable of dividing the film into one ligament per serration. Instead, a

liquid torus forms on the bell edge which eventually becomes unstable and produces ligaments which are spaced at an integral number of serrations. At bell speeds above 16,000 rpm, the fluid at the bell edge is thin enough to divide itself into a single ligament per serration. These ligaments are illustrated in Fig. 8b

where the spacing between adjacent ligaments is again about 0.3 mm. These viscous ligaments tend to be

evenly spaced and regular in shape. They also remain attached to the rotating bell for about four times the

distance of their low viscosity counterparts. This results in a considerable reduction in ligament diameter

along the streamwise direction due to the effect of centrifugal force in increasing the streamwise velocity.

The relative velocity between the ligaments and the surrounding air increases with increasing bell speed.

This causes the ligaments to become wavy and follow a sinusoidal trajectory as they move further away

from the bell edge.

The preceding discussion of bell filming and ligament formation helps to explain the dependence of

SMD on flow rate exhibited in Fig. 3. For example, in Fig. 3a the change in dependence of SMD on flow

rate at bell speeds above and below 10,000 rpm can be explained by differences in the filming

characteristics. Below 10,000 rpm, fluid IV films the bell completely, so the effect of increasing the flow

rate is to increase the film thickness at the bell edge and hence the initial ligament diameter. Since these

ligaments are fairly short, and have no chance to attenuate, this increase in initial ligament diameter will cause an increase in drop size. Above 10,000 rpm, high centrifugal forces cause stream widening to occur

as the flow rate is increased. This widening of the stream tends to offset the effect of the increase in flow

rate, so the film thickness at the bell edge remains about the same regardless of flow rate. This behavior

explains the insensitivity of mean drop size to variations in flow rate at bell speeds above 10,000 rpm.

Figure 3b illustrates the effect of bell speed and flow rate on SMD for water. It shows trends similar

to those in Fig. 3a, except that the bell speed at which the mean drop size becomes insensitive to flow rate

is around 20,000 rpm. This critical bell speed marks a change in the effect of flow rate on stream widening.

At bell speeds below the critical value, the stream widths remain unaffected by changes in flow rate, but at higher bell speeds an increase in flow rate causes the streams to widen. Since the ligaments for water are even shorter than those of fluid IV, due to its higher surface tension, it is to be expected that increases in

flow rate below 19,000 rpm should increase drop sizes, while increases in flow rate above 19,000 rpm

should have little effect on drop size. These expectations are confirmed in Fig. 3b.

Figures 3c and 3d show that the high viscosity fluids have two regions of dependency of flow rate on SMD. Below 29,000 rpm the SMD is very dependent on flow rate, while at higher bell speeds there is

little effect of flow rate on SMD. Both liquids totally film the bell surface, so an increase in flow rate always produces a thicker film at the bell edge and hence a larger initial ligament diameter. Thus, the only

way the SMD can be insensitive to flow rate above 29,000 rpm is if the final breakup diameter of the

ligaments is independent of its initial value at the bell edge. This does occur with viscous ligaments

because they are stable and can extend radially outward to a distance which is several times greater than for

low viscosity liquids. Hence, at high bell speeds where centrifugal forces are large, a viscous ligament can

thin out to a diameter which is largely independent of its initial diameter. This attenuation mechanism

appears to be present at bell speeds higher than 29,000 rpm, and is mainly responsible for the independence

of SMD on flow rate. Below 29,000 rpm, ligament attenuation is much less pronounced and an increase in flow rate raises the SMD.

124
The fluid property effects examined in this work are those of surface tension and viscosity. As shown

in Fig. 4, an increase in surface tension tends to produce larger mean drops sizes for liquids of low

viscosity. This is attributed to the effect of surface tension on stream width. An increase in surface tension

reduces the spreading capability of individual streams. This results in thicker streams that produce larger

initial ligament diameters and ultimately larger drops. Figure 5 shows the effect of viscosity at low surface

tension. Below 20,000 rpm, the expected result of a higher viscosity fluid producing larger drops is observed. Above 20,000 rpm, the mean drop size remains essentially the same for a two order of magnitude

increase in viscosity. For this to occur there must be a competing mechanism at high bell speeds to offset

the well-known effects of viscosity on ligament disintegration. The photographic evidence suggests that

the competing mechanism is the attenuation of a viscous ligament under the action of high centrifugal

forces. Thus although viscosity has its normal adverse effect of increasing the optimum wavelength for

ligament breakup, this effect is largely offset by the fact that when breakup finally occurs it does so at a

smaller ligament diameter.

The other important spray characteristic along with SMD is the drop size distribution. Figure 6 shows

how the Rosin-Rammler distribution parameter q varies with flow rate and bell speed. A general trend

exhibited by all four fluids in Fig. 6 is that as the flow rate is increased the drop size distribution widens

(i.e. q decreases), regardless of the bell speed. At high bell speeds, where SMD remains constant, the fact

that q decreases with increasing mass flow rate should be interpreted as an indication that more fine droplets

are being produced. Another general trend exhibited by all the fluids except water is that an increase in bell

speed causes the drop size distribution to broaden. This can be accounted for by considering the effects of

aerodynamic forces on the ligaments. Increasing aerodynamic forces cause the ligaments to wave

sinusoidally and interact with each another. This tends to make the atomization process more random and

thereby lower q. Photographs indicate that this ligament interaction is most pronounced at high bell speeds

because the ligaments are located in close proximity at the serration spacing of 0.3 mm and the amplitude

of their wave behavior is larger at higher bell speeds. At lower bell speeds, the ligaments are spaced further

apart which lowers the chances of ligament interactions and the amplitude of displacement is smaller. The

only fluid in Fig. 6 that shows a slighdy different trend from the other fluids is water. For water, q changes

only slightly as the bell speed is varied at a constant flow rate. High speed photographs show that water

produces very short ligaments which do not extend far from the bell edge and thus are less susceptible to the

aerodynamic forces that change appreciably with bell speed. Consequently, ligament interaction is much

less pronounced, so that q remains fairly constant as the bell speed increases.

SUMMARY AND CONCLUSIONS

In interpreting the effects of operational parameters and liquid properties on the drop sizes produced by

rotary bell applicators operating in the ligament mode of atomization, four separate processes must be

considered: (1) filming on the bell surface, (2) ligament formation at the bell edge, (3) attenuation of stable

ligaments, and (4) ligament disintegration into drops. An increase in rotational speed always produces a decrease in mean drop size by increasing the centrifugal and Coriolis forces. These forces promote

increased thinning of the liquid film on the bell surface, with the result that the ligaments formed at the

edge of the bell are smaller in diameter.

An increase in liquid flow rate generally increases the mean drop size at the lower bell speeds. At

higher bell speeds, increasing die flow rate has litUe effect. This insensitivity to flow rate of low viscosity

fluids is attributed to stream-widening, while for high viscosity fluids it is attributed to ligament

attenuadon.

Increasing the flow rate at a given bell speed always leads to a broader distribution of drop sizes in the

spray. Increasing the bell speed at constant flow rate generally widens the drop size distribution due to the

effects of aerodynamic forces on ligament trajectories. In flow situations where the bell surface is incompletely filmed, which is generally the case for low

viscosity fluids, an increase in surface tension raises the mean drop size by narrowing the discrete fluid

streams flowing over the bell surface, thereby increasing their thickness. Viscosity has an adverse effect on mean drop size for low surface tension fluids at low bell speeds, but has little effect at the higher speeds.

This is attributed to the thinning mechanisms that operate at high bell speeds.

REFERENCES

1. S. Malsumoto, D.W. Belcher, and E.J. Crosby, Rotary Atomizers: Performance Understanding and

Prediction, Proceedings of ICLASS-85. Madison, WI, pp. 1 A/1/1 -21, (1985).

125
!

!

1

2. K. Masters, Sprav Drying. John WUey & Sons, New York, NY, (1979).

3. N. Dombrowski, and G. Munday, Spray Drying. In Biochemical and Biological Engineering Science

(N. Blakebrough, ed.). Academic Press, New York, NY, vol. 2, pp. 209-320, (1968).

4. W.R. Marshal, Atomization a Spray Drying, Chem. Eng. Monograph Series no. 2, vol. 50,

A.I.Ch.E., (1954).

5. J.O. Hinze, and H. Milbom, Atomization of Liquids by Means of a Rotating Cup, J. Appl. Mech.. vol. 17, pp. 145-153, (1950).

6. N. Dombrowski, and T.L. Lloyd, Atomization of Liquids by Spinning Cups, Chemical Engineering

Journal, vol. 8, pp. 63-81, (1974).

7. A. Kayano, and T. Kamiya, Calculation of the Mean Size of the Droplets Purged from the Rotating

Disk, Proceedings of the First International Conference on Liquid Atomization and Spray Systems.

Tokyo, pp. 133-137, (1978).

8. Frost, A.R., Rotary Atomization in the Ligament Formation Mode, Journal of Agricultural

Engineering Research, vol. 26, pp. 63-78, (1981).

9. J. Swithenbank, J.M. Beer, D. Abbot,, and G.C. McCreath, A Laser Diagnostic Technique for the

Measurement of Droplet and Particle Size Distribution, Paper 76-69 presented at the AIAA 14th

Aerospace Sciences Meeting, Washington D.C., January 26-28, (1976).

10. L.G. Dodge, and S.A. Cerwin, Extending the Applicability of Diffraction-Based Drop Sizing

Instruments. In Liquid Particle Size Measurement Techniques. ASTM STP 848 (J.M. Tishkoff, R.D.

Ingebo and J.B. Kenndey, eds.), American Society of Testing and Materials, pp. 72-81, (1984).

1 1 . P.Rosin and E. Rammler, The Laws Governing the Fineness of Powdered Coal, J. Inst. Fuel , vol. 7, no. 31, pp. 29-36, (1933).

12. H.E. Snyder, Drop Size Investigation of an Electrostatically-Assisted Fan-Spray Atomizer. M.S.M.E.

diesis, School of Mechanical Engineering, Purdue University, (May 1988).

13. N.H. Goris, Operational Characteristics and Energy Considerations of Pneumatic Atomizers.

M.S.M.E. thesis. School of Mechanical Engineering, Purdue University, (May 1990).

14. R.D. Reitz, Atomization and Other Breakup Regimes of a Liquid Jet, Ph.D. thesis, Princeton

University, (1978).

15. C. Weber, Disintegration of Liquid Jets, Z. Angew. Math. Mech. . vol. 11, no. 2, pp. 136-159, (1931).

TABLE 1 Operational Parameters and Liquid Properties

Operational Parameters:

Rotational speed, rpm 2,000 - 40,000

Mass flow rate, g/s 0.67 -6.7

Liquid properties (295 K):

Surface Tension Viscosity

Fluid (mN/m) (mNs/m-^)

I. water 76 1.0

II. glycerin/water 65 100

III. com syrup/water 80 100

IV. hydrocarbon oil 32 3.0

V. hydrocarbon oil 30 100

126
80

40

120f

80

40

I oj

Q 120

80

40

8

a

oo

°4r

80

40

Q

U

A 067

167

O 3.33 O 5.00 ^ 6 67

S g FLUID 12 (a)

B g WATER (b)

Q

FLUID Y (c)

a

FLUID m (d)

0 10 20 30 40 50 X I0' BELL SPEED, rpm

Fig. 3 Influence of liquid flow rate on

mean drop size.

Fig. 2 Optical configuration for drop size measurements.

120

100

80

E

^-60

40

00

Fig. 4

rhj^ = 3.33 g/s

o

a, mN/m

O 76 (1) 32 (IV)

o

o

o

O

1 1 1 1 1

1 1 1 1 10 20 30

BELL SPEED, rpm

40 50X10^

Influence of surface tension on SMD for liquids of low

viscosity.

127
if)

120

100

80

60

40

20

Q

rrij^ = 3.33 g/s

o

1 i m W c / rr» ^ o ion (v)

3.0 (IV) - o

1 1

1 1 1 1 1 1 1 10 20 30

BELL SPEED, rpm

40 50

Fig. 5 Influence of viscosity on SMD for liquids of low surface

tension.

Ml = 10 mNs/ni2. o = 76 mN/m jWater)

(a)

HL = '00 mN,s/m2. o = 65 inN/m (Fluid ll| (b)

Fig. 7 Effect of viscosity on bell filming characteristics

at a bell speed of 24,000 rpm. nit = 3.3 g/s.

4

3

2

I 0

ui 5 ^ /I <

2

^

A

8 §1 o

3

CD

COQ

UJ5

<

or

I z

(/) o

A

D

^

A

D ^

WATER

_m| , g/s

A 0.67

167

O 3.33

O 5.00 ^ 6.67

FLUID m

0-

4'

3

2

I -

0

i4-

3-

2-

I -

o

o

A

A

FLUID m

Q

A

8

A

D

A

8 FLUID ¥

0 10 20 30 40 50 xlO'

BELL SPEED, rpm

Fig. 6 Influence of bell speed and flow rate on

drop-size distribution.

I'l. = I.OiiiNs/ii|2, o = 76niN/in |Waicr|

(^)

Ml. = ino niNs/in2, o = 80 mN/iti [t-luid lin

(h)

Fig. 8 Effect of viscosity on ligament structure and breakup at bell speed of 16,000 rpm.

128
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 9

MEASUREMENTS IN PAINT SPRAYS USING

A PHASE-DOPPLER ANEMOMETER

J. Domnick, C. Tropea and T.-H. Xu

Lehrstuhl fiir Stromungsmechanik

Universitat Erlangen - Numberg

Erlangen, Germany

ABSTRACT

Simultaneous measurements of droplet size and velocity in paint sprays using a phcise-Doppler anemometer are presented. Mie calculations of the scattered light field, based on measured refractive indices and

absorption coefficients of 2 diff'erent paints, have shown the possibility of obtaining good results with 1st

order refracted light. Furthermore, verification experiments with a monodispersed droplet generator and a

small hand-held spray gun were in good agreement with the calculated transfer function of the phase-Doppler

system. Extended measurements in a spray cabin were performed, changing a variety of parameters like gun

type, air cap and paint and air flow rates. Profiles of velocity and size were taken at 3 diflferent distances,

5, 15 and 30 cm downstream from the nozzle. The results have shown some of the expected dependencies

of the local velocity and droplet distributions on global parameters, e.g. an increase of the mean sizes

with decreasing air flow. The measured size distributions were fitted to a modified log-hyperbolic function,

thereby enabling the entire size distribution to be represented in terms of three parameters. The variations

in these parameters with changing paint flow rate are documented and are shown to adequately represent

the physical background of the phenomena under consideration.

1 INTRODUCTION

In industrial painting in Europe pneumatic spraying systems of the air-assist type with external mixing

cover most of the applications, including robot driven systems in the car manufacturing. A standard spray

gun, shown schematically in fig 1 has several holes arranged radially around the paint stream in the center,

providing high air velocity for the atomization.

The air velocity is usually close or equal to the velocity of sound at the outlet. This high air flow leads on the one hand to a good atomization and to size distributions necessary for a good quality of the paint film on the object and on the other hand it results in a loss of paint due to droplets following the air flow parallel

to the object's surface. This is the so-called overspray. For reasons of pollution control and reduction of

secondary costs for disposal of overspray strong eff'orts are undertaken to reduce the amount of overspray.

Apart from new developments in painting like ultrcisonic or rotary atomizers with or without electrostatic charge it is also necessary to improve the behavour and the efficiency of the usual pneumatic atomizer,

e.g to spray water-based paint. Here, a close look to the local size distributions and the velocity field between the spray gun and the surface to be painted is a keypoint for a better understanding of the recisons

for overspray production.

Until now, only few measurements in paints sprays have been published. Some measurements of the

velocity field have been made using Pitot-tubes [1] or Laser- Doppler anemometer [2]. Most of the droplet

size measurements reported were using Fraunhofer-Difi'ration, e.g. [3], or sampling techniques [4], [5]. None

of these papers deals with simultaneous measurements of drop size and velocity.

129
guide air

Figure 1: Pneumatic paint atomizer

This paper presents recent results of phase-Doppler measurements in paint sprays. In principle, this

technique has been proved to give very good and reliable results in any kind of spray flows, e.g. [6], [7], [8]. In the derivation of the principle of the phase-Doppler technique, it is assumed that the dispersed phase is a simple fluid with very well known optical properties like refractive index and absorption coefficient. Paint,

however, is a mixture of liquid and solid with a rather complex nature, so some basic experiments have to be made to check the accuracy of the phase-Doppler technique applied to paint drops, including calculations

of the scattered light field by means of a Mie-program.

2 OPTICAL CONSIDERATIONS

The phase-Doppler technique is suitable for any kind of spherical particles, including droplets in gas

or bubbles in liquid. Depending on the relative refractive index between dispersed and continuous phase,

measurements are possible if one of the scattering mechanisms reflection, 1st or 2nd order refraction is dominant. For this purpose programs based on Mie scattering theory [9] or geometrical optics have been

developed to enable the PDA user to adapt his system to the flow situation.

All of these programs deal only with homogeneous fluids, which have constant refractive index and

absorption inside the droplet. The application to paint, which is a mixture of solid and liquid parts, including

pigments, metallic powder, solvents and water, is therefore limited. Nevertheless it should be possible to get some first ideas about the scattered light field out of these programs. The following table shows the

measured optical properties of the examined paints; water-based silver metallic and a lacquer.

refractive index absorption coeff. lacquer

silver metallic

1.42

1.36

8.63- 10-«

5.98-10-''

In order to be able to perform Mie calculations it is also necessary to define the optical arrangement

of the complete PDA-system. A one-component HeNe system with rotating grating and 500 mm lenses on

both the sending and the receiving side was used. A 3x beam expansion was used to achieve a high intensity

in the probe volume. The specifications were as follows

probe volume diameter (/im): 134

fringe number (-): 6

fringe spacing (/im): 21.07

In the case of silver metallic droplets the maximum measurable size, using a standard two-detector

receiving optics, was 360 fim at 30** scattering angle.

130
40

20

4)

0)O

G

n

5 -20

^ -40

em -60

-80,

I I I I I I I I I I

I

I I I I I

I

I I I I I

I I I I I I

I

I I I I I -

Reflection

Porticle diameter

IS/im 50fim

Wavelength = 632.8 (nm)

Refraction Half-angle of beams = .8594 (deg) Refractive index = 1.36 Absorption coef. = 5.985E-4

Receiving-cone lialf-angle = 1.72 (deg)

Receiver elevation angle = 2.29 (deg) Polarization angle = 0 (deg)

I I I I 1 I I I I I I I I I 30 60 90 120

Off-axis angle (deg)

-I 150 180

Figure 2: Phase as a function of scattering angle (silver metallic)

W)V

DO

C

0)

41

CO

a,

-60

-120^

Wavelength = 632.8(nm)

Half-angle of beams = .8594 (deg)

Off-axis angle = 30 (dee)

Receiving-cone half-angle = 1.72 (deg)

Receiver elevation angle = 2.29 (deg)

clear— lacquer, refractive index = 1.42, absorption coef. = 8.696E—

silver-lacquer, refractive index=1.36. absorption coef.=5.985E- F I I I I I It I

I I I I 50 100 150

Particle diameter (/um)

200

Figure 3: Phase as a function of droplet size (silver metallic and lacquer)

Figure 2 indicates the phase as a function of the scattering angle for three different droplet sizes, 5, 15

and 50 /im. For clarity only the result for the silver metallic is shown in this figure. This figure also indicates a wide range between 25 and 80" scattering angle where 1st order refraction is dominant, characterized by an

increase of the phase in the negative direction. This result is not surprising, since the absorption coefficient

lies in the range of 6 • 10"'*, leading to only 5 % absorption of the refracted light in a 50 /im particle. But

one has to keep in mind, that the effect of the absorption will increase at larger droplets. This is also visible

in Figure 3, showing the response curve of the pheise-Doppler system at a scattering angle of 30". From 100

fim onwards, the curve for the silver metallic starts to deviate from a straight line, which is the result of the

beginning influence of scattered light from reflection.

From these curves one can expect a very linear relation between droplet size and phase for both silver

metaflic and lacquer. As will be shown later most of the droplets can be obtained in the range under 60

fxm, so the influence of reflection can be neglected. These results are valid only for homogeneous fluids. The

effect of the inhomogeneities inside the droplets is still to be investigated.

3 PRELIMINARY TESTS

Preliminary tests were performed to check the accuracy of the phase- Doppler technique applied to paint

sprays. A vibrating orifice monodispersed droplet generator was used to produce a stream of paint droplets

of known size. Fig. 4 shows PDA-measurements of 125 fim lacquer droplets.

Apart from the error in mean size, which is only 2.5 %, there is a small broadening of the size distribution

of about ±15//m. It is possible that this was due to the droplet generator performance since it is not designed

131
r, J T- / 1 ->c PDA ANALYSIS Elapsed Tune (s): 26. samples validated: 625 samples rejected: 4 Mean Diameters (microns)

snr: 0 DIO : 1 23 frequency: 0 D20 : 128 amplitude: i D30 '29 burst length; 0 Sauter Mean : 129 burst rate (Hz): 8938.

Particle Size Wis. Velocity Correlation

0,0 60.6 121-2 181.8 2-*2.i diameter (um)

503.

Figure 4: Measured size distribution of a monodispersed droplet generator (lacquer)

5.,. , -. , , ,c P3A ANALYSIS c-Iapsed Tune (s): 26. samples Talidated: 8514 samples rejected: 671

snr: 69 frequency: 9 amplitude: 486

burst length: 107 burst rate (Hz): 924.45

Mean Diameters (microns) DIO

D20

D30

1 1 9

15 2

19.3 Sauter Mean : 34.2

Particle Size vs. Velocity Correlation

0.0 90 2 135.3 diameter (um)

:8C.

Figure 5: Local size distribution at z = 25 mm (lacquer, low flowrate)

to produce droplets of a high viscous and nonhomogeneous fluid. In addition, Doppler amplitude variations

visible on the screen of a digital oscilloscope indicated also size variations inside the droplet stream.

Another method to clarify the response of the phase-Doppler in paint sprays is the controlled variation

of one dominant operating parameter of a spray gun for which the influence on mean size and velocity is known. The phase-Doppler technique should be able to demonstrate the corresponding change in the droplet

size distribution and the mean values. For instance Figures 5 and 6 show the influence of a change in paint

flow rate on the local size distribution at 25 mm distance from the nozzle. The measurements were made

with a simple hand-held spray gun.

There is a significant increase in the mean droplet sizes with increasing paint flow rate; the arithmetic

mean Dio is more than doubled. The mean velocity has decreased as can be explained by the higher

momentum exchange from the fast air flow to the paint stream. In both cases there exists a correlation

between the droplet size and the mean velocity.

Both experiments have been performed with a 30" off'-axis arrangement of the PDA-receiving optics,

using the normal phase factor based on geometrical optics calculations for refraction, as indicated in the

previous section. Signal processing was done with a transient recorder system using the cross-spectral density

function [10] for frequency and phase estimation. A probe volume correction was made to account for the

Gaussian intensity profile in the probe volume [11].

132
Elapsed Time (s): 20. samples validated: 4621 samples rejected: 751

snr: 1 58

frequency; 10 amplitude: 476

burst length: 1 17 burst rate (Hz): 1 109 5

PDA ANALYSIS

Mean Diameters (microns) DIO : 29.1 D20 34. 4 D30 ; 39.3 Sauter Mean 53.3

Particle Size vs. Velocity Correlation

0.0 45.1 90-2 135.3 180.4 225.4 , diameter (um) j

Figure 6: Local size distribution at z = 25 mm (lacquer, high flowrate)

Elapsed Time (s): 32. samples validated: 3890 samples rejected: 1482

snr: 172

frequency: 19

ajgipll^jde: '07

b«irst ^ngth; 1134

PDA ANALYSIS

Mean Diameters (microns) DIO 1"3

D20 24 2 D30 32 6 Sauter Mean . 59 2 bSrst^ate (Hz): ' 06381*^005

Particle Size vs. Velocity Correlation

+0.0 30.0 ;:3.o diameter (um)

:5C 0

Figure 7: Typical size and velocity distribution

4 MEASUREMENTS IN A SPRAY CABIN

After the very encouraging results of the preliminary test measurements further experiments were made

in a more practical situation inside a spray cabin, using two different spray guns designed for robot systems.

The purpose of the study was to verify existing differences between different spray guns and air caps and to demonstrate the influence of the major operating parameters, e.g. liquid and air flow rates. Profiles of local

size and velocity distributions were taken at 5, 15 and 30 cm distances from the nozzle, using silver metallic

paint throughout the experiments. Care was taken to keep all the parameters in their normal operating

range. The paint flow rate was adjusted by a speed controlled gear pump, the atomizing and guide air flow

rates were measured using an instrument based on two temperature compensated hot film anemometers.

The spray gun was mounted on a x-z traversing system, allowing vertical and horizontal movements relative

to the fixed table of the PDA-system. The flow direction was vertical downwards.

In the beginning, all spray guns were tested with water as an atomized fluid, yielding a very narrow

spray cone with small and fast droplets. The measured arithmetic mean diameters with water were in the

range of 7.0 - 10.0 //m, the corresponding velocities up to 80 m/s at a distance of 5 cm from the nozzle. In contrast, spraying paint led to a wider spray cone with reduced velocities. Fig. 7 shows a typical plot of the

size and velocity distribution from 3000 samples at one single point, indicating a Djo of 17.3 um, D^q of

32.6 //m. The velocity is around 25 m/s.

It was a common feature of the measured size distributions that they display a peak in the first size

133
Figure 8: Axial development of the velocity (spray gun 2)

70.0

radial distance (cm)

I I

I 17.5I

< I 20.0 22.5

Figure 9: Axial development of mean diameters (spray gun 2)

class, which extends up to 5 /xm. On the other hand, big droplets could be obtained up to 150 /xm and

higher, resulting in strongly skewed distributions. Since the higher moments like D30 or the Sauter diameter

are strongly influenced by the few very big droplets, 3000 samples was found to be a good compromise

between stable mean diameters and a limited measuring time.

In Figs. 8 and 9, the axial development of size and velocity of spray gun 2 is shown. The mean velocity

profile indicates a steep gradient at z = 5.0 cm, from 60 m/s in the core to 6 m/s at the edge of the spray

cone, flattening out to a nearly constant velocity of 10 m/s at z = 30.0 cm. The diameter profiles Dio and

D30 which are shown are very homogeneous except for the very edge of the spray cone where a sharp increase

of the mean diameters can be noticed. This is common for all three spray guns examined.

This result can be explained by the presence of a shield of big droplets surrounding the spray cone. The cone edge itself was defined using the arrival rate of droplets in the probe volume.

The measurements of spray gun 2 are also taken to explain the influence of a change in paint flow on

the size distribution in the spray cone. Two flow rates, 250 ccm/min and 450 ccm/min, were examined. The

result for a distance z = 30 cm is shown in Fig. 10. It is very interesting to note that the additional amount

of paint is concentrated in larger droplets in the outer region of the spray cone, expressed in an increase

of the volume mean diameter in the region x > 5.0 cm, whereas Dio is changing only slightly. In general,

higher paint flow rates lead to a more inhomogeneous distribution of the mean sizes of the spray. From this

point of view it would be preferable to operate the spray guns at a low paint flow rate, though this would

134
50.0

40.0

30.0

(U

-i-> (U^20.0

10.0

Gun 2. z=30.0 cm, ZL/HL 250/350

,H +

- fr- _ - *

*

00000 0,0 250 ccm/min AAAAA 0,0 450 ccm/min

iW*ft* D» 250 ccm/min

1 I I I 1 Djo 450 ccm/min

0.0 ~1—I—I—I—I—

I—1—I I I—I I I I—I—

I—n—I—I—

I—I—n—I—r"i—[ 1 I—

I—I—I—I—I I

I—r-i—i—[—1—1—r- 0.0 2.5 5.0 7.5 10.0 12.5 15.0 17.5 20.0 22.5 radial distance (cm)

Figure 10: Influence of paint flow rate on mean diameters (spray gun 2)

20.0

17.5

15.0 -.

m^12.5

t>,10.0

5.0

2.5

0.0

0.0

Gun 1, z=30.0 cm, 250 ml/min

oeeeo ZL/HL 200/200

O^MO ZL/HL 280/380

—I

5.0 radial distance (cm)

Figure 11: Influence of air flow rate on velocity (spray gun 1)

imply a lower travelling speed during painting.

Finally, the influence of the air flow rates on size and velocity should be discussed here. As indicated

already in Fig. 1, there are 2 independent air supplies, the atomizing air flowing through 4 or 6 nozzles

very close to the paint outlet and the guide air, flowing from 2 sides through radially arranged nozzles,

thus leading to a elliptically shaped spray cone. The appropriate choice of the air flow rates is essential for

the painting process since both the atomization of the paint stream and the air flow field in the vicinity of

the surface is strongly influenced. Higher air flow will lead to better atomization resulting in smaller mean

droplet diameters but also to higher air velocities parallel to the surface and therefore to more overspray

caused by small droplets following the air flow field.

Figs. 11 and 12 indicate the results obtained for 2 difi'erent air flow rates of spray gun 1 at 30 cm

distance from the nozzle. From Fig. 11 it can be seen that the velocity profile remains unchanged but on

a lower level. It is also clear, that the cone angle, expressed in terms of the mean velocity, has decreased,

since at a location x = 25.0 cm the velocity is nearly zero. This can be explained by the smaller guide air

flow rate. The interpretation of the behavour of the mean diameters becomes more complicated. Fig. 12 indicates, as expected, higher Dao-values for the low-flow case in the range between 0 and 10 cm radial distance, but

also higher Dio-values for the high-flow case between 12.5 and 25.0 cm. The development of D30 in the

outer region is, however, nonuniform. One possible explanation could be the combination of the creation

135
50.0

40.0

'30.0 u

20.0

10.0

0.0 25.0 radial distance (cm)

Figure 12: Influence of air flow rate on mean diameters (spray gun 1)

of smaller droplets in the high-flow case due to the higher air flow with the tendency of bigger droplets to concentrate more in the outer region of the spray cone with increased air flow rate, thus covering the effect

of better atomization. It should be emphasized that the increase of the atomizing air flow is only 40 %,

whereas the guide air is nearly doubled.

This is one of the points where more detailed measurements, e.g. with a two-component PDA-system,

are necessary to clarify the correlations between different parameters in more detail.

5 DATA REDUCTION

Up to here, all effects have been discussed in terms of the mean values of velocity and size. Two different

mean diameters, Dio and D30, have been used to show changes in the shape of the size distribution. However,

if one is interested in any other moment, e.g. Sauter mean diameter, it is necessary to store and replay or even regroup the entire droplet size distribution. Thus, it would be convenient to define a analytical function

representing the measured size distribution very accurately without any loss of information. From this

function it is possible to calculate any desired number necessary for further applications of the measurements,

i.e. calculation of mass transfer.

The function proposed here is the so called three-parameter log-hyperbolic distribution, a modification

of the usual four-parameter log-hyperbolic function. As shown in [12], this distribution gives very good

results in many different kinds of size distribution, including this paint measurements. As illustrated in

Fig. 13, the three parameters a, 6 and fi have geometrical meanings, 0 being the angle of the hyperbola

axis relative to the coordinate system, // the location parameter defining the peak of the distribution, and a

defining the slope of the asymptotes or the opening of the hyperbola.

From the values of these three parameters the original size distribution can be reconstructed, which is not possible from the mean diameters. For example, different shapes of distributions can give identical Dio

and D30 values. In addition, it is possible to deduce changes in the shape of the distribution from changes

in the parameters, making physical processes clearer.

This function was applied to the measured size distributions in the case of paint flow rate changes

(spray gun 2), which was already discussed in section 4. Figs. 14 and 15 show the corresponding change

of the three parameters of the log- hyperbolic distribution. Parameter a, characterizing the opening of the

hyperbola, is constant, whereas a slight increase of parameter fi, characterizing the location of the mode

point, corresponds to the increased arithmetic mean diameter. Most of the effect of the flow rate change seems to come up in changes of 9. Therefore it can be concluded, that the additional amount of paint could mainly be found in big

droplets, concentrated in the outer region of the spray cone. This is expressed by the decrease of 9, denoting

a longer right-hand side tail and a higher skewness to the left. Furthermore, higher ^-values near the center

136
Figure 13: The three parameter log-hyperbolic distribution

Gun 2. z=30.0 cm, ZL/HL 250/350

3.0 -it:

QQQQDa, 250 ecm/min

I I I I I 0, 450 ccm/mln

AAAAft n, 250 ccm/mln

>uuuu$/j, 450 ccm/mln

I I I

I I I I II

I I I I

I 1 I I II

I I I I

I

I I I I

I I I I I

I I I I I

I I 0.0 2.5 5.0 7.5 10.0 12.5 15.0 17.5 20.0 radial distance (cm)

22.5

Figure 14: Variation of the parameters a and /i

Figure 15: Variation of the parameter 9

137
of the spray for the high flow case indicate also a lack of big droplets in that region, resulting in a very

inhomogeneous development of the mean sizes from the center to the edge of the spray cone. Based on

this result, neglecting other operation boundary conditions, it is not recommendable to use the investigated

spray guns at higher paint flow rates.

6 CONCLUSIONS

Although the data base of the presented PDA-measurements in paint is still limited, the obtained re- sults indicate the possibility to use the phase-Doppler technique in paint sprays. This could be shown with

preliminary tests, including measurements in a monodispersed droplet generator, as well as with measure- ments in a very realistic spray inside a cabin. Using a silver metallic paint, some basic correlations between

the operating parameters, e.g. paint and air flow rate, and the local size and velocity distributions were demonstrated.

The obtained size distribution could be fitted by the so-called three-parameter log-hyperbolic distribution resulting in a storage of the local information in 3 parameters. From the changes in the parameter

values changes of the spray flow could be made much clearer.

The final aim of the presented investigations is to improve the knowledge of the flow from pneumatic

paint atomizers in order to be able to reduce the amount of paint overspray and energy consumption without

deteriorating the quality of the paint film. Therefore further detailed measurements in paint sprays will be

made, including also the flow field in the vicinity of a surface. It is also planned to enlarge the number of

verified paints suitable for PDA-measurements.

References

[1] K. W. Warnecke Air velocity measurements at the jet of a paint spray gun, Industrie-Anzeiger, 100(7),

1978

[2] D. Miiller "Untersuchung des Zerstauhungsvorganges bei der pneuwatischen Lackapplikation" Diplomarbeit, Lehrstuhl fur Stromungsmechanik, Universitat Erlangen-Niirnberg, 1990

[3] H.E. Snyder; D.W. Senser; A.H. Lefebvre; R.S. Coutinho "Drop size measurements in electrostatic paint

sprays", IEEE Trans. Ind. AppL, 24(4), 1989

[4] W.S. Janna; J.E.A. John "Drop-size distributions of Newtonian and Bingham liquid sprays produced

by fan-jet pressure nozzles with and without preorifice". Trans. ASME, J. Engng. Ind., 103(40), 1981

[5] A. Biirkholz; J. Widder; W. Miiller "Primare Tropfengrofienverteilung und atembarer Feinanteil beim

Lackspritzen durch Druckluftzerstaubung" , Farben und Lacke, 82(8), 1976

[6] K. Bauckhage, H.-H. Flogel "Simultaneous Measurement of Droplet Size and Velocity in Spray Nozzles"

Proc. of the 2nd Int. Symposium on Appl. ofLDA to Fluid-Mechanics, Lisbon 1984

[7] W.D. Bachalo; M.J. Houser "Phase/Doppler spray analyzer for simultaneous measurements of drop size

and velocity distributions", Optical Engineering, 23(5), 1984

[8] W.D. Bachalo; M.J. Houser "Spray drop size and velocity measurements using the phase/Doppler

particle analyzer" Int Conference on Liquid Atomization and Spray Systems ICLASS 85, 1985

[9] A. Naqwi; F. Durst "Computation of light scattering from a dual-beam system". Internal report

LSTM/259/T/89, Lehrstuhl fiir Stromungsmechanik, Universitat Erlangen-Nurnberg, 1989

[10] J. Domnick; H. Ertel; C. Tropea "Processing Phase/Doppler Signals Using the Cross-Spectral Density

Function" Proc. of the 4th Int. Symposium on Appl. of LDA to Fluid-Mechanics, Lisbon 1988

[11] M. Saff"man "Automatic calibration of LDA measurement volume size", Applied Optics, 26(13), 1988

[12] T.H. Xu; C. Tropea; F. Durst "The three-parameter log-hyperbolic distribution and its application to liquid sprays", submitted to Int Conference on Liquid Atomization and Spray Systems ICLASS 91, 1991

138
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 10

ATOMIZATION, CHARGE AND DEPOSITION CHARACTERISTICS OF DUAL

POLARITY ELECTROSTATICALLY CHARGED AIRCRAFT SPRAYS OF

SERRATED SPINNER CUP ATOMIZERS

K.D. Kihm, B.H. Kim and A.R. McFarland

Department of Mechanical Engineering

Texas A&M University

College Station, TX, U.S.A.

ABSTRACT

Atomization, charge and deposition characteristics of an electrostatic aircraft spray system were studied. The

system used serrated spinner cup atomizers with tap water as atomized fluid. Despite the well-known advantages of

electrostatic sprays, electrical discharge from excessive build-up of residual charge on the aircraft skin has been a major problem in utilizing electrostatic systems for aircraft spraying. An idea of bipolar electrostatic charging has

been employed which shows that the bipolar charging could effectively abate the residual charge build-up. In addition, the deposition efficiency of aircraft sprays could be significantly enhanced. Atomization and charging

behaviors of the atomizer were studied under a laboratory simulation using a Malvern diffraction technique and a Farady cage collector, respectively. The deposition characteristics were investigated by carrying out a field

experiment with an aircraft spray system and a water-sensitive dye paper sampling technique. Sampled dye papers were digitized and processed to obtain the aerosol diameters and other statical data. The bipolar spray configuration was able to effectively eliminate the electric discharge of the residual charge on the aircraft skin. Compared with a corresponding neutral spray, for bipolar charging the overall SMD was increased by over 10 % showing significant

decrease in the number of small droplets and increase in large drops. This effect would reduce the drift problem of

aircraft sprays. The total deposition mass was also increased by more than 25 % when the spray was charged by dual

polarities.

INTRODUCTION

Aircraft spray systems are extensively utilized for applying pesticides to vast agricultural fields and hard-toreach terrain like mountain areas. It is well known that electrostatically charged sprays enhance the deposition

efficiency of droplets on targets because of the Coulomb attractive force between the charge on the atomized aerosol

and its image in the collector [1]. Carlton and Bouse [2] demonstrated experimentally with a ground-based

electrostatic spray system that the deposition efficiency can exceed that of its uncharged counterpart by as much as 800 percent.

For the case of aircraft application of the electrostatic sprays, where the aircraft releases unipolar-charged

aerosols, a surface charge of the opposite polarity develops on the aircraft skin. Excessive build-up of the residual

charge creates an electric field high enough to generate localized electric discharges (Fig. 1-A). Since the electric

discharge results in undesirable fluctuations in both charge distribution and electric field around the aircraft, a buildup of aircraft surface charge may degrade the deposition efficiency of charged aerosols [3].

An innovative idea which uses alternatively charged nozzles on an aircraft spray boom (Fig. 1-B) has been

proposed and investigated as a way to minimize or eliminate the build-up of aircraft surface charge so that

electrostatic aircraft sprays can ensure an improved deposition efficiency. In addition, this approach offers a

possibility of reducing spray drift since oppositely charged aerosols will coalesce together, and smaller aerosols will

coalesce far more quickly compared larger ones because of their higher electric mobility [4]. The present work consists of two primary parts: (1) a laboratory study of atomization and charge

characteristics of spray aerosols from a serrated spinning cup atomizer, and (2) a field study of aerosol deposition and

coalescence characteristics of bipolar charge aircraft spray system. Both studies used tap water as the atomization

fluid. A spinning cup atomizer was employed as this type of system is simple, lighte and provides aerosol diameters

139
Fig. 1 Comparison of electrostatic aircraft sprays: (A) a single-polarity charged spray

system, (B) a double-polarity charged spray system.

Fig. 2 A configuration of a DC motor-driven spinning cup atomizer with serrated teeth

140
of a fairly uniform distribution with a reliably small standard deviation [5]. Spray charging was attained with an

induction charging ring installed near the periphery of the spinning cup. A He-Ne laser Fraunhofer diffraction

particle sizer (Malvern 2600 D system) was used to measure the aerosol size for the laboratory study. The amount of

spray charge was measured by collecting charged aerosols with a Farady cage. For the aircraft spray field study,

aerosols were sampled with water-sensitive dye papers, and then the marked papers were digitized and processed for

statistical data.

EXPERIMENTAL SETUP

Serrated Spinning Cup Atomizer

The general configuration of a DC motor-driven spinning cup atomizer can be seen from the cross-sectional

view in Fig. 2. The design of the atomizer was based on Uie work of Bals [6], and a laboratory type atomizer, which

had been constructed by Carlton and Bouse [2] was employed for the present study. The cup has a dimension of 55

mm in diameter with 360 uniformly shaped triangular teeth with a serration pitch of 480 micrometers. The spinning

cup, C, rotates and receives water from the two liquid supply needles, A^. Atomizer parts were made of

nonconducting materials whenever possible to achieve high-voltage insulation and to minimize the electric field

distortions and/or the electric discharge. When a positive potential is placed on the cylindrical ring, R, with the

negative terminal from the power supply connected to the needles, droplets near the serrated periphery are negatively

charged because of the electron induction mechanism and, vice versa for a negative potential.

Laboratorv Setup

Figure 3 presents a schematic diagram of the experimental configuration. The spinning cup atomizer was

affixed to a spray boom section, which was located downstream of the air blower outlet. Tap water was fed through

the boom. TTie blower outlet section had dimensions of 154 mm in diameter and 0.6 m in length, and contained a honeycomb-type flow straightener. Air was blown into the outlet through the reservoir with a 40 horse-power

electric motor-driven squirrel-cage type blower. The maximum air speed at the outlet was 175 km/hr, which was

approximately equal to tfie typical cruising speed of an agricultural spray aircraft [7].

The atomizer was installed 15 cm downstream of the blower outlet and both the atomizer and the outlet were

vertically adjustable. The Malvern system was mounted on a horizontal two-dimensional positioning system.

Thus, the whole system was capable of precise three dimensional positioning of the atomizer. The Farady cage unit

consisted of three collectors to maximize the collection efficiency: a needle collector, a plane-mesh collector, and a concave-mesh collector. The meshes were 0.3 mm diameter aluminum wires with a mesh size of 1.5 mm x 1.5

mm. All three meshes were connected to a microamperemeter to measure the spray current.

Field Setup

Figure 4 shows a schematic representation of the field setup for a ground-level sampling of droplets from

aircraft sprays. A movable sampling stand, 60 meters in length, was setup on a landing strip perpendicular to the

aircraft cruising direction. The spray aircraft was flown in the upwind direction in an attempt to maximize the

sampling efficiency with longer airborne period of aerosols. Forty water-sensitive dye papers (25 mm x 75 mm

each) were placed vertically at 0.5 m above the ground and at 0.6 m intervals along the stand. Ten atomizers were

installed along a boom section attached behind the trailing edge of the right hand wing of a Cessna Model P-206B

aircraft. Wind speed and other atmospheric conditions, such as humidity and temperature, were measured and recorded

with a portable weather station. The aircraft cruising altitude and speed during spraying were maintained at 3 meters

and 175 km/hr, respectively.

RESULTS AND DISCUSSION

The results are presented for three different aspects of the study: (a) laboratory measurements of aerosol mean

diameters for various atomizer operating conditions (e.g., spinner rpm, liquid flow rate, air velocity and charging),

(b) laboratory measurements of the amount of spray charge for different charging voltages, and (c) field measurements

of aerosol deposition and size spectra.

Aerosol mean diameters

Figure 5 shows different disintegration mechanisms of a spinner atomizer as functions of spinner rpm and

water feed rate. For a given spinner revolution rate, when the feed rate is small, the liquid can not cover the entire

surface of the spinner. The liquid leaves the edge of the spinner as a form of dripping drops (drop formation

atomization). Like usual dripping of liquid, a significant number of secondary drops are formed. As the feed rate increases, a thin liquid sheet covers the entire spinner surface and ligaments are formed at the spinner periphery which

141
Ha-Ne Laser Baam

Concave-Mash

Collector

Farady Cage

Collector Unit

Air Reservoir

Fig. 3 Schematic diagram of experimental configuration for measurements of spray size and

charge.

Wind

Fig. 4 Schematic representation of tiie field setup of a ground sampling system for

electrostatic aircraft spray.

142
disintegrate into numerous drops (ligament disintegration atomization). The distribution of drop diameters becomes

fairly uniform with a minimal standard deviation. When the feed rate exceeds a critical value, the spinner surface is flooded and the liquid sheet extends outside the periphery. The extended sheet disintegrates into drops irregularly

(sheet disintegration atomization). The drop diameters become highly polydispersed. Using a direct visualization

technique, Tanasawa et al. [8] observed that minimal drop diameters were obtained in the ligament disintegration

condition. For the present study, all measurements were carried out within the ligament disintegration region.

As schematically illustrated in Fig. 6, radially spiral spray is developed in the absence of air stream while an

axially cylindrical spray is formed under the action of convective air stream. To account for such dramatically

different spray developments, two different sampling locations were considered for the measurement of mean

diameters: point A was used when there was no superimposed air stream. This is located 150 mm in the radial

direction from the spinner axis. Point B, located on the spray axis 122 mm downstream from the spinner edge, was

used when there was a superimposed air stream. At both locations, stable and highly repeatable values of Sauter

Mean Diameter (SMD) were measiu-ed. It must be noted that inconsistent SMD values could be measured if the

sampling location is not properly selected, especially for sprays with a superimposed air stream,

.

Measured data of SMD both with and without the air flow are presented in Fig. 7 as functions of liquid feed

rate for four different spinner revolution rates. The symbols with dashed lines represent SMD measured with a

convective air stream of 175 km/hr (or equivalently 49 m/s), and the symbols with solid curves represent data taken

with no air flow. As the spinner revolution rate increases while the liquid feed rate remains unchanged (within the

Ugament breakup condition), the cross sectional area of hgaments formed at the tips of serrated teeth of the spinner

must be decreased because of mass continuity considerations. In primary atomization, ligament usually disintegrates

into drops with sizes comparable to its own diameter. Therefore, reduced cross sectional area of ligaments can

explain the reduction in SMD when the spinner rotation is increased. The reverse may be true for the case where the

liquid feed rate is increased for a fixed spinner revolution. For the case of no air flow, the increase of SMD with the

increase of liquid feed rate can be attributed mainly to the enlargement of ligaments.

When cross-sectionally enlarged ligaments are exposed to a high velocity air stream, viscous pulling action of

the air may elongate the ligaments longitudinally, which could produce longer ligaments of nearly unchanged or even

slightly reduced cross sectional area. The nearly constant or slightly decreasing SMD in the presence of the

convective air flow could be explained by this. Evaporation from the liquid surface by a high speed air stream may

also contribute to a further reduction in SMD. The ranges of the aerosol diameters were found to be smaller than the

critical diameters [9,10, 1 1], at which aerosols break up due to the inertial force of the air stream overcomming the

surface tension force. The probability of this secondary atomization will be very low for the conditions investigated.

Sprav charge

Amounts of spray current (C/s) were measured with the Farady cage collector and results are presented in

terms of charge to mass ratio Q/M (C/kg) in Fig. 8. The charging voltage was allowed to vary while the spinner

rotational rate was held at 9000 rpm for three different liquid feed rates (0.75 ml/s, 1.0 ml/s and 1.5 ml/s) and for an

air speed of 175 km/hr. No significant difference was observed for different polarity of charging voltage. The charge

to mass ratio increased with the charging voltage and showed maximum values at approximately 7 kV. The charge

to mass ratio sharply decreased thereafter due to the electrical break down as a result of the excessive electric field

between the induction charging ring and ground (the liquid in this case). As the liquid flow rate was increased, a higher amount of total spray charge was measured. The charge to mass ratio, which would be a more accurate measure of aerosol charging, showed a decrease with an increase in liquid flow rate. Since the aerosol SMD remained

nearly unchanged as the liquid flow rate was varied under a constant spinner revolution rate (Fig. 7), the amount of

surface charge of individual droplets will be directly proportional to the charge to mass ratio. Although not

presented, the charge to mass ratio demonstrated a slighUy increasing function of spinner revolution rate for a given

liquid flow rate. Since the induction charging increases with the surface area of aerosols, the increase of the charge to mass ratio is attributed to reduced SMD with increased spinner rpm and increased total surface area of aerosols for the same liquid flow rate. An excessively charged liquid drop can disintegrate into smaller droplets when the Coulomb repulsive force

(the outward pressure) produced by the surface charge equals or exceeds the holding force (the inward pressure)

produced by the surface tension. This condition is called the Rayleigh limit [1], and is defined as a limiting charge at or above which aerosols will disintegrate by the electric repulsion. When this occurs, the mean drop diameter will

be significantly reduced [12]. Drop diameters of charged sprays were measured while the air speed and the feed rate remained unchanged. Results, including the uncharged counterparts, are presented for two different spinner revolution

rates and for three different liquid feed rates in Fig. 9. Up to the peak charging voltages corresponding to the

maximum charge to mass ratio, the aerosol SMDs remained nearly unchanged for both spinner revolution rates. The

maximum value of the charge to mass ratio was measured to be on the order of 10'^ C/kg whereas the Rayleigh

limit is a charge to mass ratio of about 0.1 C/kg for the same mean diameter. The amount of aerosol charge with

143
28000

?

Formation

WATER FEED RATE (ml/s)

Fig. 5 Different atomization mechanisms of spinner atomizer functions of the spinner

revolution rate and the liquid feed rate.

Fig. 6 Spray configurations and sampling locations for the spray under convective air stream

(A) and for the spray without air flow (B).

144
Air Speed (km/hr)

175

N

rpm= 3000

, rpm= 6000

Y \ rpm» 9000

rpni' 12000

T

10 12 WATER FEED RATE (ml/s)

Fig. 7 Spray SMD both with and without the air flow as a function of water feed rate for

four different spinner revolution rates.

0.010

o.oos

M

— o.ooe oO

^ 0.004 -

0.002 -

0.000

Air 8pMd- 178 km/hr

Splnnar Revolution Roto

Wator Food Roto (ml/o)

0.78 —^— 1.0

1.8

CHARGING VOLTAGE (KV)

Fig. 8 Spray charge to mass ratio versus induction charging voltage for three different water

feed rates.

145
the present configuration was not sufficient to disintegrate the drops further by overcoming the surface tension

holding force.

Aerosol deposition

For the field experiment, the liquid feed rate was maintained at 1 mils per atomizer, the rotational rates were

set at 9000 rpm and a charging voltage of 7 kV was used to provide a maximum value of the charge to mass ratio. An electric discharge pole was installed at each wing tip to monitor the electric current discharged to the

environment A reading of 25 was recorded when five sprays were positively charged with a negative polarity.

As soon as the ten alternatively-charged sprays were activated with nearly equal values of positive and negative

charging voltages, the reading was gradually decreased and a zero reading was attained within a second. This proved

that the bipolar electrostatic charged spray system elliminates any noticeable build-up of residual charge on the

aircraft

Figure 10 presents lateral distributions of SMD and the number of deposited drops for both charged and

uncharged cases. For each test, aircraft spraying was repeated three times over the same path to average

meteorological differences. The average wind speed was 13 kmlhr and the relative humidity was 45 %. The zero

lateral location was set at the center of the aircraft. Since aerosols will spread and leave larger marks when they are

collected by the water-sensitive dye paper, a spread factor must be considered in order to determine the true aerosol

diameters. The spread factor is defined as the ratio of the marked diameter to the true aerosol diameter. For the

conditions of our experiment, a spread factor of two was used as suggested by the manufacturer. The reduction in

numbers of deposited drops from the case of uncharged spray to the case of charged spray may result from the

coalescence of bipolar charged aerosols in the aerosol state. Also, the bipolar coalescence causes the mean diameter

to increase. The charged spray was more uniformly dispersed showing less dependence of SMD and number count

upon the lateral location. It is believed that this is because electrostatic coagulation preferentially occurs between the

largest and smallest sizes.

The size spectra of all the deposited aerosols are presented for both cases in Fig. 11. The values of SMD of

all the collected drops were 92.2 fXTTI for the neutral spray and 103.3 /X/77 for the charged spray. It is apparent

that for the case of charged spray the size spectrum was shifted toward a larger mean value with a significant

reduction in small drops less than 50 fxm in diameter and a noticeable increase in large drops of 120 fxm or

more in size. Volume mean diameters (VA/Dj') were calculated to be 99.7 jim and 110.7 fim for uncharged and

charged sprays, respectively. By multiplying the VMD by the total particle count for each case, the ratio of

deposited mass with charging to that without charging was determined to be 1.27. Thus, the deposited mass was

increased by over 25 % when charged sprays were used.

SUMMARY AND CONCLUSION

Characteristics of atomization, charging and deposition were investigated for a bipolar electrostatic aircraft

spray system. The system utilizes a serrated spinning cup atomizer. Both laboratory measurement and field

experiment were carried out and major conclusions of the study are:

(1) The Sauter mean diameter of atomized aerosols remained nearly unchanged for varying liquid feed rates under a convective air stream of 175 kmlhr, which is a typical crusing speed of an agricultural spray aircraft.

(2) The maximum value of the charge to mass ratio was measured to be on the order of 0.01 Clkg with a charging

voltage of about 7 kV. The charge to mass ratio decreases with an increase of the liquid flow rate. The

maximum charge to mass ratio was found to be below the Rayleigh limit by an order of magnitude and aerosol

SMD remained unchanged with an increase of the charging voltage.

(3) A successful elimination of the build-up of residual charge on the aircraft skin was achieved with bipolarcharging aircraft sprays.

(4) Field studies showed the SMDs of deposited aerosols were 10 % larger for aerosols of bipolar charged sprays as compared with neutral sprays. A reduction in the number of small droplets and an increase in the number of

large droplets were noted when bipolar charging was used. This can be attributed to the coalescence of bipolarly

charged aerosols. The total amount of deposited mass was also increased by over 25 % when the aircraft spray was bipolarly charged.

146
100

80

E

o

o

E

CO

00 -

40

20 -

Air 8p«*d- 178 km/hr

Water FMtf Rata- 1.0 ml/a

Splnnar Ravolirtlen Rata (rpn)

9000

-O- 12000

T

2

T

8

CHARQINQ VOLTAGE (KV)

T

0

Fig. 9 Independence of spray SMD with regard to charging voltage.

I 10

125

E

o

o

E

CO

100 -

76 -

60 -

26

-4

I J3

cr

-[>• NMMral

Chara* (a M 0)

(N«aik«r) ChHfa (Naakw)

-r

2

-r

1 0 4 0 0

LATERAL SAMPLING LOCATION (meter)

200

180 COH

o

CLCO

- 100 O

c

lU

09

- BO

10

Fig. 10 Lateral distributions ofSMD and the number of sampled drops for both dual-polarity

charged and uncharged aircraft sprays.

147
Fig. 1 1 Comparison of size distributions between neutral and dual-polarity charged spray.

ACKNOWLEDGEMENTS

The financial support for the present work was provided by the United States Department of Agriculture through the

Agricultural Research Service. Authors wish to thank Dr. L.F. Bouse and Dr. J. Carlton of the USDA Agricultural

Research Service for their many valuable suggestions and help throughout the work.

REFERENCES

1 . Davies, C.N., Aerosol Science. Academic Press, Ch. 3 (1966).

2. Carlton, J.B., and Bouse, L.F., "Electrostatic spinner-nozzle for charging aerial sprays," Trans. ASAE. Vol. 23,

No. 6, pp. 1369-1378 (1980).

3. Carlton, J.B., and Bouse, L.F., "Distribution of the electric field for an electrostatic spray charging aircraft,"

Trans. ASME. Vol. 20, No. 2, pp. 248-357, (1977).

4. Oglesby, S., and Nichols, G.B., Electrostatic Precipitation . Marcel Dekker Inc., Ch. 2 (1978).

5. Lefebvre, A.H., Atomization and Sprays. Hemisphere Publishing Co., pp. 127-134, pp. 189-193, 222-228,

(1989).

6. Bals, E.J., "Design of rotary atomizers," Proc. 4th Int. Agric. Aviat. Congr. . Kingston, UK., pp. 156-165

(1969).

7. O'Neal, H., and Brazelton, R.W., "Aerial Application of Pesticides," Report No. ADP 84-016, Shell Chemical

Company, p. 22, (1984).

8. Tanasawa, Y., Miyasaka, Y., and Umehara, M., "Effect of shape of rotating disks and cups on liquid

atomization," Proc. 1st Int. Conf. Liquid Atomization and Sprav Systems. Tokyo, Japan, pp. 165-172 (1978).

9. Lane, W.R., "Shatter of drops in the stream of air," Ind. Eng. Chem., Vol. 43, No. 6, pp. 1312-1317, (1951).

10. Hinze, J.O., "Fundamentals of the Hydrodynamic Mechanism of splitting in dispersion processes," AICHE J., Vol. 1, No. 3, pp. 289-295, (1955).

11. Richardson, S.M., Fluid Mechanics. Hemisphere Pub. Corp., pp. 260-264, (1989).

12. Cross, J.A., Electrostatics. Adam Hilger, pp. 198-218 (1987).

148
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper II

A NUMERICAL MODEL OF THE CAPILLARY INSTABILITY

R.W. Sellens

Department of Mechanical Engineering

Queen's University at Kingston

Ontario, Canada

ABSTRACT

The momentum and continuity equations are reduced to a one dimensional form for the case of a

finite axi-symmetric disturbance on a liquid ligament, retaining both viscous and inertial terms. Numerical

solution of these equations yield results that are in good agreement with experimental data and the exact

solutions of Rayleigh and Weber for infinitesimal disturbances. The model carries through to the point of

breakup and predicts the formation of satellite drops between the main drops of the stream.

INTRODUCTION

The capillary instability of cylindrical ligaments of liquid is of fundamental importance in the atomization of liquids. This mechanism is key in both practical atomizers and a variety of special purpose devices

such as monosize drop generators and ink jet printers.

A cylindrical element of liquid in a quiescent environment is inherently unstable, so that small disturbances will grow larger. The first quantitative solution to this problem was presented by Rayleigh [1], who linearized the system by neglecting viscous forces and restricted his analysis to sinusoidal disturbances

of small amplitude. Weber [2], considered a viscous system, but neglected the inertial effects, again for

small disturbances. Both Rayleigh and Weber found that disturbances at unstable wavelengths would grow

exponentially so that

^ = ^oe°" where ^ = (Rmax - Rmin)/2 (1)

with the growth rate a dependent on the geometry of the system and the fluid properties.

^ , . , 5 <T (1 - Pa'^)ka Ii(ka)

Rayleigh =

' 2)

pa-* loyka)

Weber + a^Pa-" = -^(1 - Pa')k-'a'' (3)

R is the local ligament radius, a is the undisturbed ligament radius and k is the disturbance wave number,

2ir/\. By examination of equation 3 it is clear that the importance of viscosity increases with decreasing

initial ligament radius. The scale and fluid properties in most atomization flows require that the eff"ects of

viscosity be considered.

Subsequent analytical solutions (c.f Sterling and Sleicher[3]) have generally been restricted by one or more assumptions; inviscid flow, negligible inertia, or small disturbance amplitude. The assumption of small

149
disturbance amplitude prevents the application of these models to the later stages of the breakup process

where the ligament is pinching off to form separate drops, generally leaving a "satellite" drop in the pinched

off region.

The instability is driven by the surface tension force, which produces pressure variations dependent on the local curvature of the liquid surface. Thus, as surface geometry is of first order importance, the

modelling of the moving interface is critical to success. Numerical methods developed for fluid systems with

fixed boundaries, or with boundaries where surface tension may be neglected are ill suited to the solution of

this problem.

Mansour and Lundgren [4] have presented a solution method for the growth of finite amplitude,

spatially periodic, axisymmetric waves on an inviscid cylindrical ligament using a boundary integral technique

to solve the potential flow equations. The solution can follow the ligament growth to breakup and predict

the formation of satellite drops. Growth rates are in good agreement with the classic result of Rayleigh.

However, this approach holds only for inviscid flow, and it fails where small size or high viscosity violate this

condition.

Shkoohi and Elrod [5,6] have solved the complete Navier-Stokes equations for both Lagrangian and

Eulerian coordinates. However, the computational requirements of this approach have limited their studies.

The present approach considers a simplification of the equation system in hopes that the essential

character of the solution may be preserved, while dramatically reducing computation time. This will permit an assessment of the effects of a wide variety of different disturbance functions on the ligament breakup,

particularly on the sizes of drops produced. If this system is considered statistically, it may shed some light on the processes limiting the formation of small drops in low speed atomization flows. In previous work

by the author [7,8] a physical constraint limiting the formation of small drops was found necessary in the

statistical prediction of drop size and velocity distributions.

EQUATIONS OF MOTION

The surface of an axisymmetric ligament may be characterized by its radius as a function of axial

position. If the radial velocity is neglected and the axial velocity within the ligament is assumed to be

radially uniform, it may also be characterized as a function of axial position, thus reducing the problem to a transient, one dimensional problem.

Continuity for an incompressible, disk shaped element requires that the radius, R, increase to accom- modate the net mass flux into the element so that

or rearranging explicitly for the temporal change in R

dR dR Rdw

A balance of axial momentum on an element of fluid yields

dw dw IdP ^ d'^w AudRdw ...

ot oz p oz oz^ R oz oz

where the second viscous term accounts for the variation of the radius. The pressure immediately beneath

the surface may be calculated by considering the forces on an element of surface curved in both directions.

In keeping with the one dimensional velocity approximation, the pressure may be approximated as radially

uniform. Thus, differentiating equation 7 and substituting into equation 6 yields an explicit expression for

150
the temporal change in velocity.

dt ~ ^ dz^ pR? dz pR^ \dz) ^ pRdz dz^

ad^R ^ d^w AvdRdw

+ + + ——— (8)

p oz^ dz^ R dz dz

Equations 5 and 8 then provide a coupled system of nonlinear equations describing the development of the

two variables, R and w. An appropriately chosen numerical solution technique may then be used to follow

this development under varying initial conditions.

DISCRETIZATION AND SOLUTION OF EQUATIONS

The ligament is discretized into a number of elements, each characterized by values of R and w at the centre of the element. Equations 5 and 8 may then be approximated about each element by evaluating

first order central differences for each of the derivative terms. (Although various differencing schemes were

tried, no improvement on the initial scheme was found.) This approach provides estimates of the temporal

derivatives based on the current values of R and w. The problem is then solved by a semi-imphcit time marching scheme. For each time step, the following

process is employed.

1. Approximate the sets of new values wi and Ri to be equal to the sets of current values wq and Rq.

2. Evaluate (^)o and (^)o based on wq and Rq.

3. Evaluate (^)i and (^)i based on Wi and Ri.

4. Calculate w, = wo + 0.5 ((^)o + (^)i) At and R^ = Ro + 0.5 )o + )i) At

5. Repeat from step two until the values of Wi and Ri converge.

By repeating this entire process the solution may be marched forward. At at a time. At present, periodic

boundary conditions are applied, so that the difference approximations are wrapped around to the opposite

ends of the solution domain.

Initial conditions on the ligament may be set to a sinusoidal variation in the radius along the ligament, a sinusoidal variation in velocity, or a combination of the two.

Instabilities occasionally develop in the solution, particularly for cases where the viscous forces are

small. These instabihties are on the scale of the discretization and are first evident in the time derivative of

the velocity. Reducing the time step will eliminate or delay the onset of these instabilities, however it has

been found more effective to apply a 3 or 5 point moving average on the derivative field to suppress this

instability.

RESULTS AND DISCUSSION

For comparison with the classic results and experimental data, the overall progress of the solution

may be characterized by the exponential growth rate a, as defined in equation 1. Although a constant for

the classical solutions, the growth rate is found to vary with solution progress. Figure 1 shows the results

obtained for a when the solution parameters are varied. As can be seen by comparison with denser grids,

the 60 point discretization yields good results, provided that no averaging is used to stabilize the solution.

All results presented in this paper are for a 60 point discretization with no averaging, except where noted

otherwise.

The 90 and 120 point solutions are very close to the 60 point results, but averaging and/or time step

reduction were necessary to stabilize the solutions. The equations were solved on a 20 MHz, 80386 based

personal computer with numeric coprocessor. Overall solution times to carry a ligament through to the

151
340

330-

<

D

a:

320-

JZ 310

O

300

290

Water D=0.9mm ka=0.714

Varying Solution Parameters

0.0000

1 1 1 1 1 1

1

1

1

1 1 1 1 1 1 1 1 1

1 1 1

1

0.0001 0.0002

60 pts, no averaging

60 pts, 3 pt avg ***** 60 pts, 5 pt avg

60 pts, 7 pt avg

* * * * * 90 pts, no averaging

90 pts, 3 pt avg

120 pts, 5 pt avg

"T~i I I I I I I I I 0.0003 0.0004

Disturbance Amplitude (m)

Figure 1: Varying the solution parameters has little effect on the solution behaviour, as shown by the growth

rate a, except when the domain of the moving average represents a significant portion of the solution domain.

point of breakup were typically around 20 minutes, 3 hours and 60 hours for 60, 90 and 120 point solutions

respectively. Naturally, these times varied substantially with differences in initial conditions, being longer

for smaller initial disturbances.

The code presently allows sinusoidal variations in amplitude and/or velocity to be applied to the

ligament. Figure 2 shows that the solutions tend toward a common result, independent of the characteristics

of the initial disturbance. Only the wavelength of the disturbance is significant. Thus, an infinitesimal

disturbance may be simulated by any combination of amplitude and velocity, provided it is small.

A characteristic growth rate has been arbitrarily defined as the value of a when the disturbance

amplitude, peak to trough, has reached 1/4 of the undisturbed ligament radius. As can be seen from the

figures, this provides a result that characterizes the early part of the growth for a specific wave number, but

is insensitive to the other characteristics of the initial disturbance.

Figure 3 shows comparisons with the experimental dataof Sakai et al [9] for glycerine/water solutions.

A vibrating orifice generator was used to produce a ligament with imposed disturbances of known wavelength

and growth rates calculated from photographic measurements.

The results agree quite well for the lower viscosity case of 65% glycerine, but the present approach

over-predicts the growth rates for the higher viscosity case. For pure water, the present model predicts

growth rates approximately 3% higher than Rayleigh's solution. Otherwise, the model shows the same

characteristics as the classical results, yielding a maximum growth rate around ka = 0.69 for low viscosity cases and moving towards ka = 0.5 as viscosity increases.

This deviation from the classical results appears to be due to the one dimensional approximation,

which eliminates the effects of the radial velocity and its variation. From the qualitative agreement it is

clear that this is a secondary effect. It may be possible to model the radial resistance to growth in terms of

the current solution variables R and w to produce a quasi-two-dimensional model that retains the simplicity

and computational efficiency of the ID model while producing more accurate results.

The variation of a with solution progress is shown in figures 4 and 5 for water and 65% and 90%

glycerine/water solutions respectively. The growth rate remains more nearly constant for the more viscous

flows.

152
Water D=0,9mm ka = 0.714

Varying Initial Disturbance

,002 mm

im/s

0 mm/s

m/s

m/s

.020 mm

m/s

_ ...m/s

V = 8.55 mm/s

V = 10 mm/s

I I 1 1 1 1

1

1 1

1

1 1 1 1 1

1

1

1 1 1

1

1 1 O.ObO 0.005 0.010

Time

I I I I

I

I I I 0.015

(s)

I I I I I

I

I I M I I I I 0.020 0.025

340

330- D

<

CD

-I—' D

cr

XI 310:

320-

5

O

300

290

Water D=0.9mm ka = 0.714

Varying Initial Disturbance

Amplitude = 0.002 mm

° ° ° ° p V = 0 mm/s

V = 0.860 mm/s

** *** V = 1 mm/s

« » « « V = 5 mm/s

Amplitude = 0.020 mm

* * * * * V = 0 mm/s

V = 5 mm/s

V = 8.55 mm/s

* « ' ' V = 10 mm/s

I I I I J I I I I I I I I I I

I I I I I I I I I I I 0.0000 0.0001 0.0002 0.0003

Disturbance Amplitude (m)

^004

Figure 2: Variations in the initial disturbance characteristics affect only the early part of the solution. The

disturbance wavelength defines a solution that different types of disturbance move towards. The left hand

plot shows that disturbance amphtude grows exponentially with time, while the right hand plot shows the

variation of that growth rate with disturbance amplitude.

Figure 3: This figure compares the growth rates found for different wave numbers on 0.9 mm diameter

ligaments of 65% and 90% glycerine/water solutions.

153
Figure 4: Growth rate behaviour in water at various dimensionless wave numbers.

400 400

jz 300-

<

-4—' O 200

5

O

100-

o-f

90% Glycerine - D = 0.9 mm

5 mm/s Velocity Variation

0.40

0.60

0.67

0,80

0.90

0.0000 0.0002 0.0004 0.0006 0.0000

Disturbance Amplitude (m) Disturbance Amplitude (m)

T~i 1 I I (""1 I r

I

I—I—r~i—I—I—r~~i—i

\—i—i—i—i—i—r—r—i

0.0002 0.0004 0.0006

Figure 5: Growth rate behaviour in 65% and 90% glycerine/water solutions at various dimensionless wave

numbers.

154
Figure 6: Surface profiles for a 0.9 ram ligament of water at diff'erent times, ka — 0.7; lines are 2 ms apart

0.0010

^ 0.0005

O

cr

0.0000

Lines every 2 milliseconds

Water ka = 0.9

Diameter 0.9 mm

O.ObO 0.001 0.002 0.003 0.004 0.005 0.006 0.007 0.000 0.001 0.002 0.003

Axial Location (m) Axial Location (m)

Figure 7: Surface profiles for a 0.9 mm ligament of water at diff'erent times. The aspect ratio has been

exaggerated to provide more detail. On the left the dimensionless wave number is ka = 0.4, while on the

right ka = 0.9.

Figures 6 and 7 show the development of the surface profiles for ligaments of water, subject to

disturbances of different wave number. Figure 6 is plotted in the correct aspect ratio, showing the ligament

for two wavelengths and provides a clear picture of the flow. The plots in Figure 7 are exaggerated in the

radial direction to show more detail.

It was observed that all instances yield satelhte drops, but that these drops become much smaller

at higher wave numbers. This is consistent with the two dimensional inviscid results of Mansour and

Lundgren [4]. The breakup process could not be followed past the point of pinching off due to current

limitations in the code. Inclusion of a boundary condition for a pinched off end is planned to permit

following of the secondary breakup of the satellite ligaments produced at low dimensionless wave numbers.

However, this extension is not important for the higher wave number cases where both main and satellite

drops are well formed as drops by the time the ligament pinches off.

CONCLUSION

A model has been developed which incorporates both the inertial and viscous forces in the capillary

instability. Comparisons indicate a tendency to over-prediction of disturbance growth rate, particularly in

highly viscous cases, due to the approximations in the one dimensional formulation.

155
However, qualitative agreement is very good and a refinement of the model to incorporate inertia!

and viscous effects of motion in the radial direction is expected to yield quantitatively accurate results.

The computational requirements of this model are much smaller than fully two dimensional solutions,

while providing useful results.

ACKNOWLEDGEMENT

The author is grateful for the support of Queen's University for this work, through both Research

Initiation and Advisory Research Committee Grants.

REFERENCES

[1] Rayleigh, Lord, "On the Instability of Jets," Proc. London Math. Soc, Vol. 10, pp. 4-13, (1878).

[2] Weber, C, "Zum Zerfall Eines Fliissigkeitsstrahles," Z. Angew. Math. Mech., Vol. 11, pp. 138-145, (1931).

[3] Sterling, A.M. and Sleicher, C.A., "The instability of capillary jets," J. Fluid Mech., Vol. 68, pp. 477-495,

(1975).

[4] Mansour, N.N. and Lundgren, T.S., "Satelhte Formation in Capillary Jet Breakup," Phys. Fluids A,

Vol. 2, No. 7, pp. 141-1144, (1990).

[5] Shkoohi, F. and Elrod, E.G., "Numerical Investigation of the Disintegration of Liquid Jets," J. Comput.

Phys., Vol. 71, pp. 324-342, (1987).

[6] Shkoohi, F. and Elrod, E.G., "Algorithms for Eulerian Treatment of Jet Breakup Induced by Surface

Tension," J. Comput. Phys., Vol. 89, pp. 483-487, (1990).

[7] Sellens, R.W., "Prediction of the Drop Size and Velocity Distribution in a Spray, Based on the Maximum

Entropy Formalism," Part. Part. Sys. Charact., Vol. 6, pp. 17-27, (1989).

[8] Sellens, R.W., "Drop Size and Velocity Distributions in Sprays, A New Approach Based on the Maximum

Entropy Formahsm," Ph.D. Thesis, Department of Mechanical Engineering, University of Waterloo,

Waterloo, Ontario, Canada, (1987).

[9] Sakai, T., Sadakata, M., Saito, M. and Matsushita, K., "Studies of Disintegration of Liquid Column

Between Production of Uniform Drops by Vibration Method," International Conference on Liquid Atomisation and Spray Systems (ICLASS-85), paper VIIB(b)/2, The Institute of Energy, London, England,

1985.

156
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 12

ATOMISATION OF FLAT AND ANNULAR LIQUID SHEETS

PRACTICAL USE OF LINEAR THEORIES

C. Dumouchel and M. Ledoux

URA CNRS 230/CORIA - University of Rouen

Mont Saint Aignan, France

In this paper atomisation of flat and cylindrical liquid systems have

been investigated through linear theories. After an examination of the

parameters involved in the problem it is found that the behaviour of flat

systems is totaly known by the determination of a set of non dimensional i zed

numbers and that cylindrical systems behaves as flat ones in many cases.

Associated to a disintegration process that allows to calculate one drop

diameter, applications of these mathematical models on practical situations

are described. Some experimental observations concerning the influence of the

external pressure are found again and an interesting agreement between a set

of calculated and measured drop diameters is obtained.

1 INTRODUCTION

Atomisation is based on interfacial instabilities of liquid systems

either in calm atmospheres or induced by air streams. This physical process is

not perfectly known and its complexity makes theoretical investigations tricky

to set up.

Even though it can be critized the linear theory based on the development

of a dominant wave length allows a quick approach to the problem and supports

comparisons with experimental data.

Examples of linear theories are presented here for flat and cylindrical

liquid systems. It is attempted to allow the exploitation of such theoretical

models by designers through a synthetic presentation of the results of

computations and to show the possibilities as well as the limits of such

theories.

2 INSTABILITY OF A FLAT LIQUID SHEET

2. 1 Analysis

(i-2)

Fig. 1. Flat liquid sheet in air at

rest ( undisturbed flow)

( perturbed flow)

M)

The disintegration of an inviscid

flat sheet of liquid is studied through

the set up of a linear theory. This

approach assumes that a superposition of

sinusoidal perturbations takes place on

the interfaces of the liquid system (see

fig. 1). Some of these perturbations will

grow sufficiently to lead to the break

down of the sheet. However^ among them^

one will have a growth rate that is

larger than the others and the

157
disintegration of the system will be dominated by the growth of this

particular perturbation.

The mathematics involved here are similar to those used by Crapper et al

[1] and Dumouchel et al [2] for the case of a cylindrical liquid sheet and

they are not reported in detail here.

The initial flow is steady but any perturbation induces a non-steady flow

assumed irrotational and thus associated to a perturbation potential <j>. In

each phase the continuity equation allows the functions 0^ to be written as :

<t>^= & {ke^^ + B^e"''^) expj^Kkx - wt)j (1)

where i refers to the regions indicated on fig. 1. The displacement of the

upper and lower surfaces shown in fig. 1 are taken under a similar form :

T)^ = € exp^iCkx - £.)t)j (2)

The pulsation w may be complex and its imaginary part is the temporal

growth rate corresponding to the imposed perturbation defined by a wave number

k. The fluid particles that compose the interfaces are supposed to be the same

all over the time and this is expressed through four kinematic conditions.

They allow the determination of the coefficients and B^. Two dynamic

conditions expressing the balance between the pressure and the surface tension

forces are written on each interface and the resulting system of equations

conducts to the dispersion equation that has the form :

C + C + C 0)^ + C a> + C = 0 (3)

4 3 2 1 0

(the coefficients may be obtained from the authors). For a given liquid

system, equation (3) is solved to find the value of w of the imposed

perturbation which will grow if the solution found is complex. According to

Hagerty and Shea [3] and Dombrowski and Hooper [4] remarks, one has only to

consider the antisymmetric mode of perturbation which is predominating for

small gas densities (p << p ).

2. 2 Results

Figure 2 shows dispersion

diagrams ie versus k, over a range

of velocity U for a sheet of water. In

each case it is seen that the

characteristics of the dominant wave

ie k and w both increase with

amax i amax the velocity U. For the same cases

fig. 3 shows the propagation velocities

Cel of the growing perturbations

versus the wave numbers k.

The growth rate of the

perturbation is only controled here by

the aerodynamical and surface tension

forces. On fig. 2 and 3 we can see that

aerodynamical forces are destabilizing

1000

750

(0

500

250

U-IO in/s U-7.5 m/s

. U-5 m/s

800

k

1200 1600 2000

(m"^)

Fig. 2. Dispersion diagram.

h=300jam (water/air)

158
12.5

10.0-

(0e

0) lu

0.0.

U-IO m/s

7.5

5.0L_..

2.5

U-7.5 m/s

U-5 m/s

400 1200 1600 2000

-1

800

Fig. 3. Propagation wave velocity.

h=300/im (water/air)

The dominant wave results

and proportional to the difference of

velocity between the external medium and

the perturbation propagation speed (for

small wave numbers u) and Cel both

lamax

increase with k).

The surface tension forces are

indirectly proportional to the curvature

radius of the interfaces and therefore

more effective for high wave numbers. For

these values of k the growth rate

decreases as k increases showing the

stabilizing effect of the surface tension

forces that increases with k. This effect

becomes more and more effective in a

domain where the aerodynamical forces are

maximum (Cel=U) but constant and that is

why w

iamax

decreases until it becomes

zero,

from the best compromise between the

aerodynamical and surface tension effects.

We studied the influence of each parameter involved in the problem on the

characteristics of the dominant perturbation ie k

amax

u> and Cel. Using a

iamax

set of non dimensionnal parameters the different results showed only one

behaviour presented on fig. 4 where :

2<r

We = p U^h

k <r amax

Wek = P ^

.00 a.25 0.50 0.75 1.00 1.25

We

Fig. 4. Characteristics of

dominant wave for

liquid sheets.

the

flat

Wew =

(j> cr

i amax

p u~

Cel = Cel

U

(4)

Figure 4 is relevant for any kind of flat

sheet evolving through any kind of

gaseous medium as long as the assumption

p « p is kept. It is seen that if

G L

We>l the corresponding liquid system is

stable for any perturbation. We find here

again the stability criterion found by

Squire [5]. We see as well that w

iamax

reaches a maximum corresponding to

We = 0. 17 Wew = 0. 125

(5)

Wek = 0.50 Cel = 0.75

The presence of this maximum may be

explained as previously for the maxima

shown on fig. 2.

159
3 INSTABILITY OF A CYLINDRICAL LIQUID SHEET

3.

1

Introduction

Encountered in amny practical situations, the disintegration of

cylindrical liquid sheets has been the subject of many theoretical and

experimental investigations [1], [5], [6]. Furthermore such studies provide a

first approximation of the complex problem posed by the conical liquid systems

as proposed by Dumouchel et al. [2].

Here their analysis is summarized and results are presented and compared

with those of flat systems to allow a better understanding of the process

described by linear theories in such cases.

3.2 Analysis

The cylinder of liquid is defined by a thickness h, an internal radius a

and two velocity components, U along the axis of symmetry and V along the

azimuth (as for a conical sheet). We make the assuumption that the liquid

velocity is sufficiently high and the liquid sheet thin enough so that the

sheet breakdown occurs before the tendency of the cylinder to collapse under

surface tension forces takes place. The analysis, not reported in detail here,

is conducted using cylindrical coordinates. The velocity V allows a perturbation to take place along the azimuth and this is introduced in the

propagating terms of the functions and t)^ that become :

A^I^Ckr) + B^K^(kr)j expj^Kkx + ne - wt)j (6)

7)^ = e exp^Kkx + ne - ti)t)j (7)

The internal radius and the possible azimuthal perturbation introduce new

curvature radii that modify the surface tension terms in the dynamic

conditions (for more details see reference [2]).

As for the flat sheet the dispersion equation obtained for w is of the

fourth order. Here again we only deal with the solutions corresponding to the

antisymmetric mode of perturbation.

3.3 Results

3.3.1 Without azimuthal component of velocity

As for a flat liquid sheet, the dispersion diagram for a cylindrical

liquid sheet shows a maximum corresponding to a perturbation that growths

quicker than the others. We are only interested by this particular

perturbation characterised by the numbers Wew, Wek and Cel defined by

equations (4).

A new parameter Wea is introduced to take into account the internal

radius a.

<r Wea = (8)

p U^a

Fig. 5 shows the relation between Wew and We for different values of Wea.

When Wea=0.048 we can see that the curve obtained is not very different from

fig. 4. Further calculations actually showed that a liquid cylinder having a

parameter Wea less than 0. 1 behaved like a flat system. This means that when

Wea<0. 1 the surface tension effects added by the curvature radius a are very

160
0.35

0.30

0.25

0.20

<^ 0.15

r. 0. 10|h\

0.05

0.00.

3

Hea-0.58 • • Wea-0.19

Mea-0.39 Wda-0.04B

r - 4-

1 2 3 4 5 6

We

Fig. 5. Growth rate of the dominant

wave for cylindrical liquid

sheets. Wew=f(We)

2.5

2.0

1.5

3

O 1.0

IS

0.5

O.Q

— We-0.120 - We-0.060 — We-0.030

• We-0.015

4 6 10

Wea

Fig. 6. Growth rate of the dominant

wave for cylindrical liquid

sheets Wew=f(Wea)

small compared to the aerodynamical effects or the surface tension effects due

to the perturbation and therefore they may be ignored. This criterion agrees

with the one found by Guilie in a similar study [6].

When Wea increases, we see on fig. 5 that Wew increases for any given We.

This shows that the surface tension forces introduced by the radius a are

destabilizing. We note as well that when We increases Wew reaches a limit

different from zero as for a flat system and that increases as Wea increases.

In this domain the disintegration is especially controlled by the surface

tension forces introduced by the radius a and it corresponds to the domain of

very thin systems as those studied by Martinon [7].

Fig. 6 shows that the relation between Wew and Wea is linear, the

expression of which, for We<0.6, was found to be :

Wew = 0.8 /we Wea -1. 1 We" + 0. 11 We + 0.07 (9)

As for a flat sheet it can be seen from fig. 5 and equation (9) that the

maximum growth rate of a cylindrical system reaches a maximum for given We and

Wea.

Fig. 7 shows the values of Wek and

Cel for Wea=0. 58 and it is seen that

both of them reach a limit for high

We. However we can see that the limit

of Cel is very near zero that confirms

that in this domain the disintegration

is controled by the presence of the

internal radius a. From these different results we

can also draw the double effect of the

surface tension on the dispersion

diagram. When the surface tension

increases the growth rates increase

(see fig. 5 and 6) but the range of

wave numbers of the growing

perturbations decreases due to the Fig. 7.

more effective surface tension forces.

0.2-

0.0.

Wek and Cel for cylindrical

liquid sheets with Wea=0.58

161
3.3.2 With an azimithal component of velocity

Now the velocity U introduced in each non-dimensional i zed Weber numbers

(equations (4) and (8)) is replaced by the total velocity defined by :

U = / + V'

T V

2.0

(11)

As found by Dumouchel et al. [2]

the dominant perturbation may be

characterised by two orthogonal wave

lengths associated to two wave

numbers; k along the axis and n/a

along the azimuth. Then, the direction

of propagation of the perturbations

makes an angle with the axis of

symmetry. Over a range of Wea tgO^)

is compared in fig. 8 to tgO^)= VAJ.

a

U-4 m/s

V-3 m/s

For Wea>l the

longitudinal

dominant wave

Op=0) and

is

its

Fig. 8.

0.4 0.6 0.8 1.0

Wea

Direction of propagation of

the dominant wave for

cylindrical liquid sheets

with two components of

velocity.

characteristics are exactly the same

as those of the same liquid cylinder

with the axial velocity component U

only. For these cases an azimuthal component of velocity has no influence at

all. This comes from the fact that an azimuthal component of velocity tends to

reduce the surface tension forces due to the radius a. For Wea>l the

disintegration is mainly controled by the surface tension forces induced by

the radius a, therefore the dominant wave will stay longitudinal as an

azimuthal perturbation introduced by V would reduce the growth rate. However

for small Wea, the system tends to adopt the behaviour of a flat system and

therefore the presence of V will be regarded as an increase of the

destabilizing aerodynamical effects. The dominant wave is no longer

longitudinal O is different from zero) and when Wea tends toward zero ^

tends toward ^ the system behaves like the flat sheet with the velocity

component U

3.3.3 Discussion - Application

In practical situations two effects of the azimuthal component of

velocity can be therefore dissociated : first, a spin velocity modifies the

liquid geometry giving birth to a thinning conical system (as for a swirl

atomizer for instance) and secondly, when the internal radius has become large

enough, the spin becomes an effective component of velocity for the

disintegration increasing the wave number of the dominant wave. Dumouchel et

al [2] showed that a drop diameter d^ could be calculated from this

investigation :

T

Although it is generally not straight-forward such investigations may be

used as predictive tools. For instance, Dumouchel et al [2] adapted this

approach to the case of swirl atomizers that deal with conical liquid sheets

162
1.5

2r

0.8

0.4

30 40

ilJim)

50 8

U-IO m/s -

water/air — - Pg=l-29 kg.B - p^«=2.50 kg.B'

- p^«=5.00 kg.m"

high Injection, pressures

V

loM injection

pressures

± ±

0 0.2 0.4 0.6

We

0.8 1.0

Fig. 9. Calculated drop diameters d^

versus measured diameters

dqp for a set of swirl

atomizers (from ref. [2]).

Fig. 10. Influence of p on the drop

G

diameter for a flat system.

and an interesting agreement was found between measured and calculated drop

diameters (see fig. 9).

It is possible as well to bring out the influence of parameters involved

in the process of drop formation as for instance the influence of the external

pressure P on the drop diameter for a thinnig flat liquid sheet. Figure 10 amb

shows d odtained from equation (12) as a function of We for three values of

G

p . Let's call Wer the value of We at the break-up. When P

G ^ increases, p amb G

increases and fig. 4 shows that w increases then War decreases. Fig. 10

iamax

shows that an augmentation of P will make the drop diameter decrease if Wer

amb

is high (corresponding to low injection pressures) and will increase d if Wer

G

is low (corresponding to high injection pressures).

A similar behaviour has been experimentaly found by Tabata et al. [7] in

a slightly different case. It appears here that the explaination

of the opposite influence of the external pressure according to

the injection pressure may be linked to the fact that the liquid

system is thinning.

4 CONCLUSION

In this paper the instabilities of liquid systems (flat and cylindrical)

in an atmosphere at rest have been investigated by the use of linear theories.

The behaviour of flat systems was found to be totaly determined by a

set of non dimensional ized numbers and cylindrical systems were found to be

affected by the destabilising effects introduced by the internal radius.

However cylindrical systems behave as flat ones in many cases.

It is point out here that these approaches may be easily adapted to

different situations and used as predictive tools. For instance, calculated

mean drop diameters are presented for a set of swirl atomizers. A second

application shows the influence of the external pressure on the drops formed

by a thinning liquid sheet.

Most of the time, the results obtained are not accurate but the tendances

they bring out are of practical interest and may be used with confidence.

163
NOMENCLATURE

a internal radius of the cylindrical Iquid system

A, B constants

Cel propagation wave speed = w /k

r coefficients of the dispersion equation

d^ calculated drop diameters

dqp measured mean drop diameters

h thickness of the liquid systems

I (kr), K (kr) modified Bessel functions of order n

n n k ' longitudinal wave number

k k for the dominant wave

amax / 2 2 K total wave number =v k +(n/a)

T

n azimuthal number of perturbation

( r , 9 , x) cylindrical coordinates system

t time

U component of velocity along the axis Ox

V azimuthal component of velocity

(x , y) cartesian coordinates system

a amplitude of the functions t) angle between the axis of symmetry and the direction on

propagation of the perturbation

3^ angle between and the axis of symmetry

0 perturbation potential functions

T) displacement of the interfaces

p , p gas and liquid densities respectively

G L

(T surface tension

w pulsation

growth rate of the perturbation (imaginary part of w)

w maximum growth rate

iamax

REFERENCES

[1] G.D. Crapper, N. Dombrowski and G.D. Pyott, "Kelvin-Helmholtz

wave growth on cylindrical sheets", J.F.M. vol.68, p. 497-502 (1975)

[2] C. Dumouchel, M. Ledoux, M. I.G. Bloor, N. Dommbrowski and D. B.

Ingham, "The design of pressure swirl atomizers", 23 Symposium on

combustion Orleans July 1990

[3] W.W. Hagerty and J.F. Shea, "A study of the stability of the

plane fluid sheets", Journ. Appl. Mech. p 509-514 (Dec 1955)

[4] N. Dombrowski and P.O. Hooper "The effect of ambient density on drop

formation in sprays." Chem. Eng. Sci. , vol.17, p 291-305 (1962)

[5] H.B. Squire, "Investigation of the instability of a moving liquid

film." Brit. Jour, of Appl. Phys. , vol.4, p. 167-169 (1953)

[6] C. Guilie, "Etude experimentale et theorique de 1' injection

aerodynamique. " , Thesis ENS Mines de Paris Jan. 1987

[7] J. Martinon, " Instabilites d'une nappe liquide cylindrique en

presence de deux ecoulements gazeux. " Int. report ENSM Paris (1982)

[8] M. Tabata, M. Aral and H. Hiroyasu, "Effect of fuel viscosity and

surface tension on diesel spray drops." I CLASS 85 paper II/B/1

164
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 13

FLOW CHARACTERISTICS OF ATOMIZING LIQUID JET BY

SUPERSONIC FLOW AND RELATED SHOCK WAVE/

BOUNDARY LAYER INTERACTION

T. Arai, H. Sugiyama, N. Uno and T. Takahashi

Muroran Institute of Technology

Muroran, Hokkaido, Japan

ABSTRACT

Flow characteristics of atomizing liquid jet by supersonic flow and related shock

wave/boundary layer interaction were observed by flow visualization technique of

schlieren method with spark flash. Two types of liquid injector were used : a flat

plate injector and a backward step injector. The liquid was injected vertically into

supersonic flow of Mach number 2 from one orifice of 1 mm diameter at the flat plate

or at just behind backward step. In the case of a flat plate injector, the liquid was

atomized by boundary layer flow, and the liquid injection produced the bow shock

wave or pseudo-shock wave (multiple normal shock train). In the case of a backward

step injector, the liquid was not only atomized by boundary layer flow but also

supersonic main flow, since the low pressure part was formed outside the boundary

layer due to the existence of the expansion fan at the backward step. Therefore,

more fine particles will be produced by a backward step injector than a flat plate

injector. Interaction between the shock wave and the boundary layer leads to flow

separation and this might promote to mix of the fine liquid particles into the main

flow.

INTRODUCTION

There are a number of practical applications of transverse injection of a liquid

jet into a high-speed gas stream. Among a large variety of applications, fuel

injection for scramjets, ramjets, and afterburners, thrust vector control in rockets,

cooling sprays in turbines, and localized cooling on re-entry bodies exemplify this

important field of study. In all of these applications, cross-stream penetration and

the droplet size resulting from the break up of the liquid jet was items of primary

engineering interest. These general problems are still very important and were

investigated by many researchers*^"^*. The flowfield in the inlet/combustor isolator

section of scramjet engine is more complex due to the liquid injection/supersonic

flow interaction. The performance of the scramjet and ramjet engines strongly

depends on the flow characteristics. However, detailed information concerning with

the change of the flowfield due to the liquid injection into supersonic flow is not

available yet. Furthermore, the atomization mechanism of liquid fuel and the

following process of gas-liquid fuel mixing in supersonic flow is very important for

developing the scramjet engines.

The purpose of this paper is to clarify the liquid atomization process with

165
following gas-liquid mixing process in supersonic flow and shock wave/boundary

layer interaction due to the liquid injection.

EXPERIMENTAL APPARATUS AND PROCEDURE

The tests were performed in a 50 X 50 mm supersonic suction type wind tunnel at

Muroran Institute of Technology. Figure 1 shows the schematic diagram of the

experimental apparatus and measuring system. The supersonic flow trials were at

atmospheric ambient stagnation pressure and temperature. The symmetric Laval

nozzle is used so that the upper and lower wall boundary layers entering the test

section have identical properties. The side wall of the test section is provided with

an optical glass to measure the flow characteristics optically. Two symmetric Laval

nozzles were provided to obtain the nozzle exit Mach numbers of L8 and 2.0. In the

present study, two types of liquid injector were used as shown in Fig. 2. One was a

flat plate injector, and the other a backward step injector. Water and ethanol were

used as the injectant. The liquid was injected vertically into supersonic flow from

an orifice of 1 mm diameter. The driving force for the liquid injection was the

pressure difference between ambient pressure and local pressure at exit of the

orifice in the wind tunnel. The velocity of the liquid injecting was about 6~8 m/s.

The flow characteristics were investigated by using the flow visualizations

technique of schlieren and shadowgraph method with Q.S fis flash light. The wall

static pressure distributions were measured at upper wall of the square duct using

mercury manometers. The velocity profiles, in the case of no liquid injection, were

measured using a fiber optic dual-beam Laser Doppler Velocimeter (LDV) (TSI). The

TSI six jet atomizer using water was used to generate seed particles with a

polydispersed size distribution with mean diameter of about 2 fim. It was confirmed

by the preliminary experiments that the incoming supersonic flow showed a high

degree of two-dimensionality. All measurements were made on the midplane of the

wind tunnel except for a limited series of spanwise traverses made to check the

degree of two-dimensionality of the flow.

RESULTS AND DISCUSSIONS

Velocity distribution of main supersonic flow

Figure 3 shows the velocity distributions of the main supersonic flow. The top

view is the case of flat plate injector type and bottom view the case of the

backward step injector type. The thickness of the boundary layer just ahead of the

injector for the flat plate type was about 5 mm and the boundary layer thickness

just ahead of the step was also about 5 mm. It was confirmed by the measurements of

the velocity distribution in the boundary layer that it was turbulent.

Flow visualization of atomizing liquid and flow pattern

Flat plate injector. Figure 4 shows the liquid injected vertically into

supersonic flow of Mach number 2.0 with following mixing process and related flow

pattern. The top and bottom views show the water and ethanol injection,

respectively. The vertical line in the photos indicates the marker of the injecting

location. It is seen from the top view of Fig. 3 that the thickness of the fully

developed turbulent boundary layer was about 5 mm. The top and bottom photos of

Fig. 4 show the bow shock produced by the liquid injection. The bow shock reflects

at the upper and lower walls and propagates downstream. It is seen from the top and

bottom views that the liquid was atomized by the boundary layer flow so that the

liquid particles scarcely diffused into the main supersonic flow.

Figure 5 shows the flow patterns of Mach number L8. The top view of water

injection case shows that the pseudo-shock wave^®' (multiple shock train^®^) was

166
formed by the liquid injection/shock wave/boundary layer interaction. In the case

of ethanol injection, the pseudo-shock wave was formed far downstream, which was

found by the wall static distribution. The pseudo-shock wave causes the loss of

pressure and oscillation phenomena of main flow so that the generation of the

pseudo-shock wave might have bad influence on the supersonic mixing and

supersonic combustion mode. However, as seen from the top view of Fig. 5, the

separation region under the first shock wave of the pseudo-shock wave contributes

to the diffusion of the liquid particles.

Backward step iniector. Figure 6 shows the schematic view of the flow field of

the main flow Mach number M=2.0. The orifice was located 5.4 mm from the backward

step. Figure 7 shows the schlieren photographs of air flow patterns and liquid

injected by backward step injector. The left and the right side photos are taken,

respectively, with a horizontal knife edge and a vertical one. From the top view of

the left side, it is seen that the turbulent boundary layer was separated at a

backward step and formed a free shear layer, which went through a reattachment

process. The average reattachment line is located approximately 2~2.5 step heights

from the step. The expansion fan at the backward step and reattachment shock wave

are also observed. The middle and bottom photos show the flowfield in the cases of

water injection and ethanol injection, respectively. The weak bow shock wave

produced by the liquid injection can be seen in the expansion fan. The strength of

the reattachment shock wave with liquid injection grew weaker than the case of no

liquid injection. Figure 8 shows the behavior of the liquid atomization by backward

step injector, which was taken by backlight photograph method. As compared

between the spatial position of the atomizing liquid in Fig. 8, the boundary layer

thickness and station of the bow shock wave in top view of Fig. 7, it seems that the

liquid was not only atomized by boundary layer flow but also the supersonic

mainflow. Therefore, liquid particles produced by a backward step injector may be

more fine than that of a flat plate injector because the liquid is atomized by higher

velocity gas flow.

Wall static pressure distributions

Figure 9 shows the time mean wall static pressure distributions on the upper

wall along the duct for the case of the flat plate injector. The symbols , O and

A indicate the conditions of no liquid injection, water injection and ethanol

injection, respectively, and Pa is the atmospheric pressure. The abscissa x means

the downstream distance from the injection hole. The time mean wall static pressure

increased abruptly from x=350 mm for all three cases. This indicates generation of

pseudo-shock waves (shock train). Figure 10 shows the time mean wall static

pressure distributions at the upper wall along the duct for the case of the backward

step injector. The symbols in Fig. 10 have the same meaning as before. The high

pressure region are not observed in all cases in Fig. 10 and the pseudo-shock wave

was not generated. The liquid injection has little influence on the time mean wall

static pressure distribution for the two types of injector, in the present experiments

using the mercury manometers system.

Effect of pre-injection on liquid particles mixing into main flow

Figure 11 shows the schlieren photograph of the liquid (water) atomization from

the flat plate injector with two holes located 15 mm apart. The liquid injected from

the second hole has a higher penetration height than that from the first injection

one (pre-injection). From the above experimental result, it may be said that the

injection of liquid just upstream to the main injection controls the turbulent

boundary layer and promotes the mixing between the main flow and the liquid

particles.

167
Penetration height of liquid jet for the case of flat plate type iniector

The penetration height was determined by the observation of the amount of

light that scattered when a beam of light passes through a suspension of droplets.

Figure 12 shows the variations of the penetration height for the case of water

injection and ethanol injection. It is considered that the penetration height

corresponds to the mixing layer thickness. The symbols 0 and A mean the water

and ethanol injection for the case of Md=2.0 (Md : design Mach number of the Laval

nozzle, free stream Mach number M=1.8), respectively. Also, the symbols O and A

mean the water and ethanol injection for the case of Md=2.2 (free stream Mach

number M=2.0), respectively. The mixing layer for the case of Md=2.0 was thicker

than that of Md=2.2. In the case of the water injection for Md=2.0, the supersonic

mixing was not happened by the generation of the pseudo shock wave in the duct

(as it was seen in Fig. 5). It can be seen that the mixing layer thickness by ethanol

injection was larger than by water injection case. The reason for this result has not

been clarified.

CONCLUSIONS

The liquid atomization process with following gas-liquid mixing process in

supersonic flow and shock wave/boundary layer interaction by the liquid injection

were investigated by the schlieren photograph and shadowgraph method. It was

cleared that the liquid injection in supersonic flow produces the bow shock wave

and/or pseudo-shock wave (multiple shock train), resulting supersonic flows more

complicated. The behavior of the liquid atomization process and mixing process are

strongly influenced by the shock wave/boundary interaction. It is shown that the

production of more fine liquid particles and the promotion of the gas liquid mixing

can be expected if the reflection of the shock wave and the expansion fan at the

backward step would be effectively used. Therefore, the change of flow pattern,

which includes the generation of the shock wave, the development of the boundary

layer and the separation of the boundary layer, should be used effectively for the

liquid atomization technique by the supersonic flow.

REFERENCES

L Schetz, J., Kush, E. and Joshi, P., Wave Phenomena in Liquid Jet Breakup in a

Supersonic Crossflow, AIAA Journal, Vol.18, No. 7, pp. 774-778 (1980).

2. Reichenbach, R. and Horn, K., Investigation of Injectant Properties on Jet

Penetration in a Supersonic Stream, AIAA Journal, Vol.9, No. 3, pp. 469-472

(1971).

3. Schetz, J. and Padhye, A., Penetration and Breakup of Liquids in Supersonic

Airstreams, AIAA Journal, Vol.15, No. 10, pp. 1385-1390 (1977).

4. Less, D. and Schetz, J., Transient Behavior of Liquid Jets Injected Normal to a

High-Velocity Gas Stream, AIAA Journal, Vol.24, No. 12, pp.1976-1986 (1986).

5. Nejad, A. and Schetz, J., Effects of Properties and Location in the Plume on

Droplet Diameter for Injection in a Supersonic Stream, AIAA Journal, Vol.21,

No. 7, pp. 956-961 (1983).

6. Heister, S., Nguyen, T. and Karagozian, A., Modeling of Liquid Jets Injected

Transversely into a Supersonic Crossflow, AIAA Journal, Vol.27, No. 12,

pp. 1727-1734 (1989).

7. Thomas, R. and Schetz, J., Distributions Across the Plume of Transverse Liquid

and Slurry Jets in Supersonic Airflow, AIAA Journal, Vol.23, No. 12, pp. 1892-1901

(1985).

8. Aral, T., Sugiyama, H., Abe, H., Takahashi, T. and Onodera, 0., Internal

Structure of Pseudo-Shock Waves in a Square Duct, AIP CONFERENCE

168
PROCEEDINGS 208 (Current topics in shock waves, 17th International Symposium

on Shock Waves and Shock Tubes, Bethlehem, U.S.A.. 1989). pp.850-855 (1989).

Carroll, B. and Dutton, J., Turbulence Phenomena in a Multiple Normal Shock

Wave/Turbulent Boundary Layer Interaction, AIAA Paper No. 90-1455 (1990).

1 :Schlieren Systan

2:Cainera

3: Pressure Transducer

4 :D.C. Amplifier

5 : Syncliroscope

6: Mercury Manometer

7: Shock Location

Controller

8: Valve

9: Vacuum Tank

Fig.l Schematic diagram of experimental apparatus

Supersonic flow^

<t)l

4)6

Supersonic flow

5^

(ill

<i>3

5.4

(a) flat plate injector (b) backward step injecter

Fig. 2 Schematic view of injectors

'///// '/////J7i ///////

1 .0 —Vo

o

oo

o

0

oo

o

o

o

1 .0 u/u„ o

o

oo

o

o

o

oo

o

o

15

/777y77T777T77777730

//////////

(a) flat plate injector

1 .0

<- u/u„

b) backward step injector

Fig. 3 Velocity distributions of main supersonic flows,

Ujomeans the velocity of free stream flow outside

boundary layer and corresponds Mach number

M=2.0.

169
- > injection- (^f^er

Fig. 4 Schlieren photographs of liquid injected

vertically into supersonic flow of M=2.0

\ \ \ — ^ K

^^

1

^"X \ \ \ \ \ \ \ V

Fig. 5 Schlieren photographs of liquid injected

vertically into supersonic flow of M=1.8

5.4

K ^ In^ ector

A 1 ^

-5 0 5 10 15 ^

Fig. 6 Schematic view of flow field at backward step

injector

170
/ Y y y y^ y y y / y y y / / x x x x x x x x x

(a) without injection (d) without injection ^^^y yyyyyy/yyyyyy / yy/yyyyyyyj

(b) w^ti: ^ injection y y X y X , s

. '^ injection' y y y ^ * (e; water ^^^yyyyyyyyyyyyj y yyy y y y y y y y y y y y_

TJ'. //////

(c) ethanol ' i"3^=tion ^ ^ . x

,^^,^^4^^^ t injection - - - '

Fig. 7 Schlieren photographs of air flow pattern and

liquid injected by backward step injector at

M=2.0 ; left side photos are taken with a

horizontal knife edge and right side photos with

vertical one.

Sb^^'*., ^

^ ^ t injection (water) ^ ^

Fig. 8 Liquid atomization bybackward step injector

taken by backlight

method at M=2.0

0.7

0.6

<2 0-5

&4 0.4

0.3

0.2

0.1

0

0: Water

A: Ethanol

without Injection

a 0

QgoQoBeOSeooD

200 400 600

X mm

Fig. 9 Wall static pressure

distributions for flat

plate injector at M=2.0

800

171
0.6

(0 0.5

0.4

0.3

0.2

0.1

O without injection

A water

ethanol

P:wall static pressure

Pa: atmospheric pressure

DO O —'—'

'—1—I—I I

i 200 400 600

X mm

800

Fig. 10 Wall static pressure distributions for backward

step injector at M=2.0

-f ^ injection^ (water) \ \ \

Fig. 11 Schlieren photograph of double injections into

supersonic flow for flat plate injector at M=2.0

30

20

10

X

A

o

6 0

X X

Q 0 •

o O o

M

' Inj ector

• • Md=2.0 , Water

a: Md=2.0 , Ethanol

O : Md=2 . 2 , Water

A: Md=2.2 . Ethanol

100

X

200 mm

Fig. 12 Penetration heights of liquid jet for flat plate

injector

172
ICLASS-91 Gaithersburg.MD. U.S.A. July 1991 Paper 14

SHOCK WAVE INTERACTION WITH LIQUID

LIQUID ATOMIZATION AND ANALYSIS SCHLIEREN, LIGHT SHEET AND

HOLOGRAPHIC RECORDS

N.M. Fargere

Centre d' Etudes du Bouchet

Vert-Le-Petit, France

ABSTRACT

Though subaquatic explosions have been the subject of many studies, the interaction of the subaquatic shock

wave with the liquid air interface is not well known, particularly in its earliest effects. In particulary, microscopic

and precocious events occur during the acceleration of the free siuface of the liquid.

We used several techniques (fast camera, image convector, schUeren and ombroscopie records...) to record the

water surface motion and the main phenomena above it. For the description of the first phenomena, holographic

techniques appeared very instable, and we observed the free surface movements using light sheet for the latter events, like plume.

GENERAL DESCRIPTION

The experimental set up is a test tank (1 x 1 x 0,5 m3 P.M.M.A. of optical quality) filled tap water [1].

A 215 mg PETN detonator is ignited at several immersion dephts, basically 0,25 m, the depth for which all the main characteristics of the problem have been studied. This in order to study the influence of the shock

intensity on the surface effects. The detonator is a very simple way to generate a spherical shock wave far enough

from the detonation point the profile of which is checked by pressure transducers and compared with the theorical

calculations [2]. The correlation obtained allows us to calculate the rough parameters of similitude in Cook

formula

.

A typical case of the shock front properties impacting against the water-air interface is a peak pressure lower

than 100 bars, usually 50 bars, a shock velocity in water : 1500 m/s, a maximum free surface velocity : 7 m/s.

Several optical methods are used

:

1) Direct camera records.

2) Schlieren and ombroscopie cinematography using an argon laser modulated light (10 |is duration, 900-1000

Hz frequency) and a drum camera.

3) Holographic technique using a pulsed laser. 4) Light sheet laser techniques.

THE OBSERVED MACROSCOPIC PHENOMENA

The observed effects at the water surface are classified in two main categories : macroscopic and microscopic.

During its detonation in a liquid medium, the solid explosive is rapidly transformed into gaseous products at high temperature and pressure. The so created bubble expands rapidly and its inner pressure decreases down to a

pressure slightly lower than the hydrostatic equilibrium corresponding pressure. The bubble then shrinks until the

limit of gas compressibility is reached, the movement inverto and a secondary shock wave is generated.

This phenomenon occurs after the second shock and so on. Just after detonation, the first shock wave separates

from the bubble and spreads out in surrounding water at a velocity which falls rapidly to a slightly supersonic

one. Its peak pressure decreases rapidly, a great part of the shock energy being dissipated during the propagation in

173
water. Coming up at the surface, it transmits a wave in the air and a rarefaction wave goes back towards the

detonation source. The above parameters highly depend on the depth and on the charge weight It is known that for

small charges, at relatively important depth (ko/do «1, Ro radius of charge, do depth), the reflection can be

considered as total and the shock transmitted into the air is weak [(3), (4), (5)]. In the case we are interested in (Ro/do « 1), the first macroscopic effect is bound to the arrival of the shock

wave itself. It is the shock black ring which spreads on the water surface. Then we observe the dome, mass of

water accelerated by initial shock and sub surface movements. Its initial speed is very close to the speed of matter

bound to the shock. The maximum heigth of the water above the surface is reached when the forces of air friction

and gravity balance those of upward motion. In our case, the secondary dome, bound to the going up of secondary

shock wave and to the bubble second expansion, adds to the principal dome when it's going down. There is then a break and the beginning of the water plume. It is noted that the coming out of the column (plume) is closely

linked to the depth of the detonation and to the used explosives characteristics (mass...). In a general way the

column is the result of either breaking of dome or coming up in the air of detonation products when these are still energetic (first period (s) of pulsation interacting with the surface) [(6), (7), (8)]. These phenomena have been

observed by ombroscopy and schlieren techniques on high speed camera. The microscopic effects show the

coming up at the surface of small jets (which have been observed by means of an image convector tube) and a

suspension of mistlike droplets. A first approach of jets formation mechanism has been made thanks to a rapid camera and ombroscopy records.

The description of the microscopic effects and their origins is the aim of this paper. The results described are the one obtained with the laser means of recording. Lasers are used as energy sources (light sources) as well as

specific precision means. They offer several technical advantages : three dimensional image (holography), high

energy (light sheet), great precision of synchronisation and a short exposure time. We must underline the great

possibilities in magnification and the excellent definition.

CAVITATION

Liquids can more or less withstand tensions. Untreated, their continuity happens to be broken up by formation

of cavities. This cavitation tendency is especially increased by presence of impurities. In the case of sea water or

industrial water, the ambient pressure under which the cavitation increases indefinitely is very close to saturated

vapor pressure. This approximation does not take into account the size of impurities seeding. We admit that if the

pressure nearby a bubble exceeds saturated vapor tension, the cavity disappears. There is an implosion with

creation of a jet characterized by a high speed and the emission of a compression wave [(9), (10), (1 1)].

The formation of bubbles, their gradual growth are also influenced by water physico-chemical properties

(inclusions, surface tensions, specific viscosity....). The modeling of the phenomenon is not easy, especially in

presence of a free surface.

After their formation, the bubbles of an initial diameter of some micrometers can be submitted to pulsation,

according to the flow where they happen to be [12]. This pulsed movement is characteritic and can lead to maximal diameters of several tenth millimeters associated to a life time from some hundred of nanosecondes to several milliseconds, limited by a relatively violent implosion [13].

Assuming that the cavitating pressure is saturated vapor pressure we decided to apply first the Cole

formulations [14]. The experimental cavitation zone was not compliant with the one we numerically found. We

did not take into account the bubble erosive action on the expansion front Using Dubesset et Lavergne [15]

method and a trick in mathematics we found to introduce time, we obtained the specific shape of the cavitation zone at different instants and depths. An analysis of the numerical results (minimum and maximum depth) shows

a relative agreement with the experimental results, although by excess, they are very closed to Kedrinskii's [16].

The experimental cavitation zone has been deduced from under water surface holographic records.(Fig.l)

INSTABILITIES AND BREAK UP OF LAMINAR JETS

Basically, Plateau [17] attributed the main mechanism to the interfacial tension. He showed that if there is an

axisymetric disturbance in the jet, it reduces the free energy of the jet if its wave length is greater than the

original undisturbed jet circonference. The system attempts to minimize the free surface energy, aiming at a minimum surface area per volume unit, reaching then the break up in spherical drops. Rayleigh [18] gave the

equations and correlated the amplitude growth rate with the perturbation wave length. He found that if all axisymetric disturbances - which are characterized by a wave length greater than the circumference of the

undisturbed jet - are unstable, only one has got an amplitude growing faster than the others, dominating the break

up and fixing the jet's length and drop's size.

174
The first studies are based on the assumption that there is a pressure initiated perturbation originated at the

jet's surface which equation should be

:

(1)

5=5oexp(qt+ikz)

Weber [19] also solved the Navier-Stokes equation for a liquid jet. He showed that the perturbation

characterized by the wave number k* rises rapidly and makes the jet break up into drops. Tomotika [20] solved

the equation in his study of jets in 1966. The dimensionless growth rate and the wave number of the disturbances

that grows most rapidly are given by, respectively [21].

Q* = q* 'y^27^ (2) and x* = ak* (3)

Assuming that break up occurs when the disturbance grows up to be comparable with the jet radius, break up

time T is given by. _

T- L" (a/5o)

Assuming that jet's radius and velocity are quasi constant we can say that break up length is given by T x u.

L = Ln

l5o

The liquid in a column of one special wave length turns into one drop of corresponding volume, the diameter

of drops formed from jets of those type is :

3\l/3

(6)

HOLOGRAPHY

General Principle

General Principle of classical photography is based on the record of light, intensity and wave length.

Moreover holography calls for light phase, which allows to integrate for recording, the relative position of

different parts of the object. The light diffused by the objects is recorded on a very high resolution power plate on which is superposed a reference wave. The complex diffraction so formed net diffracts light when placed in

reference wave, then restoring a real image and a virtual one of the object.

Microholography

Because of the phenomenon low rate of darkening (< 10%) we used GABOR mounting. The separate reference

assembling regains an interest for recording the phenomenon after 2500 |xs. We then got rid of field observations

(astigmatism, coma...) but not of the spherical one [22].

Restitution is obtained by the analysis of the real image recovered on a TV set. Reconstruction is made semiautomatically by means of an image analyser [(23), (24)].

A synchro-optic cell is adjusted to obtain a double exposition, allowing to record of the object in two

consecutive positions, and to calculate the speed calculation.

We used a ruby oscillator laser (50 mJ, 15 ns) for recording with an optic system [Div. Doublet (20, 30 mm),

Conv, lens (800 mm)]. For reconstruction we used an unfocal system (65 mm) and an objective. Taking into

account the errors in the different positionnings of detonator and hologram (recording and reconstruction) the

relative error is never more than 10 % in depth, and 5% in the observing plane.

Hologram assessment and results for jets

General remarks. It is difficult to assert the occurence of such and such mechanism. We will see that the

observations allow the appraisal of

:

1) the periodic instabilities along the jet,

2) the forces of tension which depend upon the velocity gradient in the jet.

175
These two mechanisms take place separately, the instabilities being able to grow and to set only if the whole jet

exists. However, the conjugate mechanisms are sometimes simultaneously observed. The velocity gradient works

in terms of stretching and shearing, the break up is then irregular.

Regular drops generated by jets are mostly relevant to Rayleigh's instabilities.

Specific results assessments and exploitation gave us the following set of average results, where d and T are measured and calculated directly, d' an T are deduced from calculations.

n 1 2 3 4 5 6 7

d{[m) 194 279 320 410 450 200 260

d'(um) 198 250 240 310 310 140 270

% 2 11 28 27 36 33 4

T(us) 190 790 528 434 368 475 2570

r (jis) 405 353 670 360 460 439 2500

A(%) 66 72 25 15 24 8 3

The variation of relative difference between calculated and measured diameters is in the range of 2 to 36 %.

If we exclude two results (n = 1,2) that we assume to correspond to a different break up mechanism (striction),

the times T and T are in agreement in the range of 3 to 25%.

Error can be imputed to the specific experimental measurement error (about 8%) and to the

fact that Tomotika theory is relevant to very constant laminar jets (created in reproductible conditions)

characterized by constant speed. If the following set shows that we realize one condition (one jet drops production)

speed is not totally constant and we know its importance in the rupture time calculation.

n 1 2 3 4 5 6 7 8 9 10 11 12 13

dOim) 89 74 89 81 71 77 89 77 71 71 89 74 75

LIGHT SHEET EXPERIMENTS

Light sheet laser assembling - Dispersion laws

The light sheet laser mounting is very closed to the microholography one. The main difference is due to the

specific optics used. We calculated the numerous parameters aiming at a maximum energy in the conic sheet

created by the quartz bar (the Neodym laser used was characterized by 1,15 J (1050 nm) and 250 nrJ (532 nm) for a 25 |is duration. The beam diameter is 7 mm).

Exploitation was built over the use of an image analyser. For several depths and times we analysed the photos.

We were then able to appraise MMD (Mass Median Diameter) and NMD (numerical Median Diameter) for each

trial.

The theorical laws are

:

Drop's volume . V = ^ 7t R^ = J- Ji <1>^ (7)

3 6

Drop's mass : M = Vd (8)

dM - MT p^n/ In (!) - In MMD?

Mass variation : , . , ^ r-— .

• ^''V \ ^rr ^^ „ \ W # { In c f . Tin . (|) \ y2 . In O /

(in (t) - In MMDf . Mi

(in af = -i (10)

176
Numerical variation

NT

# (In af . V2TT . 0

exp - In (> - In NMD

V2 Incj

(lnaf = X ( In <t) - In NMDf. Ni

IN

(11)

(12)

MMD and NMD relation : Ln MMD = Ln NMD + 3 (In a)^ (13)

Results

Comparing curves of theorical dispersion and tlie experimental one we found that if we are very closed as far as we compared NMD's results, we weren't not so acute in MMD. Knowing that the NMD's laws are the more

followed, we used it with the aim of comparing the theoric MMD and the experimental one (Fig.2, Fig.3)

.

Depth

mm

Time

ms

MMD

measured

mm

NMD

neasured

mm

LnaMMD LnoNMDc

MMD

alculated

mm

(1)

250 26 1.6 0.7 0.4982 0.5239 1.59 0.33

250 24 1.5 0.5 0.5448 0.5338 1.17 22

250 18 1.2 0.6 0.5055 0.4929 1.24 3.6

250 9 1.6 0.6 0.5030 0.5298 1.39 12.9

225 4.837 1.5 0.5 0.5245 0.5248 1.14 24

175 4.837 1.5 0.6 0.5014 0.5698 1.59 5.9

160 4.837 1.5 0.6 0.4614 0.5363 1.42 5.2

(1) A MMD/MMD measured %.

We can notice that the acciuxacy is somewhat variable, but generally closed to 10 % and less. We encountered some difficulties with the beam's width and its influence on the light transmission.

SCHLIEREN AND OMBROSCOPY

We used an Argon laser, an acousto-optic modulator and a Callus, drum camera to realize the first schlieren

views (interframe time 1,1 ms, impulse time 10 \is, rotating frequency 396 Hz). They enabled us to identify with

certainty the jets we already saw with the image tube convector. But wc had some difficulties with a too weak

light intensity so we decided to use ombroscopy with a very close setting. We realized different trials and we were able to quantify the peales and

dome behavior (position and velocity). We identified the jets instabilities but the accuracy was not so good, but

the instabilities under the surface were identified too

.

177
CONCLUSIONS

The free surface motion of water-air interface accelerated by the impact of a shock impulse generated by an

explosive has been observed by means of holographic technique.

The results of the holograms analysis allowed us to confirm and to clarify the role of cavitation phenomena

which :

- is identified as soon as the reflection of the shock wave takes place at the interface,

- creates micro jets in the air, the distribution of which is correlated with the propagation of the cavitation

phenomenon,

- is in relatively good agreement with a numerical model taking into account the erosion of he

expansion front

Holography appears to be a very suitable technique to prove the microjets existence and to visualize the jets

instabilities.

The microjets deformation and rupture seem to be consequence of:

- periodic instabilities along the jet,

- forces of tension which depends upon the velocity gradient in the jet.

Dispersion laws are followed somewhat accurately. The light sheet results are very interesting.

Shlieren and ombroscopic records enabled us to characterize the phenomena occuring at the surface between

the very early instants (bubbUng and peacks birth) and the whole effects (plume).

Shock wave interacting with a liquid air interface is able to generate cavitation and small bubbles, the

implosion of which is responsible of microjets and small droplets production in air. New studies of the

implosion of the bubbles are needed to contribute to the behavior of cavitated liquid supporting chemical reactions

[25].

NOMENCLATURE

a = radius ofjet (m),

d = diameter of drop (m),

X = dimensionless wave number.

k = wave number (m"l),

L = break up length of jet (m),

M = drop Mass

MT = Total measured drop's mass N = Drop Number

NT = Total measured drop's number

Q = dimensionless growth rate, q = growth rate of disturbance (s'^)

R = Drop radius

T = break up time (s),

t = time (s), z = axial coordinate (m),

V : Drop Volume

8 = disturbance on jet (m),

((» =Drop diameter

p = density of liquid (kg/m3),

a = interfacial tension (N/m),

178
BIBLIOGRAPHY

1- Jonhson J.N.: Calculated shock pressures in the aquarium test - Los Alamos National Laboratory NM 87

545. 1982 (US Dpt. Of Energy).

2- Cook M.A, Keyes T.T, Ursenbach W.O.: "Measurements of Detonation Shock and impact pressures"

University of Utah - Salt Lake City.

3- Mader C.L."Compressible Numerical Calculations of Underwater explosion". Los Alamos Scientific

Laboratory University of California (1971).

4- Snay H.Cand Rosenbaum S.H.: "Shock wave parameters in Fresh Water for pressures up to 95 kbars"

Naval report 2383 US Naval Ordnance Laboratory - White Oak, Maryland (1952).

5- Wang Xu, De Yang Yu, Yun-Xiang Xu, Xiuny-Fei Zeng : "Effects of charge diameter on detonation,

Pressure Measured by aquarium technique". Lanzhou Modem Chemistry Research Institute. Chine.

6- Sakurai A and Pinkson J.M, US Army Engineering : "Water ways Experiments Station" Technical Report n°

1-771 (1967).

7- Zaslavskii BT : ZH Prick Mekls - Telek - Fig 4.50 (1963).

8- Holt M."Underwater explosions" - Ann Rev. Ruid - Mech § 187-214 (1977).

9- Flynn H.G.: "Physics of Acoustic Cavitation in Liquids". Physical Acoustics. Col. Wamen, P. Manon, Vol

IB (1964).

10- Weston D.E : "Underwater Explosions as Acoustic Sources" A.R.L. Teddington, Middlesex U.K.

11- Wentzell R.A : "Cavitation phenomena due to Shock Pulses Reflected from the Sea Surface". University J. Waterloo, Ontario, Canada.

12- Bangah et Sarh : "Cavitation induite par une onde de detente dans un fluide au repos". Rapport interm^diaire.

Contrat DRET 81-500. (1981).

13- Giovanneschi P.: "Contribution h I'^tude de la cavitation h buUes isol6es, initios par un rayonnement laser

de grande intensity". Universite d'Aix-Marseille II. UM CNRS n° 34 - Thfese (1987).

14- Cole R.H. "Underwater Explosions" 1949.

15- Dubesset et Lavergne "Calcul de cavitation due aux explosions sous-marines h faible profondeur". Acustica

Vol. 20 (1968).

16- Kedrinskii V.K.: "Surface Effects from an Underxater Explosion". Zhumal Prikladnoi Mekkaniki i Tecknicheskoi Fiziki n° 4 pp.67-87 (1978).

17- Plateau M.T.- Phil Mag. 12 P. 286 (1856).

18- Raileigh J.W., Strutt, Baron - Proc. London Math. Soc. 10, p.4 (1878).

19- Weber C- Zeit. Angew Math. Mech : "Disintegration of Hquid jet" Math Mech - 11/136-54 (1931).

20- Tomotika S."The Instability of CyUndrical Column of a perfect liquid surrounded by another perfect fluid"

Proc. Physico Math Soc. Japon 18.550.62 (1936).

21- Abott W. - "Analysis of the Instability of laminar, Newtonien liquid Jets in Air" - The Louisiana State

Agricultural and Mechanical Col. PHD New Orleans (Louisiane) (1982),

22- Royer H.: "La microholographie et ses applications" Institut Franco-AUemand de Recherche de Saint-Louis -

France. Rapport interne.

23- Farg^re N : "Contribution k I'automatisation du d6veloppement d'hologrammes de GABOR". GAMS

COFERA, "Joum^es d'6tude sur les A6rosols" (12/87).

24- Farg^re N : "Ddpouillement semi-automatique d'hologrammes de GABOR et programme d6grad6". Data

processing on Interpretation in Holography 2 nd Franco-German Congress on AppHcations of HolographySaint-Louis - France (1988).

25- Suslick K. "Les effels chimiques des ultrasons" - Pour la science n° 138, p. 89 h 94 (1989).

179
Fig.1 Experimental and Theorical Cavitation Zone

Zone of cavitation = f (Time)

rrrr,

E

E

-40-

-«0- a

0)

-80-

-100-

-120-

mm, Experimental zone I

Theorical zone

Holographic measure

100 200 SCO 400 Time (microseconds)soo

Fig. 2 Drop's number versus Diameter (250 mm)

eQ

........ - - [j|an,g,8f (mm)

Fig. 3 Drop's mass versus Diameter (250mm)

600

Diameter (mm)

180
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 15

DYNAMIC SIMILARITY BETWEEN LIQUID SHEETS

AND HARD SPRING SYSTEMS

A. Mansour and N. Chigier

Department of Mechanical Engineering

Carnegie Mellon University

Pittsburgh, PA, U.S.A.

ABSTRACT

Detailed measurements of the frequency of oscillation of a liquid sheet issuing from a 2D air

assisted nozzle have been made. The measured vibrational frequencies were then correlated with

the resulting spray angle. It was shown that the liquid sheet oscillations are dynamically similar to

that of hard spring systems. The effect of introducing air in the nozzle simulates the effect of

forced vibrations on the nozzle jaws. Thus, for each air flow rate, there is a specific vibration

frequency for the nozzle. The frequency of these vibrations is proportional to the air velocity. As

the liquid sheet natural frequency approaches that of the nozzle, resonance is established. At

resonance, the maximum spray angle is achieved.

INTRODUCTION

An experimental study is being conducted on a 2D twin fluid atomizer, similar to that of

Reference 1. A liquid sheet emerges from a central slit with a high aspect ratio. High velocity air

impinges on both sides of the liquid sheet at the nozzle exit. Microphotography was used to

determine the global structure of the sprays, and a light attentuation technique was employed to

measure the frequencies of oscillations.

FREQUENCY MEASUREMENT TECHNIQUE

The experimental technique utilizes the attentuation of a collimated laser beam passing through

the liquid sheet in the transverse direction (i.e., parallel to the nozzle slit). The optical system is

shown in Fig. 1, The beam was positioned 2 mm upstream of the breakup region. This method

ensures the sampling of the most dominant frequency, that which causes the breakup of the liquid

sheet. Figure 2 shows the mean breakup length (i.e., the distance from the nozzle exit to the

disintegration point) of the liquid sheet, as determined from many photographs. Numerous tests

were performed by translation of the laser beam along the intact length. The bandwidth of

oscillation frequencies was reduced as the laser beam was translated in the downstream direction

along the intact length. Spurious modes of oscillation disappear as a dominant frequency emerges

which is sustained by the liquid sheet. Even if the breakup length recedes to the nozzle exit, this

method still proves to be reliable.

Two types of waveform are possible depending on the location of the laser beam with respect

to the Z coordinate. Figure 3 shows photographs of the waveform taken from the oscilloscope

181
1) He-Ne Laser

2) Collimating Lens

3) Nozzle

4) Optical Filter 5) Focusing Lens

6) Photodiode

7) Current to Voltage Converter

8) A-DI rConverter

9) IBM XT

Fig. 1 Optical System

E

Q. 3

Air Pressure

101 574 kPa

102 015 kPa

102 74 kPa

104.31 2 kPa m 106 22kPa

1 1 1,667 kPa

20

Fig. 2

Liquid Velocity (m/s) Mean Breakup Length as a function of Liquid

Velocity and Air Supply Pressure to Atomizer

a) Ziaser = A (amplitude of oscillation) b) Ziaser = 0

Fig. 3 Typical Wave Forms (Pq = 101.823 kPa, muquid = 15.77 g/s)

screen. When Ziaser = A (A being the amplitude of oscillation at the measuring location) the

waveform is shown in Fig. 3a. Figure 3b shows the waveform when Ziaser = 0. Particular

attention has been paid to the Z location of the laser beam because, as shown in Fig. 3, sampling at Z=0 would lead to a vibrational frequency that is double the actual frequency of the liquid sheet.

Figure 4 shows the actual waveform sampled when the breakup length recedes to the nozzle

exit. In this case the signal is produced by ligament shedding. Figure 5 shows the power spectral

densities of the liquid sheet vibration. Figures 5a through 5c correspond to Figs. 3a, 3b and 4,

respectively.

182
Fig. 4 Typical Wave Form when Breakup Length Recedes to the Nozzle Exit

Ziaser = A; Pq = 104.312 kPa; miiquid = 15.77 g/s

120-

110-

100-

90- « 80-

t 50-1

40-

30-

20-

10-

0-

Frequency (Hz)

1013 Frequency (Hz)

a) Ziaser = A; Pq = 101.823 kPa;

miiquid = 15.77 g/s => f = 508 Hz

b) Ziaser = 0; Pq = 101.823 kPa;

miiquid = 15.77 g/s => f = 1013 Hz

1034 Frequency (Hz)

c) Ziaser = 0; Pq = 101.823 kPa;

miiquid =15.77 g/s =^ f= 1084 Hz

Fig. 5 Power Spectral Density

RESULTS AND DISCUSSION

Figure 6 shows the vibration frequency of the liquid sheet as a function of liquid exit velocity

for different air supply pressures to the atomizer. Figure 6 is divided into three distinct

subregions. In regions A and B the frequency of the liquid sheet is highly stable. A sampling rate

of 20000 Hz with a time duration of 0. 1 sec was sufficient to insure perfect repeatability. In these

two regions the bandwidth of oscillation frequencies is very narrow and a perfect spike similar to

183
N

X

o

c

CDa5000

4000

3000

Air Pressure

2000 -

1000 -

o 102.015 kPa

« 102.704 kPa

103.565 kPa

104.312 kPa

108.22 kPa

1 1 1 .667 kPa

20

Liquid Exit Velocity (m/s)

Fig. 6 Vibration Frequency of Liquid Sheet as a Function of Liquid Exit

Velocity for 6 Different Air Supply Pressures to Atomizer

that of Fig. 5 is recorded. The clarity of the signal displayed in these regions is due to the fact that

the instrument is very sensitive to wave amplitudes: the greater the wave amplitude, the better the

signal.

In region C, secondary modes of oscillation appear. The transition from region B to region C

is accompanied by a widening of the bandwidth of oscillations and a considerable increase in the

measured frequencies. The repeatability in region C of Fig. 6 is poor and a maximum discrepancy

of about 15% between each run was recorded. The data points of region C correspond to the

arithmetic mean of five different runs, all of which have a sampling rate of 20000 Hz and a time

duration of 0.1 sec. Region C is characterized by small wave growth and thus small wave

amplitude at the measuring location. The signal in this region deteriorates considerably since, at

the measuring location, the wave amplitude is of the same order of magnitude as the laser beam

diameter.

Frequency (Hz) Frequency (Hz)

a) in region B of Figure 6:

Ziaser = A; Pq = 103.565 kPa;

mw = 44.16 g/s =^ f= 1250 Hz

b) in region C of Figure 6:

Ziaser = A; Pq = 103.565 kPa;

mw = 56.78 g/s f= 1836 Hz

Fig. 7 Power Spectral Density

Figure 7 shows the power spectral densities of the liquid sheet vibrations for an air supply

pressure of 103.565 kPa and for two different liquid flow rates. Figures 7a and 7b correspond to

data points sampled in regions B and C of Fig. 6, respectively. Figure 7b shows that while the

primary mode which corresponds to region B (see Fig. 7a for detail) still exists, a secondary, more

184
powerful mode with a higher frequency is displayed. Photographic visualization of the liquid sheet

shows that in regions A and B, the sinusoidal mode of oscillation dominates the liquid sheet

vibrations. In the transition between regions B and C, both sinusoidal and dilational waves exist.

In region C, the liquid sheet oscillations are predominantly of the dilational type.

Figure 6 shows that for each air supply pressure, the frequency of oscillations of the liquid

sheet undergoes sudden but continuous jumps as the liquid flow rate is increased. It is also shown

that these jump points are displaced to the right as the air pressure or the air velocity is increased.

In order to understand the jump phenomenon that were recorded, a separate test was

conducted to measure the spray angle. The criterion for measurement of the spray angle was to

follow the crests of the major disturbances. The surrounding droplet mists were not generally

considered. It should be mentioned, however, that when the breakup length recedes to the nozzle

exit, the droplet cloud was globally considered for spray angle measurement.

Figure 8 shows spray angle measurements as a function of liquid velocity for three different

air supply pressures. It is shown that for each air supply pressure the spray angle undergoes an

increase followed by a decrease as the liquid flow rate or the liquid velocity is increased. It is also

shown that the spray angle curves and the curves' maxima are displaced to the right as the air

supply pressure is increased. These maxima diminish with an increase in air supply pressure.

Careful examination of Figs. 6 and 8 shows that for low air supply pressure, namely Pair =

102.704 kPa and 103.565 kPa, the local maxima of the frequency curves in region B of Fig. 6,

closely correspond to the maxima of the spray angle curve. The reason for this correspondence

will be explained in subsequent paragraphs. There may be a mechanical element in the system

which has a resonant frequency near the maximum frequency displayed in this region. Near

resonance, the spray angle reaches its absolute maximum and the frequency of the liquid sheet

reaches a local maximum.

X3

60 -

20

Air Pressure

102 704 kPa

103 565 kPa

1 1 1.667 kPa

Fig. 8

Liquid Exit Velocity (m/s)

Spray Angle vs. Liquid Velocity for Three Different Air Supply Pressures

In order to demonstrate the phenomenon of resonant frequency and its influence on

maximizing the wave amplitude, the spray angle was plotted as a function of the vibrational

frequency of the liquid sheet. The results are presented in Fig. 9. It is shown in this figure that for

low air pressures, the liquid sheet oscillation is similar to that of a non-linear, damped, hard spnng

system.

In non-linear, non-damped, hard spring systems, it is found that the amplitude A undergoes a sudden discontinuous jump near resonance. The jump phenomenon can be descnbed as follows.

For the hardening spring system with increasing frequency of excitation, the amplitude gradually

increases until a point "a" in Fig. 10a is reached. It then jumps to a smaller value, indicated by

185
point "b", and diminishes along the curve to its right. In decreasing the frequency from point c, the amplitude continues to increase beyond point b to point d and suddenly jumps to a higher value

at point e. The shaded area in the amplitude frequency plot is unstable. The extent of instability

depends on a number of factors such as the amount of damping present and the rate of change of

the exciting frequency [2].

100 -

20

Air Pressure

° 102.704 kPa • 103 565 kPa

111,667 kPa

1000 2000 3000

Frequency (Hz)

Fig. 9 Spray Angle vs. Vibrational Frequency of the Liquid

Sheet for Three Different Air Supply Pressures

|Ai |A|

W/Wn

a) The jump phenomenon for the

hardening spring (no damping)

b) The jump phenomenon for the

hardening spring (with damping)

Fig. 10 Amplitude vs. Frequency for a Hardening Spring System

In the undamped case, the amplitude frequency curves approach the backbone curve (shown

dotted) asymptotically. This is also the case for the linear system where the backbone curve is the

vertical line at w/wn = 1 . With a small amount of damping, the behavior of the system cannot

differ appreciably from that of the undamped system. The upper curve, instead of approaching the

backbone curve asymptotically, will cross over in a continuous curve as shown in Fig. 10b. The

jump phenomenon is also present, but damping generally tends to reduce the size of the unstable

region [2]. To summarize, it is seen that for low air pressures, the liquid sheet oscillation is similar

to that of a damped, non-linear hard spring system.

As the air supply pressure is increased. Pair = 111.667 kPa, Fig. 9 shows that the jump

phenomenon is no longer present and the angle vs. frequency curve approaches that of a linear,

damped spring system. This change in character is due to the fact that at high air flow rates, the

oscillation frequencies of the liquid sheet show a monotonous increase with increase in liquid flow

rate. The cyclic increases and decreases, characteristic of low liquid flow rates, are no longer

present.

186
Air Velocity (m/s)

Fig. 1 1 Resonant Frequency vs. Air Velocity

Thus, it can be seen in Figs. 6, 8 and 9 that the spray angle reaches a maximum near the

resonant frequency. For low air supply pressures, the local frequency maxima of region B in Fig.

6 correspond to the resonant frequency. For high air supply pressures, the resonant frequency can

be inferred from Fig. 8. It is interesting to note that the resonant frequency increases with air

supply pressure or air velocity.

Figure 1 1 shows the influence of the air velocity on the resonant frequency. It is seen that the

resonant frequency increases linearly with air velocity. For each air flow there is a specific

vibrational frequency for the nozzle. The frequency of these vibrations is proportional to the air

velocity. The effect of introducing air in the nozzle simulates the effect of forced vibrations on the

nozzle jaws.

ACKNOWLEDGEMENTS

The authors gratefully acknowledge financial support from NASA Marshall Space Center,

grant number NAG 8-126, and Parker Hannifin corporation for providing the atomizer.

REFERENCES

1 . Mansour, A. and Chigier, N., "Physics of Fluids A," Vol. 2, No. 5, pp. 706-719, 1990.

2. Thomson, T.W., "Theory of Vibration," Prentice-Hall, pp. 392-397, 1979.

187

ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 16

AN EXPERIMENTAL STUDY OF AIR ENTRAINMENT BY MULTIPLE

SPRAY NOZZLES

N. Dombrowski and J. Singh

Department of Chemical Engineering, Leeds University

Leeds, United Kingdom

ABSTRACT

Gaseous entrainment by banks of multiple fan sprays is investigated. The results show

that an effectively isolated single fan spray has a higher entrainment efficiency then

within an assemblage. A hypothesis is presented that the reduction in efficiency is

caused by gas flow mutually decreasing the relative gas-drop velocity of its neighbours

and hence the drag force. Spray nozzle configurations have been devised for

maximising the relative velocity and hence entrainment. It is found that, although

entrainment is affected by nozzle size, it is primarily controlled by liquid loading.

INTRODUCTION

The process of heat and mass transfer between a continuous and dispersed phase is

involved in a wide variety of processes such as spray drying, extraction, combustion,

absorption, humidification and cooling. Sprays have particular advantages over other

methods of contacting liquids and gases. They are characterised by a large specific

surface within a relative small volume without having the pressure drop associated with

other types of contacting equipment. They also have the ability to entrain gas. These

features are examined in this work in relation to a novel forced-draft water cooling

tower, to be described later, where the entrainment property of a bank of sprays is

made use of in place of the conventional fan. This work examines gas entrainment

relative to the water flowrate per unit cross-sectional area. REVIEW OF PAST WORK

The aerodynamic behaviour of sprays is complex, and, in general, researchers have

investigated the. entrainment characterists of sprays for a specific systems using single

atomisers, Generalised experimental or theoretical analysis is complex and its

universal application has yet to be fully demonstrated.

EXPERIMENTAL

Nozzle Selection

Benatt and Eisenklam have reviewed gaseous entrainment rates for various types of

liquid sprays. Their studies showed that hollow cone swirl spray atomisers entrain

more air per momentum supplied in the liquid than from equivalent fan or solid cone

nozzles for the same length of spray. However, their study made no allowance for the

water loading for each spray type, which cooling tower manufacturers seek to

maximise to obtain a compact unit. In order to obtain a higher water loadings then

those possible with hollow cone sprays, whilst maintaining a fine spray, fan spray

nozzles, produced from single-orifice pressure nozzles, were selected for this work.

189
Bray nozzles OOOl 2,7 and 10 corresponding to flow numbers of 0.38, 1.69, 3.40 and

4.24 mls-v(kN/m^) ^'^ were selected.

Figure 1 demonstrates the process of drop formation from fan spray nozzles. The

liquid emerges as a thin attenuating sheet. Waves which are initiated by disturbances

on the sheet are rapidly amplified by aerodynamic action. When the waves reach a

critical amplitude they break down into drops via the formation of threads. The

lateral dispersion indicated in figure 1 is caused by a normal, or perpendicular, velocity

component derived from the waves.

Front View Edge View

Fig 1 . Drop formation from a fan spray nozzle

(No. 2 nozzle, 103 kN/m2)

Apparatus

Figure 2 illustrates the apparatus used to study air entrainment. Water is pumped from

a re-circulating tank through control valves to a nozzle manifold under test. The

supply pressure was taken from a tapping located just upstream of the manifold, whilst

the water flowrate was determined by means of a rotameter placed in the delivery line.

The manifold was housed in a rectangular test section with an adjustable width and/or

length of spray penetration. The nozzles were arranged on a common manifold aligned

to spray vertically downwards with the plane of the sheets lying parallel to each other.

The spray bank was arranged so that nozzle separation was accommodated by varying

the duct width; the duct breadth was maintained constant.

*Geo. Bray & Co. Ltd, Leeds, U.K.

190
The superficial duct air velocity upstream of the nozzle assembly was measured by a

hot wire anemometer close to the duct inlet. The air profile across the duct at this

location was found to be sensibly uniform over the full range of air flows encountered.

A range of nozzle supply pressures up to 240 kN/m ^ was employed which gave

superficial air velocities of up to 5 m/s in the duct.

RESULTS

Nozzle Size and Spacing

Figure 3 shows the variation of entrainment efficiency, expressed as the gas/liquid

(G/L) mass ratio, with water supply pressure for different nozzle spacings. The test

was carried out with No. 2 nozzles arranged along a spray bar (referred to as the row

configuration) with a spray length of 900 mm. The figure shows that the mass ratio

increases significantly with reduction of nozzle spacing and it may therefore be inferred

that a single isolated (or unbounded) nozzle is more efficient in entraining gas than a

nozzle assemblage. Air flow measurements for a single nozzle could not be undertaken

with the present rig because the relatively low flows could not be measured with

sufficient accuracy. The figure also demonstrates that; apart from a significant

reduction at low pressures, entrainment is effectively independent of pressure.

The effect of nozzle size on entrainment was studied with four sets of nozzles at three

water loadings, selected to demonstrate the effect of various loadings. The results are

plotted on figure 4 to show the vacation of mass ratio with nozzle size for water

loadings of 0.63, 1.3 and 2.6 1/s.m^ at a pressure of 138 kN/m^. The figure shows

that the entrainment increases significantly as a result of increased distance between

nozzles of the same size (points B, C and D for No. 2 and points E and F for No. 7).

Fig 2. Test rig

191
It also shows that at high water loadings (circa 2.6 l/s.m'^) an increase in nozzle size

from No. 2 to 7 causes a slight increase in entrainment which falls sharply with a

further increase in size to No. 10 nozzles (points D, F and G). It may be inferred that

any advantage gained from increased spacing with the larger nozzles is offset by loss in

momentum transfer, due to the relatively coarse drops produced by No. 10 nozzles.

(D32 mean drop sizes are 265 , 405 and 456/xm for No. 2, 7 and 10 nozzles

respectively). At low water loadings (0.6 l/s.m-^) No. 2 nozzles with a mean drop size

of 265/im entrain more gas than the smaller 130/im drops of the (XX) nozzle, because

the advantageous effect of increased spacing offsets any momentum transfer loss due to

the larger drop size (points A and B). A balance is reached between nozzle spacing and

drop size at a water loading of 1.3 1/s.m^ where it is seen that increase in nozzle size

from No. 2 to 7 has little effect on entrainment. It may be concluded that while air

entrainment is affected by nozzle size it is primarily controlled by water loading

irrespective of liquid pressure.

240 WATER SUPPLY PRESSURE KNm *

Fig 3. Variation of G/L with pressure and Nozzle separation

(aspect ratio = 4.5)

G/L

B66MM

A16MM

NOZZLE SPACING

C33MM

D17MM

000

0.63

WATER LOADING (LS^ m'')

1.3

E67MM

F33MM

G40MM

1.0 N0Z2LESIZE

0.0s 0.06

FLOWRATE PER NOZZLE LS"""

Fig 4. Variation of L/G with nozzle size and spacing

For row configuration

192
spray Length

The effect of spray length on entrainment was evaluated at a water loading of 1.3

Vs.m^ with number 2 nozzles spaced 33 mm apart. Two spray lengths, 450 and 900

mm were examined. The spray length was varied by spraying at different heights

above the duct discharge. The results are shown in figure 5 as a plot of the G/L ratio

against nozzle supply pressure. The figure shows that the variation of spray length in a

duct has a small effect only on entrainment, a 100% increase length resulting in an

increase in entrainment of 14%.

Fig 5. Variation of G/L with pressure and spray length

MECHANISM OF GAS ENTRAINMENT BY MULTIPLE SPRAYS

Development of Hypothesis

The above tests show that, generally, momentum transfer between air and spray

droplets can be enhanced by increasing the distance between adjacent nozzles. The

effect can be explained in terms of the likely mechanism that takes place. The

momentum transfer between the drops of a spray and the air is dependent on the

relative drop-air velocity. For single unbounded sprays, entrainment occurs throughout

the spray length from the surrounding quiescent environment. However, for single or

multiple sprays located in duct, the entrained air flows co-currently through the duct

causing a reduction of the relative drop-air velocity within the spray and hence a

reduction in entrainment. The larger the distance between adjacent sprays the lower is

the average gas velocity through the duct and hence the higher entrainment efficiency.

The following trials were carried out to test the hypothesis.

193
Two-tier Spray Arrangement

It has been demonstrated that entrainment efficiency can be improved with increased

nozzle spacing. A possible alternative method of increasing the spacing, whilst

maintaining compactness, is to employ a two-tier deck of sprays. A schematic diagram

of the arrangement is given in figure 6. Initial trials with banks of number two nozzles

in a row proved to be inconclusive as the sheet break-up at the lower tier was found to

be affected by drops from the upper tier. An alternative nozzle assembly was

employed using impinging jet nozzles (see below) where the sheets appeared to be unaffected by the falling drops. The results, shown in figure 7, indicate that the large

improvement that might have been expected by doubling the distance between the

adjacent nozzles was not obtained. These tests confirm that although entrainment by a

bank of sprays in a duct is a function of the actual physical spray separation, it is also a

function of the local relative drop-air velocity. The air velocity induced by each tier

reduces the relative air-drop velocity of the other, thus reducing the overall

entrainment.

3.8-

3.6- G/L3.4-

3.2- 3-

2.8-

2.6-

2.4- ZZ2-

Fig 6. Two-tier spray arrangement

4

1.8-

1.6- /

20 60 100 140 ICO 220 WATER SUPPLY PRESSURE KNm'

Fig 7. Entrainment by single and two-tier systems

194
Effect of Upstream Duct Aspect Ratio

To ensure adequate flow straightening, and hence, a uniform air flow pattern upstream

of the spray manifold, the experiments described above were performed in a duct of

relatively high aspect ratio (ratio of duct height upstream of manifold to duct width).

Thus, by virtue of its forward momentum, the air was constrained to flow co-currently

over the spray sheets.

It was postulated that, under these conditions, the air flow would flatten the spray

profile i.e. the spray width normal to the plane of the sheet, and hence increase the

inter-boundary spacing above that based on 'unbounded' measurements. At low aspect

ratios the flow pattern is less constrained to flow co-currently, and, the spray boundary

would be less affected with a consequent reduction in entrainment efficiency. This

reasoning is confirmed in Fig 8 which shows that a reduction in the aspect ratio from

6,8 to 1.5 causes entrainment to fall by up to 30% and in Figure 9 which shows a

comparison of the spray profiles of unbounded and bounded sprays.

3.8 -

3.6 -

G/L3.4-

3.2 -

3

-

2.8 -

2.6 -

2.4 -

1 H 1 1 1 1 1 1 0 2 4 6

ASPECT RATIO

Fig 8. Variation of L/E with aspect ratio

Fig 9. Comparison of edge profiles, LHS, single nozzle; RHS row configuration

(6.8 aspect ratio)

195
CONCLUSIONS

It has been found that the entrainment potential of a bank of sprays is reduced as the

nozzles are brought close together. The spray from a single nozzle in a free

environment entrains air along its full length normal to its axis, the air being taken

from its locality. However, for a bank of sprays in a duct, the air is forced to flow

parallel to the spray outside the drop envelope and as a result the relative drop-air

velocity within each spray envelope decreases. These air profiles reduce the drag

between the air and the spray droplets and result in lower momentum transfer and

hence lower entrainment at any given spray length.

The relative influence of a spray on its neighbours can be reduced by greater physical

separation of adjacent nozzles, although this leads to lower water loadings.

In general air entrainment is controlled primarily by water loading, nozzle size having a

less significant effect.

ACKNOWLEDGEMENTS

The provision of a resecu^ch grant from Hall-Thermotank Limited is gratefully

acknowledged.

REFERENCES

(1) Rasbash D. J. and Stark G. W. V. Some aerodynamic properties of sprays.

The Chemical Engineer. (1962)

(2) Briffa F. E. J. and Dombrowski N. Entrainment of air into a liquid spray. A.

I. Ch. E. Journal 1966 12, 708.

(3) Benatt F. G. S. and Eisenklam P. Gaseous entrainment into axisymmetric liquid

sprays. Journal of the Institute of Fuel 1966 309.

(4) McQuaid J. Air entrainment into axisymmetric sprays. Proceedings of the

Institution of Mechanical Engineers 1975, 189 28.

(5) Heskested G., Kung M. C. and Todtenkopf N. F. Air entrainment into water

sprays and spray curtains. The American Socety of Mechanical Engineers, 76-

WA/FE 40.

(6) Arrowsmith A. and Foster P. J. The motion of a stream of monosized liquid

drops in air. The Chemical Engineering Journal 1973 5. 243-250.

(7) Hedley A. B, Arrowsmith A. and Foster P. J. Entrainment and evaporation in

single streams of monosized drops. Chemeca 1970.

196
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 17

BEHAVIOR OF SURFACE WAVES ON THE BREAKUP

MECHANISM OF A LIQUID COLUMN

D.-J. Kim* and C.-W. Lee^

Dept. Mechanical Design Engineering, Pusan National University, Pusan, Korea

^Dept. Mechanical Engineering, Kyungpook National University, Taegu, Korea

ABSTRACT

In twin- fluid atomizers it is difficult to analyze the behavior of liquid

coliomn since it is affected coraplicatedly by many factors. The simplest shape of

liquid, which is injected through these atomizers is a solid jet before it

encounters with the gas stream. The liquid jet is deformed at first, and then

disintegrates into drops. The objective of this study is to investigate

experimentally the basic disintegration mechanism of the liquid column in coaxial air flow. The behavior of the liquid jet ejected from the circular

capillary tubing into co-axial air flow was observed by the high-speed camera.

Breakup patterns, surface wave length, amplitude, and growth rate of the

disturbance were measured to explain the behavior of the jet. The wave length

of the disturbance formed on the liquid surface becomes shorter at the downstream

and the trend of change of amplitude as time passes deviates from the linearity.

The growth rate of the disturbance is not much influenced by the liquid flow

rate. And an empirical equation of the breakup length is obtained as a function

of air Reynolds number, liquid Weber number, and initial breakup length,

INTRODUCTION

Lately, the twin- fluid nozzle, in which combustion air is used to atomize

liquid jet, has been used in the combustor of jet engines, because of the

characteristics of fine atomization optimized distribution, and strong mixing of

air and fuel. But the limit ranges of ignition and flame hold in the combustor

with this nozzle become narrower than those with pressure nozzle. Owing to these

facts the characteristics of atomization, fuel dispersion, and fuel distribution

become worse at low speed operation conditions.

In order to eliminate the defects it is necessary to study the effect of

air flow on the basic breakup mechanism of liquid jet, the evaporation process

of droplets, and the mixing process of droplets and air. For the first time,

Rayleigh [1] developed a theory about the breakup mechanism of liquid jets.

Casleman [2] reported that the breakup of liquid resulted from the growth of

unstable waves formed on the liquid jet surface by the aerodynamic interaction

at the liquid-air interface. Sato [3-5] suggested that the breakup of smooth

liquid jets came from the dispersion of propagation of nonlinear waves. Situ [6]

197
pointed out that growth rate of swell and neck formed on the liquid jet surface

was changed with the characteristics of nonlinear effects. He also investigated

theoretically the effect of heat and mass transfer on the instability of liquid

jets. These studies were performed on the breakup mechanism of the liquid jet

itself at different experimental conditions. A systematic view between these

studies has not been developed yet. Most studies inclined to qualitative

analysis. Especially , the effect of the disturbance behavior on the breakup of

liquid jets has not been made clear. So, experimental equation proposed for

breakup length was applied to limited nozzle only.

The purpose of this study is to investigate the disintegration mechanism

of liquid jets, with change in surface wave length, amplitude, growth rate of

disturbance, and maximum growth rate to dimensionless wave number, which were

examined by using a high speed camera. The empirical equation of breakup length

as a function of air Reynolds number, liquid Weber number, and initial breakup

length is obtained.

EXPERIMENTAL APPARATUS AND PROCEDURE

An apparatus was newly designed in order to study the disintegration

mechanism over wide ranges of operating conditions. The experimental apparatus

consists of the air and liquid supply system, injection system, and high speed

photographing equipment.

When liquid, compressed by a compressor, is injected vertically through an

injection nozzle into the stagnation air or the co-axial air flow, the liquid

flow rate is controlled by the injection valve. Air used to atomize the liquid

is pressurized by a compressor and then flows vertically in co-axial liquid flow.

The diameter of the liquid nozzle, used in this study, is 0.7mra. Diameters

of the air orifice are 4mm, 6mm, and 8mm. Liquid flow rates are 0.69 g/s , 0.89

g/s . and 1.14 g/s. When the air flow was increased gradually for each different

liquid flow rate, wave length, amplitude, and growth rate of disturbance were

measured with photographs by the high speed camera.

RESULTS AND DISCUSSION

Breakup Patterns

Figure 1 shows photographs by using the high speed camera in order to

compare the formation process of the liquid droplet in stagnation air with that

in co-axial air flow. Figure 1(a) shows the photographs of a liquid jet in the

absence of air flow. This also shows the surface wave formed on the liquid column

symmetrically. The length of the ligament between swell and neck is slightly

longer before producing droplets. The ligament is disintegrated into small

satellite droplets at the downstream. Figure 1(b) shows the photographs of

liquid jet in co-axial air flow. The amplitude of surface waves becomes shorter

than that of Fig. 1(a). The reason of the generation of satellite droplets from

the ligament is considered as follows. Before the longer ligament was formed, the

shorter ligament was broken up, and then absorbed into the main droplet due to

surface tension at the downstream. In the case of Fig. 1(a), the preliminary

droplets were observed before the droplet generation. This droplet was a result

of the balance of gravity and surface tension. But, in the case of Fig. 1(b),

the preliminary droplets were not observed and were formed at the tip of the

smooth jet just before the preliminary droplet formation. The effect of air flow

198
0-1

10-»

mm

i 1

i ,

1

I

(a) Ua = 0 m/s (b) Ua = 10 m/s

Fig. 1 High speed photographs of liquid jet

(dl=0.7mm, Ml=0.89g/s, At=0.67ins)

3 r

E

Ml = 0.39 g/s

0 U a = 0 m/s

A 1 0 • 1 4.

1

c

4>

>5

2 - t

o o o o Q

aa o

0 5 10 1&

t/T * 100

Fig. 2 Variation of the wave length with elapsed time

for different air velocity ( Ml = 0.89 g/s)

199
on the liquid breakup is stronger than many other parameters. Both cases show

that the amplitude of surface waves increases at the downstream. The reason is

that the first formed surface waves were pushed by following surface waves, and

that the gravity and the surface tension act on the liquid jet. Thus, the effect

of co-axial air flow on disintegration of the liquid jet, as expressed in Weber's

equation, results in changes of relative velocity of the liquid and air. initial

amplitude of the disturbance, and growth rate of the disturbance.

When the liquid is injected into stagnation air, droplets are slowly

generated because of the action of surface tension and gravity, even though the

initial disturbance is small. But the liquid jet in co-axial air flow is more

largely subject to the effect of an increase in growth rate of the initial

disturbance due to aerodynamic interaction than the effect of surface tension.

Surface Wave Length

Figure 2 shows that the wave length on a liquid jet through a nozzle is

changed with time at the liquid flow rate of 0.89g/s. Time, T, represents the

value that breakup length is divided by liquid velocity. t is time that is

passed. When the liquid jet was in co-axial air flow, the surface wave length

of disturbance decreases with time, and the reduction rate of it increases as air

velocity increases. The reason is that the aerodynamic force acting on it becomes

stronger than surface tension by the balance of forces acting on the disturbance.

And the instantaneous time that the surface wave length decreases rapidly, as

indicated by the arrow in Fig. 2, makes earlier with increasing air velocity.

This instant is nearly equal to the time that the velocity of the first formed

surface wave and the following surface wave is inverted.

Growth Rate and Amplitude of Surface Wave

Figure 3 shows that the amplitude of the disturbance on a liquid jet is

changed with the ratio of the time to the production period of droplets at Ua=0

m/s . The slope represents the growth rate of the disturbance. The amplitude

increases linearly for all liquid flow rates. The flow condition of liquid flow

rate of 0.69 g/s and 0.89 g/s corresponds to the applicable region of Rayleigh's

theory (so called smooth region). The disintegration of the liquid jet results

from the growth of axial symmetry caused by surface tension and gravity.

Figure 4 shows the amplitude of the surface wave for the liquid flow rate

of 0.89 g/s. The shape of the curves at Ua>0 m/s deviates from the linearity as compared with Ua=0 m/s. The absolute values of amplitude were slightly

increased. The slopes of the curves showed the similarity to each other until

the first formation of droplets, but change irregularly with time. The reason

is that the frequency in generation of a droplet changes with the increase of the

air velocity. The change in slope corresponds to the time that the velocity of

the surface wave changes rapidly.

The maximum growth rate occurs at the dimensionless wave number of 0.697

in the case of Rayleigh's linear theory [7], less than 0.697 of Wang's analysis

[8] , the lower wave number of Yuen [9] . Finally the nonlinear effect at above

dimensionless wave number used in the instability theory dominates the

disintegration processes. But the maximum growth rate occurred at near 0.7 as

shown Fig. 5. This means that Rayleigh's linear theory has the limit of

application for liquid jets in co-axial air flow.

200
oo

o

o

•o

•o

in

»o

<1

o

o

o

VI V

m

E

\u>

o

O

at

00 II

re

o

It

o < • - ^

lO

(ului) ^

in

M

«o

OO

j:: OO

+-> .-,

o

S II c ^ a >x

QJ +-> O "H

CO

U

—I

O

a -H

•-H >

T3> c o u

CO QJ

=4-1 <4-l 3 —I

CO T3

CO

0)

U ^ O

+-> V-i

tt-i 0) O B CO

—I T3

+-> QJ

CO tn u CO

CO —I > 0)

130

o 8

<1 o<i

<1

ss/ui 0)

<o

91

OO• o o* o •

II

II

re

o < •

_I I L

•<i

o I o

10

'O

N

in

(N

If)

•O

aoIICO

C CO

Q)

U

B

QJS

UO

CO -H —I ^ a

CO -a

-a 3

QJ> —I

CO

c

QJ QJ U t_

CO QJ 3 -H

CO "D

QJ I- JZ o ^ t+_

'4-. QJ O B O— T3

+-> CD

CO COa U CO

QJ

CO

OO

201
—

E

_E

O)

je

E z

,001

OOo • .07 3 o o ob

oO

ino

E3e>

raM

10«e2wc9E

no

+->

u 3 0} O X!B

oo 3C O Q) c >

CO CO J3 ==

=!

L. CO

(-> (0 W 0

T3co

C

0) CO

0)c

0)e

0) T5 c T3 o C

CO

4-> CO 0)

»-> (0 CO

oo

in

o

00

9i9t i||/wojB •3Ui?qin»i!a

202
Empirical Equation of Breakup Length

Grant and Middleman [10], Sato [11], etc., proposed the empirical equation

of the breakup length of a liquid jet in stagnation air. Especially, Sato had

developed the equation in terms of Weber number and air velocity for a nozzle

diameter of 2.09 mm in co-axial air flow. Figure 6 shows the change of breakup

length for various diameters of the air orifice. When the equation of the

breakup length is formulated, the effect of air flow rate with the change of air

orifice diameter must be included. The empirical equation for breakup length

includes parameters such as the initial breakup length, liquid Weber number, and

air Reynolds number including air velocity and air orifice diameter as shown in

Fig. 7.

The equation is set up as follows

.

L = (T^e)O°°^5v^exp(-0.0143N/^) (1)

^0

This equation means that the breakup length of liquid column changes

complicatedly under the influence of parameters such as the aerodynamic

interaction, the geometrical shape of air orifice, and of liquid column.

CONCLUSIONS

To investigate the disintegration mechanism of a liquid jet in co-axial air

flow, breakup patterns, surface wave length, amplitude, and growth rate of

disturbance were measured with photographs by a high speed camera.

(1) The wave length of the disturbance formed on the liquid surface becomes

shorter at the downstream and the trend of change of amplitude as time

passes the disturbance deviates from the linearity.

(2) The growth rate of the disturbance is not influenced much by the liquid

flow rate

.

(3) An empirical equation of the breakup length is obtained as a function of

air Reynolds number, liquid Weber number, and initial breakup length.

REFERENCES

1. Rayleigh, Lord., "On the Instability of Jets", Proc. Lond. Math. Soc,

Vol. 10. pp. 4-13 (1878).

2. Casleman, R.A. , "The Mechanism of Atomization Accompanying Solid

Injection", Report National Advisory Committee for Aeronautics, No. 440,

pp. 735-746 (1932).

3. Sato, H., Makai, T. and Saito, M. , "Non-linear Waves on Smooth Jet of

Liquid Column" . 8th Japan Conference on Liquid Atomization and Spray

Systems, pp. 7-12. (1980).

4. Sato, H, and Saito, M. , "Non-linear Wave Propagation and Disintegration of

a Circular Liquid Jet" , International Conference on Liquid Atomization and

Spray Systems. 1-3. pp. 19-27 (1982).

5. Sato, H. , Sakai, T. and Saito, M. , "Nonlinear Waves on Smooth Jet of

Liquid Column II", The 10th Conference on Liquid Atomization and Spray

System in Japan, pp. 19-24, (1982).

203
6. Situ, M. and Schetz, J. A., "Computational and Experimental Study of the

Effect of Mass Transfer on Liquid- Jet Breakup", AIAA Journal, Vol. 23,

No. 2, pp. 254 -261, (1985).

7. Rayleigh, Lord, " Theory of Sound". Macmillan, London, Vol. 2, pp. 43-375,

(1926)

.

8. Wang, D.P. , "Finite Amplitude Effect on the Stability of a Jet of Circular

Cross -Section", J. Fluid Mech. , Vol. 34, pp. 299-313 (1968).

9. Yuen, M.C. , "Non-linear Capillary Instability of a Liquid Jet", J. Fluid

Mech., Vol. 33, Part 1, pp. 151-163, (1968).

10. Grant, R.P. and Middleman, S., "Newtonian Jet Stability", AIChE Journal,

Vol. 12, No. 4, pp. 669-678, (1966).

11. Nagaosa, S., Matsui, H. , Tokuoka, N. and Sato, G.T., "A Study of the

Disintegration of a Liquid Jet in an Air Flow", International Conference

on Liquid Atomization and Spray Systems . 2-1, pp. 29-36. (1978).

a

eo

1.0

0.8

0.6

0.4

a

I 0.2

0.1

40

M 1 (9/S)

0 (mm]

4 S 3

0.30 O o •

0.42 o

0.50

I,

o

s

l/lo = weO-007S

^

Exp(- 0.0143/R?)

60 30 100 120

Air Beynolds numbee Re

O.S

1 40

Fig. 7 Breakup length ratio of smooth jet as function of Reynolds

number, air orfice diameter and liquid flow rate

204
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 18

LIQUID ATOMIZATION WITH HOLLOW CONE NOZZLES

H.D. Dahl and E. Muschelknautz

Lehrstuhl und Institut fur Mechanische Verfahrenstechnik

Universitat Stuttgart

Stuttgart, Germany

1. ABSTRACT

Hollow cone nozzles consist of a spin chamber with an excentrical or axial inlet generating rotational flow. Because of

the great centrifugal forces a thin liquid film is formed in the

outlet.

For the calculation of the flow in the nozzle cyclone theory is

used. Cyclone theory consists of simple, physically founded

equations for outlet velocities and pressure loss, which are

taking into account the influence of wall friction.

We found, that the mechanism of disintegration of the liquid

film depends on the outlet velocity of the nozzle. For lower

velocities the breakup of the film takes place under the

influence of aerdoynamic forces forming sinuous waves according

to the model of Dombrowski and Johns /I/. If the outlet velocity

surmounts a certain value the disintegration is due to turbulent

liquid forces, i.e. the film is atomized. The maximum drop size

is then a function of the Weber and Ohnesorge numbers.

All calculation methods are based either on physical balances or

on a complete set of similarity numbers. They have been verified

by a systematical variation of nozzle geometry and of liquid and

gas properties. Therefore they can be applied for nozzle design

and scale-up.

2. FLOW IN THE NOZZLE

2 . 1 Theory

Cyclone theory /2/ takes into account the influence of the wall

friction of rotational flow. Figure 1 shows a hollow cone nozzle

with axial inlet.

205
1 X X

I^Tq —

Ur

Figure 1: Angular velocity in a hollow cone nozzle

After having entered the spin chamber through the grooved core

the pressure gradient of the rotational flow is impressed on the

liquid. Due to this the flow contracts and accelerates to the

outer angular velocity ua. The contraction coefficient a is

defined as the angular momentum's ratio before and after the

contraction of flow:

a = M u_ r_

e e ^ "a ^ae e

^a ^a

(1)

The tangential stress due to wall friction is

T = - A u (2)

206
with geometric mean values for tangential velocity and radius

the loss of angular momentum results as

"r = dA^r = X Aj, -2^1^^ (3)

The balance of momentum equation (4), leads to the equation for

the angular velocity in the outlet of the nozzle:

Mur = Mu. r. + M„

a a 11 TR (4)

u r /r.

a a' 1

u. = (5)

^ ^a

2 V

^a/^i

The pressure loss of the nozzle consists of a part caused by the

acceleration of the liquid and of a part caused by wall

friction.

The wall friction pressure loss results from a balance of

energy

:

Ap = X

^

(u u.)^/2 (g)^

The acceleration pressure drop is calculated by the Bernoulli

equation:

By means of experiments this pressure drop can be as well expresised as a function of the angular velocity in the outlet:

with

2

u . u

.

C. = (1 + 2 + ) (9)

V. V.

207
and

V.

1

V

1

(10)

In comparison to hollow cone nozzles the pressure loss coefficient in cyclones is higher, because the swirl flow in the

cyclone is reversed.

With equations (7), (8) and (9) the axial velocity is

^ax = 2 u. V. (11).

The film thickness in the outlet is given by continuity equation

as

= - J (12). " ^ax

2.2 Experimental Results

The examined nozzles had ratios of spin chamber radius to outlet

radius from 2 to 5. The spin chamber radius varied from 6 to 12

mm. The physical properties of the liquids which were used for

experiments are displayed in Table 1.

[kg/m"^] [10"-^Pa s] [lo"-^Nm]

Water

Glycerin/Water

Ethanol/Water

Suspension of lime 35%

Table 1: Liquid properties

1000 1 72

1215 75 65

972 1.9 42

1280 2.2 49

With an experimentally determined contraction coefficient a the

wall friction coefficient X could be calculated out of the experimental data by means of equation (5). A decreases for

laminar flow with increasing Re-number until a constant value

for great Re-numbers (turbulent flow) is attained (Figure 2).

208
0 001 ' I 1 I—I

I I I1 1

1 1 '—t—I—I

I 1

1

1000 10000 100000

n 9l U

Re =

Figure 2: Friction coefficient X

3. FORMATION OF DROPS

3 . 1 Theory

Sinuous wave breakup, which is due to aerodynamic forces, has

been closely examined by Dombrowski and Johns /I/. From their

model, equation (13) for a nonviscous, conical sheet was

derived.

1/6 a 1/3 r. 1/3

d - 8 i-^) ( 5—) ( ) (13)

PQ p. 8 8 tany

Li

As this model does not take into account viscous forces of the

gas, we found, that it had to be supplemented by a Reynolds

number for the gas flow in contact with the liquid film :

209
p w 5

Re^ = (14)

With equation (13) and (14) a complete set of similarity numbers

according to similitude theory is given. Our experimental

results justify the omission of the liquid viscosity in this

context

.

At higher outlet velocities the liquid film is atomized by

turbulent liquid forces (/3/, /4/). This is described by two

similarity numbers, the Weber number and the Ohnesorge number:

p d

We = (15)

Oh = (16)

J cr d^^^ ' "^L max

3 . 2 Measurement Techniques

The maximum drop size was measured by sampling the largest drops

in oil. A Malvern particle Sizer was used for the determination

of the drop size distribution.

3.3 Results

The application of equation (13) on experimental data leads to

Figure 7. Hence the maximum drop size for wave breakup is

calculated by equation (17):

1/6 cr 1/3 r^ 1/3 w 6 1/5

Pq Pi,^^ 5 tanr

The exponent of the Reynolds number was found out by means of

correlation of experimental results.

210
2000

dmax

l|im]

1000

800

600

4004

Liquid film disintegrt

by aerodynamic wav

ition

es /o*

O X

cVy^

2-10

-4

PQ

0

Pl w26

5- 10-'

1/3

) (

10- 2-10

-3

Figure 3: Maximum drop diameter for sinuous wave breakup

Equation (17) is valid for sinuous wave breakup. If the velocity

of the liquid film is raised up to a certain point the turbulent

forces in the liquid film will surmount the aerodynamic forces,

which cause sinuous wave breakup. The drop size is then smaller

compared to calculation with equation (17). The stability

criterion for maximum drop size is given by Figure 4:

We = 4,1 10^ Oh-"-/^ (18)

for 0,003 < Oh < 0,03. The experiments showed, that gas density

and liquid film thickness have if any only small influence.

Table 2 gives characteristic parameters of the measured drop

size distributions:

wave breakup atomization

^32/^max 0.35 0.31

d^^/d 0.48 0.42

50' max

Table 2: Characteristic drop diameters

211
5-10'

X

COE

T3

c\j D

II

10'

5-10'

Atomization:

We = 4.1-104 Ohi/6

0.002

a

as

p G = 1-2 kg/m3

Water 93= 1.2 kg/m^

Water 93= 1.2 kg/m^

Glycerin / Water

Ethanol / Water

Suspension of lime

H 1 1—4 1 1 1 1 »-

0,01 0.1

Oh = Vo Pl dmax

Figure 4: Atomization of the liquid film

4 . NOTATION

A

d

^50

d

max

M

M

Oh

APi

^Pe

r

r

u

u

m

area

drop diameter

Sauter diameter

mass medium drop diameter

maximum drop diameter

angular momentum

mass flow

Ohnesorge number

acceleration pressure loss

friction pressure loss

radius

r r

.

a 1

m

mean radius

angular velocity

\fu u7 : mean angular velocity

a 1

m

m

m

m

m

Nm

kg/s

N/m^

N/m^

m

m

m/s

m/s

212
V. : mean velocity in the outlet m/s

: axial velocity of the liquid film m/s

ax

V : volume flow m^/s

w : velocity in the outlet m/s

We : Weber number

H : spray angle °

5 : film thickness m

: pressure loss coefficient

Tj : dynamic viscosity Pa-s

X : friction coefficient

p : density kg/m^

(T : surface tension N/m

T : tangential stress N/m^

Subscripts

:

a : on spin chamber radius

e : inlet

G : gas

i : on outlet radius

L : liquid

R : friction

5 . REFERENCES

/I/ Dombrowski ,N. , Johns, W.R. :

;

Chem.Eng.Sci. 1963, Vol. 18, S. 203 - 214

/2/ Muschelknautz , E. , Krambrock, W.

:

Chem.-Ing.-Techn. 42 (1970) 5, S. 247 - 255

/3/ Troesch, H.A.

:

Chem.-Ing.-Techn. 26 (1954) 6, S. 311 - 320

/4/ Grassmann, P.:

Physikalische Grundlagen der Vefahrenstechnik

Salle u. Sauerlander, 3. Aufl. 1983

213
i

i

1

1

I

I

I

i

I

t

i

i

I

i

i

i

!

1

I

r

I

I
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 19

DROP SIZE-VELOCITY CORRELATION AND DROP FLUX IN THE SPRAY

CONE OF A FAN SPRAY ATOMISER

G. Schulte and I. Ozdemir

University of Bremen, Verfahrenstechnik

Bremen, Germany

ABSTRACT

Sprays of liquids are of high importance in many technical processes.

Therefore the measurement of drop size and velocity have always been of interest

to research work [1-5]. In such work laser measurement instruments give the

advantage of non-invasive access to the wanted location in spray cones [6-9].

However, although a lot of work is done in this field there still seems to be

sometimes unclearness in respect to interpretation of measurement results in

terms of the parameters of the dispersed flow. The work here reports on an

investigation of a spray by means of the ^^ase Doppler anemometer (PDA)

intrinsically giving the flow parameter "particle flux" and additionally by means

of diffraction s.ize meter (DSM) intrinsically giving the flow parameter "particle

concentration" . The relationship between these flow parameters is given by the

drop size-velocity correlation at the respective location in a spray. This

correlation and the development of the flow parameters mentioned above will be

discussed in the following paper.

GENERAL REMARKS

In investigation of drop flows i.e. sprays are two points of main interest.

On one hand the description of the dispersed flow situation on the other

knowledge of the individual elements, the drops. If we look at sprays, by the

view of particle measurement technology [10], we look for a physical property

that characterizes the particles - mostly the "size" given as an equivalent drop

diameter - and we look for the particle velocity as a second item which

characterizes the particle motion and causes their location to depend on time.

So we not only look at particles in terms of their physical properties but also

in terms of flow parameters. A common parameter related to space is the

"particle number flow concentration" c^. Another parameter related to time is the

"particle number flow" . The particle flow often will additionally be related to

a flow cross -sectional area so the resulting parameter is the "particle number

flux" fg. Basically in a steady spray these parameters can be defined for each

location in the spray cone and at a point in time ((f^-f^, (x,y,z,t);

Co=Co(x,y ,z , t) ) . These flow parameters normally vary with the different spray

locations because of the spreading character of a spray and ac- or deceleration

of the drops. Particle measurement methods give distributions of the measurement

215
quantity according to size (and velocity) and intrinsically measure a flow

parameter either drop "flux" or drop "concentration". Normalized distributions

from spray measurements and mean diameters derived from therefore always should

be marked as either flux (index "f") or concentration (index "c") related. If

correlated values of size and velocity are known these results are convertible

[9] . It depends on the field of application whether one is more interested in

the drop "flux", i.e. in spray-coating or in the drop "concentration", i.e. in

burning sprays of fuel.

The following report on the experimental analysis of a spray of water is

organized as follows: First a description of the experimental rig and the

measurement locations. It will be followed by the presentation of profiles of

medians of drop size from PDA and DSM measurements. Thereafter 2-dimensional

number distributions are analyzed with respect to the drop size -velocity

correlation. And at last we will look at drop number and drop volume flux

through the spray planes under observation.

EXPERIMENTAL PROGRAM

Water was sprayed in a vertical arranged flow channel of 1000 mm length

having a cross section 900 mm to 750 mm (Fig. 1). Grids at top and bottom of

the chamber provided a block profile of the superimposed air flow of v = 0,2

m/sec. The spray investigated was a flat full cone spray generated by a fan

spray atomiser. The nominal operating pressures in our experiments were 2 bar,

3 bar and 4 bar. This corresponds to theoretical spray angles of 51°, 65° and

69°. Figure 2 shows the shape of the spray. Drawn are three horizontal planes

E2 at z=20 cm, E3 at z=30 cm and E4 at z=40 cm. The experimental program

reported here covered phase Doppler (PDA) measurements at 350 locations (65 in

E2; 125 in E3 and 160 in E4) and diffraction measurements (DSM- "line of sight")

at 120 locations 30 in E2; 40 in E3; 50 in E4)

.

The phase -Doppler- instrument was built by components from Dantec and a selfmade processor:

He-Ne laser; 25 mW wave length: 632,8 mm

intersecting angle: 2,5° elevator angle: 5,8°

off -axis angle: 60°

A measurement location for PDA in a spray plane z is defined by PDA (x,y).

The diffraction size meter (DSM) was a Helos 12 LA:

He-Ne laser: 5 mW diameter of laserbeam: 13,28 mm

size range: 7,9-1750 mm focus length of receiver: 1000 mm

In a given spray plane z the DSM-measurement location is given by the "line of

sight" parallel to the y direction defined by DSM (x)

.

DISCUSSION OF RESULTS

Figure 3 shows the profiles of the volume medians d^Q 3^ obtained by the

PDA. The number of counts included in the evaluation was 4000, however, we found

that even a number of 2000 counts leads to a sufficient accuracy if we look at

the marginal number distributions because some larger drops have here no

significant influence. Working out volume distributions causes that the number

of large drops will become a most important influence. On the other hand, a

counting method like PDA may become a time consuming method depending on the

drop flux at the respective measurement locations . So in our case , the

measurement- time for 4000 drops reached from 50 sec in the centre on the spray

axis up to 500 sec at the spray boundary with corresponding acceptance rates of

216
0,5 up to 0,9. For balance calculations the results were extrapolated assiiming

the distributions of drops not included in the evaluation are the same as from

those included. So the number of 4000 counts seems to be a reasonable

compromise. This the more as we have a large number of measurement locations so

that significant deviations from the general trend are easier recognizable as

unsteadiness in the profiles of the parameters of interest. The top of the

profiles look very smooth and show that in the core of the spray the medians are

lowest as expected. The shapes of the profiles are very similar. The step in

the djo 3£ profile of spray plane E4 was caused by an adjusting error which

affected all the measurements at locations x=-l up to x=-12 cm spray plane

z=40 cm. The profile of djQ along x-axis (run a in Fig. 4a) show more clear

the effect of the adjustment error. To find out whether this unsteadiness was

due to an asjrrmnetry of the spray pattern or due to a supposed adjustment error

we ran an additional series of measurements by PDA (run b in Fig. 4a) and as

reference by DSM (Fig. 4b). We found that the PDA results fit very well with

those obtained in the first series except these values which we assume to be due

to an adjustment error. So the conclusion: before believing in absolute values

of point measured results it is recommended to look at the results from locations

in the vicinity to find out whether the results fit with the general trend of the

parameter in question. Figure 4b gives the volume median diameter d^Q obtained

from DSM "line of sight" measurements at different x-locations. The "line of

sight" represents the axis of the intersecting volume between the laser beam and

spray, and was parallel to the y-axis. The DSM-result gives an average of all

drops in the intersecting volume. A transformation of these integral data into

point data is possible by a Fourier transform method [11]. However because of

the shape of the spray and the alignment of the "line of sight" overall and point

data show the same behavior [6] and therefore for the aim of this work the

overall "line of sight" data are sufficient. The curves of djg 3^, show for the

PDA results a clear symmetry with respect to the spray axis represented by the

ordinate x=0.

Figure 5 shows the 2 -dimensional number distribution of size and velocity

hQ£(d,v) on the spray axis and at the boundary of the spray cone in 3 spray

planes in question. From these again we recognize the drops on the spray axis

are smaller. The difference between velocity of small and large drops will

become smaller with increasing distance from nozzle tip and from spray axis or

in other words more general: the correlation of size and velocity of the drops

as we see is dependent on the location in the spray. As 2 -dimensional

distributions contain a large amount of information it is difficult to get

clearness in looking at a number of them at the same time. So it is usual to

restrict oneself to looking at the key parameters of interest. The drop sizevelocity correlation seems to be a parameter worth looking at respectively to be

meaningful to describe the spray cone in terms of it. Looking at the so called

empirical regression curve v^^ = v (dj^) which is often given as standard software

of PDA- instruments we mostly realize unsteadinesses at the tails where the

distribution counts lower numbers of drops in its bins, v(d^) in Fig. 6 shows

a typical shape of such curves. Computing the linear regression curve v^.^, of

the v(d^) -values this gives only a poor fitting although there seems to be a linear correlation for the main bulk of drops. The result will become much

better if we include only this main bulk of drops into the calculation of the

linear regression. The resulting straight line Vq 2 i^ow fits very well with the

empirical regression curve besides at its ends. Included in the calculation are

here only these size bins of the boundary distribution hQf(d^) in which there

were counted numbers of more than 0,2*\f-j„^^ drops meaning high enough by view of

217
statistics. This procedure of data preparing results in a linear correlation

coefficient of rg 2=0>994 which points out a really good fitting whereas only

about 7% of all drops are excluded from the calculation. If the bulk contains

more than 90% of the drops this can be seen as representative for the sizevelocity correlation, so we looked for the linear regression in this way for all

measurement locations and we found fairly good results for the ro 2 -values all

over the spray, seen in Fig. 7a, which means the correlation of the main field

of drops is given fairly good by a linear regression. No significant differences

are seen which depend on the measurement location. For comparison in Fig. 7b

are given the r^ ^-values in spray plane E2 which confirm the result of Fig. 6:

the correlation coefficient of all drop size bins are included is mostly less

than 0,8 saying that not linear correlation is given if the tailvalues are

included. Figure 8 gives the linear regression coefficient b^^jg 2

" the measure

of ascent of the regression line- for the x-axis location corresponding to the

ro 2-values of Fig. 7a. The basic information is that there is a clear

dependency on the measurement position recognizable. As expected in general at

downstream positions b^jjo,2 smaller: the angle of ascent of the regression

line becomes smaller. The situation on the spray axis is different dependent

on atomiser distance. So the air entrainment seems to cause an equalizing effect

on the drop velocities which is seen at the boundary of the spray cone and on

the spray axis here more evident at downstream positions where the b^^jg 2 lower on the spray axis. The development of the linear regression coefficient

^vdo,2 °f respective coefficient of correlation tCq 2 along the spray axis

is seen in Fig. 9a. We recognize increasing of b^^g 2 and after getting the

highest value at z=15 cm going down to b^^jg 2=0,01 at z=50 cm. Again the

correlation coefficient rg 2 points out the good quality of the linear regression

for all locations. The b^^jg 2 curve shows - if extrapolated to a starting point - that first all drops seem to have the same starting velocity and that very soon

because of deceleration of the small drops we have a significant increase of

^vdo 2- The development of the arithmetic means d^ gf (Fig. 9b) is affected by the

air entrainment effect and has its lowest value at z=22,5 cm. At the same time

the velocity means Vj^ g^ are represented by a steady line lastly reaching the

velocity of the air stream.

We will now leave the distribution parameters and look at results computed

from the distribution as spray parameters (Fig. 10) . First the drop number flux

fo as profiles in the observed spray planes. We see a very clear characteristic

spray pattern of the atomisers and the development of the spray with growing

distance. In the core of the two phase flow of the spray there are by far higher

values of drop number flux f^ caused by the entrainment effect mentioned before,

which shifts the smaller drops to the centre of the spray. Again the symmetry

of the spray pattern in the x and y directions as well (the view shown here is

only +y) . This is confirmed and the spray situation will become clearer from

Fig. 11 where the drop volume flux fg is seen. The peaks especially in E2 and

E3 are caused by the weighing effect in volume distributions: the comparatively

low number of larger drops at the spray boundary have a larger contribution to

the volume flux than the high number of small drops at the spray centre. Working

out the drop volume flux balance is a principle possibility given by the PDA

results. So in Fig. 11 the drop volume flow through spray planes is given as

well. But one should not forget that although PDA gives the results as absolute

values these results are affected by a number of error sources reaching from the

optical characteristics of the process fluid and bias effects caused by size

and/or velocity up to the signal processing which can cause a "complex" influence

on the results. So at least PDA results give values of important parameters in

218
sprays at different operating parameters for different locations as information

relative to each other so compensating potential bias influences. The balance

here is worked out based on a large number of measurement locations and on

results which are confirmed by a reference method (DSM) . This allows a more

extensive data preparation than without additional information. The high

knowledge of the process in question last but not least leads to the good fitting

of the balance calculation results presented here.

SUMMARY

PDA and DSM were used for a descriptive analysis of a spray of water. PDA

and DSM deliver intrinsically flux related or concentration related results which

are in principle convertible by using the drops velocity, however here measured

only by PDA. The spatial averaging effect of DSM is possibly not wanted. In

this case, the procedure of getting "point" results from different measurements

is necessary but also time consuming. If as in our case, the geometry of the

spray allows to have similar DSM-measurement volumes for different locations also

the DSM deliver a detailed analysis of the spray in terms of drop distributions

according size in a less time consuming procedure in comparison to PDA. The PDA

however allows a far more differentiated analysis of sprays in measuring also

drop velocities and because of its high spatial resolution. From this arises the

possibility of working out the drop size-velocity correlation as a quite

meaningful spray parameter and derived quantity like the drop volume flow.

REFERENCES

1. Bachalo, W.D., Houser, M.J., Smith, J.N., Atomisation and Spray

Technology, Vol. 3, pp. 53 (1987).

2. Bauckhage, J., Dannehl, M. , Fritsching, U. , Schone, F. , Schulte, G., 2nd

Intern. Cong, on Laser Anemometry - Adv. and Appl. , Strathclyde/UK (1987).

3. Dannehl, M. , Schone, F. , Schulte, G., Liquid Particle Size Measurement

Techniques: 2nd Volume, ASTM STP 1083, Hirlemann/Bachalo/Felton, Eds.

American Society for Testing and Materials, Philadelphia, pp. 246-256

(1990) . 4. Dodge, L.G., Rhodes, J., Reitz, R.D., Applied Optics, Vol. 26, No. 11,

pp. 2144 (1987).

5. Leschonski, K. , Part. Charact. , Vol. 1, pp. 89 (1984).

6. Meyer, P., Chigier, N. , Atomisation and Spray Technology, Vol. 2, pp. 261

(1986).

7. Raasch, J., Umhauer, H. , Chemie Ingenieur Tecknik, Vol 49, No. 12, pp. 931

(1977).

8. Schulte, G., ICLASS 85, LondonA^K, Proceedings (1985).

9. Schulte, G., Bachkhage, K. , Flogel, H.-H., Fritsching, U. , 3rd Intern.

Symp. of Appl. of Laser Anemometry to Fluid Mechanics, Lisbon/Portugal

(1986) . 10. Young, B.W. , Bachalo, W.D., in The direct comparison of three 'in-flight'

droplet sizing techniques for pesticide spray research, in Sym. Optical

Particle Sizing-87, pp. 42-1 - 42-15, Rouen, France (1987).

11. Zhu, H.M. , Sun, T.Y. , Chigier, N. , Atomisation and Spray Technology,

Vol. 3, pp. 89 (1987).

219
220
I

o o o o o o e o o o o o s e V n 01 o

O lO O lO O lO

o

CO

O O O o o > O lO o to

esi — »-

221
HI

CMo

0}

(0

e 0)

0) $ _ ^ i<

o i C

C

J 0)

o

O CM ? 5: o" o

CDO

EX « 5S S «I oc o

.2 CM = o

5:O

o

•• iS ^ t

O)

o

iZ

o

sX

222
V • 3324mm*3/s

V 3280mm*3/s

V • 3265mm*3/s

Fig.10: Drop number flux f^^ Fig. 11: Drop volume flux fg

Profiles from PDA-measurements Profiles from PDA-measurements

223

ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 20

NUMERICAL ANALYSIS OF PRESSURE SWIRL ATOMIZERS

C.-P. Mao and S.G. Chuech

Delavan Inc. West Des Moines, lA, U.S.A.

ABSTRACT

This study investigated the liquid breakup process in pressure swirl (simplex) atomizers. A "FILM" code

which combined a film dynamics model with wave instability theory was developed to describe the flow

characteristics and primary breakup process. This analysis led to the prediction of film trajectory, angle, breakup

length, thickness, velocity and drop size as a function of flow conditions, fluid properties and nozzle geometry.

Experimental data obtained with a high-magniflcation 4x5 camera, a Phase/Doppler Particle Analyzer, and a PMS imaging probe were used for model verification. The predicted Sauter mean diameters showed good

agreement with the PMS data.

INTRODUCTION

Pressure swirl atomizers are widely used in industrial, agricultiu'al and propulsion applications. These

atomizers are very popular because of their simple design, low cost and effective atomization.

In most pressure swirl atomizers, the liquid is injected under high pressure through tangential slots or passages into a small swirl chamber. Prior to discharge through a circular oriflce, the liquid forms a hollow fllm as a result of the vortex and air core developed within the swirl chamber.

As this fllm emerges from the exit oriflce, wave instability immediately occurs, leading to the

disintegration of the liquid sheet into ligaments and droplets. Following initial breakup, the ligaments and large

drops undergo a secondary breakup process. During this stage, many drops collide and coalesce. The

atomization process is Anally completed further downstream in the dispersed two-phase regime where spherical

drops exist. The characteristics of the liquid fllm are considered to be extremely important in determining spray

quality. A considerable number of studies have been conducted on the formation of liquid fllm and variations

in fllm thickness and velocity as a function of nozzle design dimensions and fluid properties. It is well known

that the initial axial and tangential velocities at the exit oriflce determine the shape and angle of the spray and

that the initial fllm thickness is related to drop size.

Despite some minor differences among previous theories [1-7], the investigator made important

contributions to the understanding of the flow phenomena in pressure atomizers. These theories permit the

prediction and correlation of initial fllm properties as a function of physical dimensions of the nozzles, fluid

properties and flow conditions. However, there is a common deflciency in most existing theories; namely, the

lack of information concerning the process of fllm breakup outside the nozzle. It is essential to determine the

fllm properties at the nozzle exit, but it is equally important to understand the breakup process. Spray

characteristics such as symmetry, uniformity, pattern, drop size distribution, velocity, and volume flux are closely

related to the fllm trajectory, breakup length, and fllm thickness outside the nozzle. Also, the interaction

between the liquid fllm and ambient air can be extremely critical in determining the downstream spray

characteristics. In the present study, a simple fllm dynamic model which, with wave instability theory, was developed to

describe the fllm characteristics outside pressure swirl atomizers. This analysis led to the prediction of film

trajectory, breakup length and initial drop size distribution. Experimental data obtained with a Phase/Doppler

Particle Analyzer, PMS imaging probe, and a high-magniflcation camera were used to verify the model

predictions.

225
EXPERIMENTAL APPARATUS

Three pressure swirl atomizers were selected for the study. These nozzles have similar designs, but

slightly different dimensions in the fuel distributor and exit orifice. Nozzle flowrates were measured on a

calibrated test stand equipped with turbine flow meters. All tests were conducted using MIL-C-7024 Type II calibration fluid having the following physical properties: density = 0.769 g/cc; kinematic viscosity = 1.15

centistoke; and surface tension = 28.2 dynes/cm. Each nozzle was tested within a 20 to 1380 kPa pressure range.

A high-magnification 4x5 format camera was used to determine the spray angle, film trajectory and

breakup length. The camera was equipped with a variable-length bellows extension to permit close-up views near

the breakup zone. All photos were taken with a back-lighting technique. Spray motion was frozen by means

of a high-intensity, short-pulse (0.5 /isec) flash unit manufactured by EG&G (Model 549). Kodak T-Max 400

film was selected for this study.

An Aerometrics Phase/Doppler Particle Analyzer (PDPA) was used to measure the drop size

distributions and mean diameters. A thorough calibration study using a single- drop generator and a rotating

disc was conducted to determine the accuracy and limitations of the instrument. For calibration, a piezoelectric

generator was used to produce uniform size drops ranging from 150 to 1000 microns.

During data acquisition, the PDPA instrument was used to obtain both local and global drop size

information. The atomizers were mounted on an automated traversing mechanism to allow movement at fixed

increments for the purpose of obtaining local point data. In addition, the traverse device could also sweep

continuously across the probe area from one side of the spray boundary to the other. Using this continuous

traversing method, it was possible to obtain a global mean drop size which closely represents the entire cross- sectional area. All drop size measurements were made 10 cm downstream from the nozzle face.

A Particle Measuring Systems (PMS) optical array spectrometer probe (OAP-260X) was utilized for

direct measurement of global mean diameters. The probe covers a range of 21 to 1875 microns in 62 chzmnels,

with approximately 30 microns per channel. A collimated laser beam exiting from a 2 mW He-Ne laser unit is directed perpendicularly across the main flight direction of the droplets. Droplets are sized using a linear device

with masked diode elements to sense the shadows projected on the diode elements by the moving droplets.

FILM DYNAMICS AND WAVE INSTABILITY THEORY

In the analysis of the primary breakup process, two basic models are required to describe the physical

phenomena. First, a film dynamics model is needed to simulate the motion of the liquid film as it emerges from

the orifice. Second, instability theory is required to describe the growth and propagation of waves on the film

surface.

The present study is concerned only with primary breakup and the mean diameters that represent the

entire spray. Other physical sub-models, such as secondary breakup, evaporation and drop history, were not

considered. The theoretical analysis focused only on the breakup regime.

Film Dynamics Model

In the present analysis, the hollow film formed outside the nozzle was assumed to be a steady-state

structure with no pressure gradient in the circumferential direction. Viscous forces within the thin liquid film were neglected. However, viscous friction at the interface was taken into account.

The governing equations for the hollo\y film were derived in a curvilinear coordinate system conforming

to the curvature of the liquid sheet. An integral method was used to establish the mass and momentum

equations. Primary forces acting upon the control volume included: convection, centrifugal motion, interface

friction, gravity, pressure gradients and Coriolis forces. The governing equations for the swirling film are described in detail in reference [8]. In the present model, the pressure differences between the inner and outer film surface were represented

by the capillary force (i.e., surface tension). The interaction between the liquid film and surrounding air was

calculated from the interface friction on both sides of the film surface. The friction coefficients are non-linear

functions of the Reynolds number based on the local film velocity and thickness. The three momentum

equations were solved simultaneously to yield information on the film velocity and trajectory (film shape and

angle). The variation of film thickness along the sheet surface was then determined from the mass continuity

equation.

226
Wave Instability Theory

Instability theory has been successfully employed to describe the breakup process in flow systems such

as liquid jets and fan sprays. From photographic studies, the theory of wave instability also seems to be suitable

for describing the behavior of a hollow film. The amplitude of the waves gradually increases as they progress

toward downstream along the film surface. Breakup occurs at a point where the wave ampUtude exceeds a

certain critical value.

Based upon Meyer and Weihs [9], a characteristic equation for hollow films can be derived from the

temporal growth rate of the unstable waves, according to the following equations:

Gja^+G^a = G3

where Gj = p/i+ipjk)

G3 = p/i/'-o^^ (1)

Equation (1) relates the growth rate to the local film properties and the surrounding air flow conditions.

It can be solved numerically to find all potentially unstable waves which, in turn, determine the breakup length

and drop size distribution.

In determining the breakup length, it is crucial to find the proper criteria for disintegration of a swirling

film. A spin frequency defined by the tangential velocity at the nozzle exit was found to be directly related to

the breakup process. Two characteristic parameters for determining breakup length are defined as follows to

determine breakup length:

(2)

fc = ^2/^2

(3)

K = V/c

It was assumed that liquid disintegration occurs when the wave amplitude exceeds one-half of the

characteristic wave length or one-half of the local film thickness.

When the fastest growing wave propagating in the streamwise direction meets the breakup criteria, a

circular ribbon of liquid detaches from the leading edge of the surface. Meanwhile, unstable waves keep growing

in the circumferential direction of the circular ribbon. Further breakdown occurs in the tangential direction as a result of flow instability and stretching.

One important assumption was made here for the prediction of mean drop size and distribution. It was

assumed that, after breakup, the drop size distribution would follow the same distribution function as that

determined by the unstable (positive) growth rates. A probability density function (PDF) can be constructed

from the solutions of Equation (1). This PDF was then used to calculate the drop size distribution and Sauter

mean diameter.

Methods of Estimating Initial Film Properties

Previous studies [10-11] indicated that the initial film properties are extremely critical in determining the

downstream spray characteristics. Hence, it is necessary to make an accurate estimate of the initial film

characteristics to ensure a valid prediction of the spray breakup process.

In the present study, two methods were used to estimate these initial parameters. (A) Giffen and

Muraszew's approach and (B) Lefebvre's correlation [12]. These two methods provided slightly different initial film properties.

Method (A) is based upon the concept that, anywhere inside the nozzle, the fluid mass, angular

momentum and total momentum are conserved. In the analysis by Giffen and Muraszew, the following equation

for the nozzle discharge coefficient was derived by assuming maximum flow through the exit orifice:

(4)

1+X

227
Because Q can be determined from flowrate measurements, Equation (4) may be solved for the value

of X (area ratio of air core and exit orifice). The film thickness can then be calculated from the following

equation:

After obtaining the initial film thickness t, the axial and tangential velocities are determined from angular

and total momentum balance.

The detailed description of method (B) for estimating the initial film thickness was presented in

reference [12]. Unlike other investigators, Lefebvre indicated that the initial film thickness does not remain

constant for a given nozzle, but depends on the operating pressure and fluid properties. He derived the following

equation for predicting the film thickness in simplex nozzles:

(6)

, = 2.7[M!^f]0.«

APp,

In method (B), Equation (6) was used to calculate the film thickness at the nozzle exit. The axial and

tangential velocities were then determined the same way as method (A).

Although the initial film thickness calculated by the two methods was slightly different, the results seem

to be in the same order of magnitude, and follow the same qualitative trends. It was observed that the initiid film thickness gradually decreased with an increase of the liquid pressure. As the pressure increased, however,

the air core began to develop and gradually increased in diameter. At a certain high pressure, the diameter

approached a constant value and the film thickness therefore also remained constant. As the pressure

approached 1380 kPa, there was better agreement between the two methods.

RESULTS AND DISCUSSION

Comparison of Initial Film Angle and Film Trajectory

By using 4x5 high-magnification photography, it was possible to observe the detailed variations in film

trajectory. A series of pictures taken at different liquid pressures was compared with the model predictions.

Present results indicated that the initial film properties estimated by method (A) agreed quite well with

the experimental data. The predicted trajectory gradually became wider and more coniced as the liquid pressure

increased.

Although the FILM code prediction agreed reasonably well with the actual film shape, the code cannot

predict liquid breakup at extremely low pressures. It predicted multiple "onions" that could not break up if the

Weber number is less than 4.0 x 10^. For lower Weber numbers, there is probably a different brezikup

mechanism which the present model does not take into account.

The most encouraging feature of the FILM code is its ability to predict the transition from an onion

shape to a fully developed conical film as the flowrate gradually increases. Hence, this computer model is

sufficient to describe the primary mechanisms at normal operating conditions.

Comparison of Breakup Length

Figure 1 compares the breakup length for predicted and measured values for the 2 and 10 GPH nozzles

at several liquid pressures. Breakup length was defined as the film length along the nozzle axis, not the film

surface. The values from method (A) agreed better with the experimental data. Within the pressure range

under study, it was found that the breakup length generally decreased with an increase in liquid pressure. Both

the experimental and predicted results approached asymptotic values. Thus, the breakup length will eventually

remain constant after a certain high pressure is reached. For the 10 GPH nozzle, the discrepancy between the

predicted breakup length from methods (A) and (B) diminished at liquid pressures above 1380 kPa.

The variation of breakup length appears to be closely related to the initial film thickness and angle. For

a given flowrate, short breakup lengths are associated with thin films. Wide film angles also result in a shorter

breakup lengths. The breakup length approaches a constant value at high pressures because of the constant film

thickness and velocity at the exit orifice.

228
2 GPH Nozzle

Aiiol Distance (mm)

Fig. 1 Comparison of breakup length between predictions Fig. 2 Variation of film thickness, streamwise and

and measurements for 2 and 10 GPH nozzles at tangential velocities with axial distance for

several fuel pressures. 2 GPH nozzle at several fuel pressures.

Variation of Film Thickness and Velocity

To date, experimental and analytical techniques can only provide information beyond the locations where

film breakup occurs. The accuracy of existing computer models depends on the upstream initial conditions

measured by the drop sampling instruments which normzilly have probe volumes located far away from the

breakup region.

The present FILM code ansdyzes the variation of film thickness, angle, and velocity in the film breakup

zone. This information is extremely useful because it permits more accurate prediction of drop behavior in the

dispersed two-phase regime.

Figure 2 shows the variation of film thickness, axial and tangential velocities for downstream locations

at several fiowrate conditions. The film thickness generally decreased with an increase in downstream distance.

The streamwise velocity accelerated with an increase in axial distance. However, the tangential (swirling) velocity

decreased. A common feature was observed in the variations of all three parameters. That is, these parameters

all tended to approach constant values prior to film breakup. It was also observed that, at higher pressures, the

film thickness and velocities changed at a faster rate as the downstream distance increased.

229
2 GPH Nozzle 2 GPH Nozzle

EI

10'

Fuel Pressure : 136 KPo

0.8-

U.D

c

c \

0.4-

c a \

^\ PDPA Meflsurenmt

0.2- Prediction

0.0- no 200 300

Drop Diameter (|un) 400

Drop Diameter ()unj

Drop Diameter Oimf

250

200

ISO

no

50

0

— Prediction PDPA Meosurement

0 PMS Measuretnent

a

0 a

200 , 400 600

Fuel Pressure (KPa)

5 GPH Nozzle

600

200 400 600

Fuel Pressure (KPo)

n GPH Nozzle

800

200 ^ 400 600

FudPresswe(KPi)) 800

Fig. 3 Comparison of predicted and measured drop size

distributions for 2 GPH nozzle at several fuel

pressures.

Fig. 4 Comparison of predicted and measured Sauter

mean diameters for 2, 5 and 10 GPH nozzles

at several fuel pressure conditions.

Comparison of Drop Size Distributions and Mean Diameters

Although methods (A) and (B) gave slightly different values of film trajectory, angle, and breakup length,

they provided similar drop size distributions and mean diameters. In the following discussion, however, the

predicted drop size values will be presented only for method (A).

Drop size distributions were calculated from a PDF constructed from solutions of the characteristic

equation. The PDF obtained at each flow condition was compared with the drop size distributions measured

by the Aerometrics Phase/Doppler Particle Analyzer (PDPA) and the PMS imaging spectrometer.

Because the drop size distribution predicted by the FILM code represents the particle spectrum for the

entire spray, the local mean diameters measured by the PDPA can not be directly used for comparison.

A continuous traverse method was used to measure the Sauter mean diameter along the center line

where the local point samples were taken. The intersection point between the radial SMD distribution and the

SMD measured by the continuous traverse will be determined as the location where the drop spectrum matches

that of the entire spray. Figure 3 shows a comparison of predicted and measured drop size distribution at several flowrate conditions for the 2 GPH nozzle. The predicted shape of the distribution function agreed quite

well with the PDPA measurements. Notice that the agreement appears to be much better at 138 kPa condition

where the peak point is fairly close.

230
Generally, the present model overestimated the percentage distribution of the large drops. In the

prediction, the size class of the peak point for the distribution function significantly dropped as the liquid

pressure increased. However, the peak point of the measurements did not drop significantly.

The predicted drop size distribution function was later used to calculate the SMD values for the

comparison of mean drop sizes. Figure 4 shows the comparison of SMD between predictions and measurements

at several liquid pressures for 2, 5, and 10 GPH nozzles.

Shown in Figure 4, good agreement was observed between PMS data and model predictions at liquid

pressures of 138 and 345 kPa conditions for 2 and 5 GPH nozzles. It appears that the agreement is better at lower pressure conditions for a low capacity nozzle. However, for a higher flow capacity nozzle, the agreement

is better at the high pressure conditions. This seems to indicate that the present model cannot predict well when

the initial film thickness is too thick or too thin. It could also be caused by the improper estimation of initial

film properties at those conditions used in method (A).

SUMMARY

The present study has demonstrated a realistic approach for analyzing the liquid breakup process and

flow characteristics in pressure swirl atomizers. Reasonably good agreement was observed between the model

predictions and experiments. The following major conclusions can be drawn from the present study:

(1) The initial film properties are extremely important in determining the spray characteristics. FILM code

predictions using the initial film properties estimated by method (A) showed better agreement with the

experiments.

(2) Prior to breakup, film properties such as thickness, axial and tangential velocities tend to approach to constant values. For a given nozzle, the gradient of film property variation as a function of downstream

distance increased with an increase in liquid pressure.

(3) In pressure swirl atomizers, the variation of breakup length is closely related to the initial film thickness

and angle. For a given flowrate, a thinner film thickness and wider film angle resulted in a shorter

breakup length. Similar to initial film thickness, breakup length tends to approach a constant value cifter a certain high pressure is reached.

(4) The present model prediction agreed better with the PMS data. The SMDs measured by the PDPA

were consistently lower than those of the model predictions and PMS data.

(5) The present analysis successfully predicted the primary breakup process in the pressure swirl atomizers.

Reasonably good agreement was obtained for the predictions of drop size distribution function.

NOMENCLATURE

discharge coefficient

Da orifice diameter

fc critical spin frequency for breakup

h half of film thickness at downstream locations

k wave number

M mass flowrate

AP liquid pressure drop

r radial distance in y direction

t film thickness at nozzle exit

U streamwase velocity in the f direction

Ua axial velocity at nozzle exit

W. Weber number, (p ,U2^D2/a)

tangential velocity at nozzle exit

X area ratio of air core and exit orifice

li liquid dynamic viscosity a temporal growth rate

P liquid density

a liquid surface tension

critical wavelength

231
SUBSCRIPTS

g gas phase

1 Hquid phase

r relative value between gas and liquid

2 location at the nozzle exit orifice

REFERENCES

1. Voros, I., Royal Hungarian University of Engineering, Ph.D. Thesis, Budapest, 1935.

2. Novikov, 1. 1., Journal of Technical Physics (Russian), XVIII, No. 3, 1948.

3. Doumas, M. and Laster, R., "Liquid-Film Properties for Centrifugal Spray Nozzles," Chemical

Engineering Progress, October 1953.

4. Giffen, E. and Muraszew, A., The Atomization of Liquid Fuels, Chapman «fe Hall Ltd.,

London, 1953.

5. Dombrowski, N. and Hasson, D., "The Flow Characteristics of Swirl Spray Pressure Nozzles

with Low Viscosity Liquids," AIChE Journal, Vol. 15, No. 4, July 1969.

6. Fenger, N. P., "Dimensional Analysis of Swirl Atomizers," Journal of Engineering, Vol. 216,

pp. 896-899, 1976.

7. Rizk, N. K. and Lefebvre, A. H., "Internal Flow Characteristics of Simplex Swirl Atomizers,"

AIAA 22nd Aerospace Science Meeting, Paper No. AIAA-84-0124, Reno, Nevada, 1984.

8. Chuech, S.G., and C.-P. Mao, "Numerical Analysis of Pressure Swirl Atomizers," Delavan

Technical Report No. 565, 1990.

9. Meyer, J. and Weihs, D., "Capillary Instability of an Annular Liquid Jet," Journal of Fluid

Mechanics, Vol. 179, pp. 531, 1987.

10. Mao, C.-P., Chuech, S.G., and Przekwas, A.J., "An Analysis of Pressure Swirl and Pure

Airblast Atomization," To be published in Journal of Atomization and Sprays, 1991.

11. Mao, C.-P., "An Analysis of Film Thickness Inside Airblast Atomizers," ILASS-Americas, 4th

Annual Conference, Hartford, Connecticut, 1990.

12. Lefebvre, A. H., Atomization and Sprays, Hemisphere Publishing Corporation, 1989.

232
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 21

THE SPRAY STRUCTURE FROM SWIRL ATOMIZERS

(Part 1: General Structure of the Spray)

N. Tokuoka, Y. Yamaguchi*, M. Takada and F. Zhang

Dept. Mechanical Engineering, Keio University, Yokohama, Japan

*Tonen Corporation, Tokyo, Japan

ABSTRACT

Spray structure is one of the major factors to control the combustion and

its simulation. Several papers have been reported in this field. But those

only engaged in the macro viewpoint because of the limit of measuring devices.

In this paper, micro behavior of drops such as the radial distributions of

the dropsize, velocity and volume flux etc. were measured by Phase Doppler

Particle Analyzer (PDPA) and the entrained air also measured to clarify the

general structure of spray from the swirl atomizer. The spray consists of

larger drops in the sheath, which have larger momentum and entrained ambient

air into spray, and smaller drops in the central part which are induced by air

flow. Therefore, mean drop velocity in the sheath decreases remarkably but one

in the center decays scarcely toward downstream. The volume rate of entrained

air is in proportion to axial distance Z and the spray momentum over a cross

section decreases with Z. The trajectory angle of drops is affected by its

size.

INTRODUCTION

The spray characteristics are of similar importance not only in

evaporation processes, combustions and chemical and food engineering etc. but

also on their simulation studies.

The relation between the atomizer and the spray characteristics have been

a fruitful subject for study. But in most researches on this subject, the

spray characteristics were clarified using a special atomizer.

The researches that disclosed the relations between the spray

characteristics and the geometrical shape of the atomizer were done by

Tanasawa( 1 ) , Kobayasi(2) (3) (4-) ,ect.

Tanasawa analyzed theoretically the stream in the swirl chamber as

potential flow and showed the relation between the geometrical size, the

discharge coefficient, and the spray cone angle. On the basis of Tanasawa 's

theory, Kobayasi designed and produced several atomizers for his systematic

experiments. The result of which was the establishment of the empirical

formula about the spray cone angle and the Sauter mean diameter (SMD).

Further, its accuracy was improved by examining surface tension and viscosity.

Nevertheless all of those studies were done in the macro viewpoint.

233
Recently, it becomes very important to know the structure of spray and

reciprocal actions between drops and air on the study of spray to establish

the requirement for energy saving and environment preserve.

The purposes of this research is to generalize the structure and

characteristics of spray from pressure atomizers and deliver the effective

data for the design of atomizers serving its purpose. At moment, swirl

atomizers are used. Centrifugal and single hole atomizers will be also

examined.

From this point of view, the spray structure and relations between the

dimensions of atomi zer, operating conditions etc., macro— and micro—scopic

characteristics such as spray angles, dropsize and drop velocity

distributions, trajectory, behavior of entrained air etc. will be made clear.

This paper (Parti ) makes clear the general characteristics of spray from

a swirl atomizer using only the typical atomizer

under constant operating conditions.

APPARATUS AND EXPERIMENTAL METHODS

Tested Atomizer

In this paper, the swirl atomizer is

used because it is used the most widely in

the continuous combustion. It is well known

that properties of atomizer are affected

by its dimensions such as diameter of inlet

holes into swirl chamber, orifice diameter,

diameter of swirl chamber and its shape.

The atomizer used in this research is

shown in Fig.1. It is designed for changing

the dimensions of each parts systematically.

The dimensions used in Parti are listed in

Tablel

.

Table 1 Dimensions and Flow Rate of Atomizer

Inlet holes Diameter 0.0011 m

Number 4- Swirl chamber Diameter 0.0104. m

Height 0.0030 m

Orifice Diameter 0.0007 m

Thickness 0.0010 m

Flow Rate 6.9cc/s at P=1 .0 MPa

COVER

SWIRL CHAMBER

CAB

SWIRL CHAMBER

ORIFICE PLATE

PACKING

CASE

Fig.1 The Structure

of the Tested

Swirl Atomizer

Experimental Apparatus and Measuring Methods

The schematic diagram of apparatus to produce a spray and to measure the

dropsize and its velocity is shown in Fig. 2. The sample liquid, stored in an

accumulator, is compressed by pressurized nitrogen gas and injected into a

still air through the tested atomizer. The liquid and nitrogen gas in an

accumulator are separated by a rubber balloon to prevent gas from dissolving

into liquid. The atomizer is set on a positioning device in an injection room

which has 1.200m width, 1.200m depth and 1.000m height. The top and bottom

walls of the room are constructed by honeycomb board and the room is slightly

blown out by a blower to prevent ambient air from circulating. So that the

234 «
spray is not affected by ambient

air. It is possible to move the

atomizer in the axial direction

and radial direction.

The dropsize and velocity

are measured simultaneously by

Phase Doppler Particle Analyzer

made by Aerometrics Inc..

The flow rate of entrained

air is measured by the 'porous

vessel' method which is similar

method by Ricou et al (5), as

shown in Fig. 3. (a) shows a

schematic diagram of apparatus

and (b) shows the details of the

porous vessel.

When a spray is injected

into the atmosphere, the ambient

air is entrained into it. This is

due to the pressure difference

between the inside of a spray and

a surrounding atmosphere.

If the pressure at points

beyond spray is kept atmospheric,

the spray characteristics are

similar to unbounded spray,

although it is formed in a closed

space. Base on this preceding

principle, a spray is formed in a

closed vessel which is

sufficiently larger than the

reduced pressure field and has a

permeable inner wall with a

circular hole exit at the front

wall of diameter such that the

spray just touched the perimeter

of the hole. All walls are porous

except those at the downstream

end where the aperture is

adjustable so that entrainment is

prevented and air is supplied

only from the sides and upstream.

The front wall was made of

transparent plastic, so that the

spray and its impact with the

perimeter of the hole could be

observed.

The static pressure was

measured at the point just inside

the permeable wall and outside of

the reduced pressure field by a

Betz type micro-manometer with a

sensitivity of 0./^9 Pa. During

operation of the spray, quantity

ii.i.iii..i.iiiiiiiiiiiiitiiii /\^\

mini R |- 09

PRESSURE

GAUGE

CITY WATER

Fig. 2 Schematic Diagram of Experimental

Apparatus to Measure the Dropsize

and its velocity

CITY UfiTER

AIR

-z—i

(a) Measurement of Entrainment

o.67om

(b) Details of Prorous Vessel

Fig. 3 Schematic diagram of Experimental

Apparatus to Measure the Rate of

Entrainment air

235
of air was supplied to the chamber through the permeable wall until the static

pressure inside the wall reached atmospheric. This amount of air supplied is

equal to the amount of air entrained through the spray surface between the

tip of the atomizer and the front wall. The total amount of entrained air' is

measured at Z = 0.020-0. 200m. The amount of air entrained through a spray

surface between any two sections can be obtained as the difference of the

total amount of air between those two sections.

Experimental Conditions , Measured Items and Analysis Method

Experimental conditions and measured points are shovm in Table 2.

Dropsizes and their velocities are measured at each point. The distribution of

mean dropsize, velocity, volume flux, momentum etc. in a space and at any

cross sections of spray are calculated from them. Trajectory angle and

velocity are obtained by measuring the radial and axial velocity separately.

Table 2 Experimental Conditions

Tested Atomizer Swirl Atomizer (see Table 1 and Fig.l)

Sample Liquid City Water

Pressure P = 1 .0 MPa

Measuring Device PDPA for Dropsize and its Velocity

Porous Vessel Method for Entrained Air

Measured Items Dropsize , Drop Velocity

and Flow Rate of Entrained Air „ Measured points Axial Distance Z=20,40,60,80, 1 00 (xlO'-^m)

Radial Distance R=0-0.100(m) l-5(x10 '^m)pitch

EXPERIMENTAL RESULTS AND DISCUSSION

Volume Flux Distribution and Spray Angle

Volume flux distribution in the radial direction at any axial distance Z

are shown in Fig./^.. The volume flux in the central part of spray becomes

lower at any Z, increases towards the boundary surface and decreases again

2.0

^5 1.5

1.0

Io

3

qjO.

I

1 i\

\

O

1 = 0.020rT O.CMOm

:

0.060^1 ® Z = 0.060m: 3 Z = O.IOOn

0 10 20 30 40 50 Radial Di stance [xw'^m]

60 70 10 20 30 40 50 Radial Distance [xio'^m]

60 70

Fig./4 Volume Flux Distribution Fig. 5 SMD Distribution

236
after reaching the maximum value. The radial point showing the maximum volume

flux is defined as the spray sheath, and the radial distance as the sheath

radius of spray, Rsheath. The spray angle is defined by the outline of spray

from photograph in general. However, it is difficult to decide the outline

exactly. The half spray angle, a'/2, in this research is defined as the angle

constructed by Z axis and the line connected the orifice and spray sheath, viz

a ' =2-^ar ctan ( Rsheath/Z )

.

The spray angle a' defined in this research shows 4-8 degree between

Z=O.OiiOm and 0.100m (see Fig.8). The volume flux at Z less than 0.04.0m can not

be measured by PDPA because drops in this region are not spherical. According

to the observation by photographs, disintegration completes around at Z=0.020m

and deformation of drops follows it up to Z=0.030m. The spray cone angle

(angle of liquid sheet) shows several degrees larger than the spray angle.

SMD and Mean Velocity Distribution

Figure 5 shows the distribution of Sauter Mean Diameter (SMD) in the

radial direction at any Z. SMD increases towards the sheath at any section and

reaches maximum in the sheath. It is also known that SMD in the sheath

increases with Z. This is due to the small drops being induced by entrained

flow, leaving the larger drops which have larger momentum moving linearly

along the sheath.

The mean velocity shows maximum near the sheath and it decays with

increasing Z as shown in Fig. 6. The mean velocity here is calculated

arithmetically. This tendency is similar to the SMD distribution. On the

other hand, the decay of the velocity in the center of the spray is smaller

than one near the sheath. It is considered that ambient air is accelerated by

the drops near the sheath. Smaller drops flow with the entrained air and

concentrated in the center of the spray. The decay of small drops in the

center of spray becomes less as they are accelerated by entrained air.

40

35

30

1^25

f,20

0

£io

• Z = 0.020-11

;

O Z = 0.040m4 ! e z = o.ceom

i

® Z = 0.080m! i O Z= 0.100m i

Qi,„^gf:::y,. j

:

10 20 30 • 40 50 Radial Di stance ixio-^m]

60 70 10 20 30 40 50 60 Radial Di stanceixio'^m]

Fig. 6 Mean Velocity Distribution Fig. 7 Momentum Flux Distribution

Momentum Flux Distribution

The momentum flux distribution of drops is shown in Fig. 7. The momentum

flux is calculated by the mass and velocity of independent drop. This tendency

shows more similar to one of the volume flux than to one of the mean velocity

because the momentum is largely influenced by the larger drops with larger

mass and the mean velocity is calculated arithmetically.

237
Axial Distributions

The SMD and the momentum over each cross section are calculated byassuming that the spray from a swirl atomizer is Z axial symmetry. The axial

distribution of SMD, momentum and flow rate of entrained air are shown in

Fig. 8.

The theoretical SMD is calculated by means of the empirical formula (3)

and the theoretical momentum is obtained from the axial velocity by

Tanasawa(l) multiplied by the flow rate at the orifice. SMD at each cross

section except Z=0.020m becomes constant and coincides well with the empirical

formula by Kobayasi(3). SMD at Z=0.020m is smaller than others. In this

region, disintegration has not completed yet and the value is calculated by

only spherical particles at the point.

The total amount of entrained air in volume increases proportional to

the axial distance Z , but the total momentum of drops in cross section. Ml,

is inversely. The induced air velocity just before entering the sheath Vin is

in inverse proportion to Z because the volume fraction dQa entrained from dZ,

dQa/dZ, is constant and surface area at dZ is proportional to Z.

This figure shows that the spray angle at Z=0.020m is larger than one for

Z>0.02mm and it becomes constant at Z>0. 04.0m. This fact also suggests that

disintegration has not completed yet at Z=0.020m.

20

15

m 10

'o

X

m

C3

60

Z [XlO^m]

2.0

1.6

I/)

1.2 e

OB

-0.4

Jo

Fig. 8 Axial Distribution of SMD, Ml, a' and Qa in section

Drop Velocity and Trajectory Angle

Figure 9 shows the drop mean velocity and the trajectory angle of each

dropsize range at local points. The trajectory angle is calculated by using

the mean velocity of each dropsize in radial direction and one in Z axial

direction. It is considered that small drops flow with entrained air so the

velocity of entrained air can be known roughly from the _gelocity of small

drops. From the figure for the dropsize range of 0-10 xlO m, the air flows

faster at the central part than at the outer part. It is observed that the

238
decrease of drop velocity near the spray axis is smaller than that at the

sheath. And an ambient air seems to be entrained at Z=0. 0/^0-0.060 m

vigorously. Further, the velocity of large drops is faster at the outer part

and slower at the center than one of small drops. In the sheath, the number of

small drops having low velocity decrease and the number of larger drops having

high velocity increase with Z, because of the total amount of entrainment

increase with Z as shown in Fig. 8.

I

! I 0-10[xl0"^m]

Z = 0.020 m —

I

—In Z = 0.040.m m

Z = 0.060iT>_

Z = 0.080 ^

Z = 0.100

30-40C xlO"^] -90[xlO-6m] 130-14C[xlO-6m]

—+-

Fig. 9 Number Mean Velocity and Trajectory Angle

Figure 10 shows the trajectory angle distribution of drops at Z=0.060m

referring to the size ranges. Though all drops flow parallel to the spray axis

in the region of less than R=0.010m ^ut at the outer part the trajectory angle

of small drops of less than 4-0x10 m becomes smaller than the spray angle.

Large drops flow along the sheath, that is the half spray angle.

40

30

o

o

y 20

10 <

OIL

10

1 — 1

I -- 0.060 m

100-110 XlO"^m

\ r

A"^, 130-140 xlO"^m

40-50 xlO-^m

\ 10-20 XlO'^m

: ^ 6 : :

: 70-80 xlO" m ; ;

10 10 20 RsHEATH* 30

Radial Distance [x10"^ m ]

40

Fig. 10 Trajectory Angle Distribution of Each Dropsize

239
CONCLUSIONS

In this paper, distributions of mean dropsize, velocity, volume flux,

flow rate of entrained air, trajectory angle etc. were measured by using the

typical atomizer under constant operating conditions to make clear the general

characteristics of spray from the swirl atomizer and to generalize the

structure. The results are summarized as follows;

(1) The spray consist of large drops with large momentum in the sheath and

small drops induced by an entrained air in the central part.

(2) The mean drop velocity at the sheath is larger than one in the central

part of spray.

(3) The decrease of the mean velocity in the central part is smaller compared

with one in the sheath. This fact suggests that drops accelerate

entrained air at the outer part of spray and entrained air does drops at

the central part.

iU) The trajectory angle of drops depends upon the dropsize. Though all

drops flow parallel to the spray axis in the central part, at the outer

part, the trajectory angle of small drops is smaller than large

drops.

(5) The total amoimt of entrained air is in proportion to axial distance Z

and the induced air velocity before entering the sheath is inversely

proportional to Z.

(6) The total momentum of drops over a cross section is inversely

proportional to axial distance Z.

ACKNOWLEDGMENT

The authors would like to express their deep gratitude to Mr. T.Hyodo,

Mr. K.Fukuzato from Seika Sangyo Co. Ltd. and staffs of Aerometries Inc. who

offered valuable advises for this study.

REFERENCES

(1) Tanasawa Y. et al., "Design Method For Swirl Injection Nozzle Using

Potential Flow Theory", JSME, Vol 17, No. 58, pp. 107-1 11 (1951).

(2) Kobayasi K,, "Atomizing Characteristics of Swirl Injection Nozzle (1st

Report)", JSME, Vol 17, No. 58, pp.ll2-lU (1951).

(3) Kobayasi K., "Atomizing Characteristics of swirl Injection Nozzle (2nd

Report)", JSME, Vol 17, No. 58, pp.lU-118 (1951).

(4-)Kobayasi K., "Atomizing Characteristics of Swirl Injection Nozzle (3rd

Report)", JSME, Vol 19, No. 80, pp,60-6A (1953).

(5) Ricou, F.P. and Spalding, D.B., "Measurements of entrainment by

axisymmetrical turbulent jets", J. Fluid Mech., Vol 11, No.l, pp. 21-32

(1961).

(6) Benatt, F.G.S. and Eisenklam, P., "Gaseous entrainment into axisymmetric

liquid spray", J. Inst. Fuel, Vol A2, No. 8, pp. 309-31 5 (1969).

240
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 22

THE SPRAY STRUCTURE FROM SWIRL ATOMIZERS

(Part 2: Effect of Ii\jection Pressure and Atomizer

Characteristics on the Spray Structure)

F. Zhang, Y. Yamaguchi", M. Takada and N. Tokuoka

Dept. Mechanical Engineering, Keio University, Yokohama, Japan

Tonen Corporation, Tokyo, Japan

ABSTRACT

It is knovm that the spray characteristics change according to the

structure of atomizer and operating conditions. The swirl atomizers whose

dimensions of each parts can be changed systematically are designed and the

relations among atomizer's characteristics, operating conditions and spray

structure are made clear experimentally.

The boundary of the sheath region and central part of spray becomes

clearer and the thickness of sheath region decreases with an injection

pressure. The flow factor B and the characteristic factor K affect on a spray

angle. A spray angle increases with the increase of B and with the decrease of

K. Empirical formula for spray angle a' is presented as a function of Re. The

flow rate of entrained air is also affected by B, K and injection pressure.

INTRODUCTION

Spray characteristics are changed by the structure of atomizer, its

dimensions, operating conditions, liquid properties, ambient air conditions

etc.. It is known that the characteristics of swirl atomizer are affected by

its dimensions such as an area of inlet hole into a swirl chamber, a diameter

and a height of swirl chamber, orifice diameter and its length etc..

Many studies(l )-(3) have been reported about the relations between the

structure of atomizer and characteristics of spray. However, those made clear

only macro characteristics of spray such as spray cone angle, Sauter mean

diameter, discharge coefficient, etc. because of the lack of measuring

technique. Fortunately, a phase doppler system is development and it becomes

possible 'to measure the velocity and size of each drop simultaneously. This

study aims to generalize the structure of spray from the microscopic point of

view.

The swirl atomizers of which dimensions of each parts can be changed

systematically are designed. The general structure of spray from a swirl

atomizer was made clear in Parti (4-). In this paper, the relations among the

characteristics of atomizer, an injection pressure and spray characteristics

such as spray angle, spatial distributions of mean dropsize, velocity,

momentum and volume flux, trajectory angle of drops, flow rate of entrained

air etc. are examined to generalize the characteristics of spray.

241
EXPERIMENTAL METHODS AND APPARATUS

The details of experimental apparatus, measuring item etc. are shown in

Parti (^). The velocity and diameter of drops, volume flux are measured by

Phase Doppler Particle Analyzer (PDPA) from Aerometries Inc. and the flow rate

of entrained air by the "porous vessel" method(5).

The dimensions of each parts of tested atomizers(4-) are shown in Table 1.

Those are designed to be able to examine spray characteristics under the

constant characteristic factor K, the flow factor B, etc. by selecting some of

atomizers. The city water is sprayed into a still air under an injection

pressure of 0.7, 1.0 and 1.5 MPa.

Table 1 Dimensions and Properties of Atomizers

Atomizer Measured Values Theoretical Values by (1)

Number do di

[xlO"^

de K ao k dc B

] [-] [deg] [-] [10 -^m] [-]

S-01 1.28 10.

A

0.96 0.66 0.61 0.58 0.49

S-02 1 .28 10.

A

0.70 0.90 65.3 0.54 0.38 0.49

S-03 1 .28 10.

A

0.51 1 ,2k 55.9 0.47 0.24 0.A9

S-OA 1.10 10.

A

0.96 0.A8 83.2 0.67 0.64 0.42

S-05 1 .10 10.0.70 0.66 74.1 0.61 0.-^2 0.42

S-06 0.93 10.

A

0.96 0.35 92.2 0.72 0.69 0.36

S-07 0.93 10.^ 0.70 0.4-8 83.7 0.67 0.47 0.36

S-08 0.93 10.4- 0.51 0.65 74.7 0.61 0.31 0.36

ao: Theoretical spray cone angle (=2*arctan(k/(1-k**2)**0.5) )

; B:Flow factor

(=2^m**(1/2) -dO/di) )

; dc:Air core diameter at the orifice; de: Orifice diameter

di: Diameter of swirl chamber; do: Inlet hole diameter to swirl chamber;

k: Cavity factor (=dc/de); K: Characteristic factor (=m*do*^2/(de''^di)

)

RESULTS AND DISCUSSION

Effect of Injection Pressure

Distribution of SMD and Mean Drop Velocity . The difference of SMD

distributions under various injection pressure Pi is shown in Fig. 1(a) and of

number mean velocity in Fig.l(b). Both are at an axial distance Z=0.060m

and an atomizer is S-05.

SMD decreases with Pi at each radial point and the radial distance where

SMD shows the maximum increases with Pi.

The mean velocity decreases with radial distance R and increases towards

the sheath after showing minimum. Those tendencies do not affected by Pi but

the velocity difference of drops along R becomes more distinct with Pi being

higher. It is considered that the increase of drop velocity near the sheath is

due to the increase of initial injection velocity under higher Pi but one in

the central part is due to higher entrained air velocity caused by the

decrease of pressure within a spray led by larger momentum of drops at sheath.

Distributions of Volume Flux and Momentum Flux of drops . Fig.1 (c) and (d)

shows the distributions of volume flux and momentum flux of drops at Z=0.060m

respectively. Those indicate that the maximum values and radial distances

showing maximum increase with the injection pressure Pi. The fact means that

the spray angle increases with Pi. The distributions near the sheath becomes

more peaky and the thickness of sheath region reduces with increasing Pi.

242
150-

125

vo^OO

I.o

Pressure [MPa]

O 0.7 ~

1.0

1.5

Radial Distance [xlO m]

(a)

Radial Distance [xlO m]

(b)

Fig.1 Effect of Injection Pressure on Distribution

The total momentum of drops and the

flow rate of entrained air are shown in

Fig. 2. The latter means the total volume

of air entrained through spray surface

between Z=Om and Z=0.060m. The flow rate

of entrained air is roughly proportional

to the total momentum Ml when an

injection pressure increases.

Effect of Dropsize on Velocity

Distribution . The mean velocity and

trajectory angle of each dropsize are

shown in Fig. 3. The line length in Fig. 4- shows the mean velocity and the

20 r

m c^5

a

Io

xlO

a)

Z = 0.060 m

0.5 0.7 0.9 1.1

Injection Pressure

1.3 1.5

Pi [MPa]

Fig. 2 Effect of Pi on Qa and Ml

243
direction a trajectory angle. When the injection pressure ingreases, the

velocity also increases. The area where drops og more than 80X10" m decreases

with Pi though one for drops less than 4-0X10 m does not change and a path

width passing through large drops becomes narrow. This fact also indicates

that the thickness of sheath region reduces with increasing Pi.

0-10 [xlO ^m] 30-40 [xlO" m]

Llj

I 20 m/

80-90 [x10"^m] 130-U0[x10 ^m]

LiJ

o

II •H

Z = 0.020 m

Z =7 0.060 m I

Z = 0.080 m I llllliiiiiiiuiKnn"- ~jl[THinnnnin>inTM"""

Z = 0.100 m

s

}]||||lllllininniiiiiiiiin|Hilll —

I

mm

11I11P1!»

—II ,->TOTO

—

+

—-1-

"11

—

I

•H

lLi

I Z = O.O^Oji

Z = 0.040 _m

Z = 0.0^ _m

Z = 0.080 ja

I Z = 0.100 m —]jjII[Hmnil'"ininMnt'""'''

I

—

I

cd

Z ^0:,020_!n

' " ' Z = 0^40_in

Z = 0^060jn

•H -H Z = 0.08p_m

Z = 0.100_m

1^

1

—

I

-^

1^

—

f

mm"

Fig. 3 Number Mean Velocity and Trajectory Angle

Effect of Atomizer ' s Characteristics

The spray produced by different atomizers with various dimensions of each

parts are examined under the constant injection pressure of Pi=1.0MPa to

clarify the effect of atomizer's characteristics on spray.

Effect of Orifice Diameter de. To make clear the effect of an orifice

diameter on spray characteristics, the sprays from atomizers with various

244
25 Pi = 1 .0 MPa; B = 0.4.9

do = 0.00128in; di = O.OIO/^m

de[m] K

O 0.00096 0.66

€i 0.00070 0.90

0.00051 ^.2A

Z=0.060m; Pi=1.0MPa; B=0.ii9

cy 0 .50 0.70 _^ 0.90

de [10 ^m]

150 Fig. 5 Effect of Orifice Diameter de

on SMD, Qa, Ql, Ml, to

orifice diameter de but constant

inlet hole diameter do and diameter

of swirl chamber di are examined.

In the case that do and di are

constant and de changes, the

characteristic factor of atomizer K

is in inverse proportion to de and

the flow factor B is constant.

Figure 4- shows the

distributions of mean velocity, SMD

and volume flux at the axial

distance of Z==0.060m. The flow

factor B is 0.49 and K changes

between 1.24 and 0.66. Those

figures suggest that the maximum

values of mean velocity, SMD,

volume flux, spray radius, and

spray angle a' increases with the

increase of orifice diameter de,

that is with the decrease of K.

This is because the radial velocity

component at the orifice increases

relatively compared with the axial

velocity component when de is

larger, viz K is smaller.

SMD(D32) over a cross section

increases with the increase of the

orifice diameter. It is also

predicted by a theoretical work(1).

According to it, the thickness of

liquid film at an orifice,

to, increases with K and shown in Fig. 5 with SMD.

The ambient air is entrained by the momentum of drops flowing in the

sheath region and. Therefore, the flow rate of entrained air increases with de

o

10 20 30

Radial Distance

AO

[x10"\]

50

Fig./i Effect of Orifice Diameter

on Radial Distribution

245
because the momentum of drops Ml in the

sheath region Is larger when de Is

larger as shown in Fig. 5.

The number mean velocity in the

sheath of all atomizers are almost same

at Z=0.060m. This is because that the

reduction rate of velocity is smaller in

the sheath region because the size of

drops there are relatively large. On the

other hand, drop velocity in the central

part of spray becomes smaller with de.

It is considered that the mean velocity

of air within a spray decreases with de,

as the spray radius becomes larger

though the Qa increases.

Z=0.060in; Pi=1.0MPa; K=0.66

I^100

Z. 90 Q

^ 80

I

120

015

Effect of Flow Factor B. The flow

diameter of an

chamber. To make

a

a

factor is affected by

inlet hole and of swirl

clear the effect of B, atomizers with

constant characteristic factor K and

I

JO

5-

required.

of do and

and those

OL"

0-3 0.1+ Flow Factor B

I 0.5

Fig. 6 Effect of Flow Factor B

on SMD, Qa, Ql, Ml

different flow factor B are

Atomizers with suitable sets

di are selected from Table 1 are S-01 ,-05,-08.

The flow rate of liquid Ql, the

flow rate of entrained air Qa, SMD etc.

are shown in Fig. 6. All of them increase

with the increase of flow factor B. The

spray angle a' also increases with B

because the radial distance indicating

the maximum of volume flux increases

with B as shown Fig. 7.

According to the potential theory,

the characteristics of spray depend upon

K but the viscosity affects on the flow

in a swirl chamber when real liquid is

injected. Tanasawa(l) also suggested

that the flow in a swirl chamber

approximates the potential flow most at

B=0.5. When B is less than 0.5, a

friction due to the viscosity near the

wall increases and the swirl velocity

decreases. This fact leads a spray angle

decreases and a discharge coefficient

increases. On the other hand, when B is greater than 0.5, the radial and axial

velocities are affected a little by friction and the discharge coefficient

became smaller while the spray angle became larger.

In case of S-01 ,-05,-08, though the theoretical cone angles calculated by

the potential theory are all 75 deg. but spray angles measured from Fig.

7

become 60, 50 and 32 deg. respectively. This result indicates that the spray

angle decreases with the decrease of B.

CO

B

? 6

o

x3

3

i-l

(x<

<D 1°

O

>

B

0.36—o—

- 0.A2—b

0.^9-0—

1

/• A

10 20 30 i^O Radial Distance [x10~^

50

m]

Fig. 7 Effect of Flow Factor B

on Volume Flux

Effect of Inlet Port Diameter do. The sprays from atomizers, of which an

inlet hole diameter do varies and diameter of an orifice de and of a swirl

246
Fig. 8 Effect of do On SMD,Qa,Ql,Ml, to

chamber di are kept constant, are

examined. In this case, both the

flow factor B and the

characteristic factor of atomizer K

change, K is in proportion to the

square of do and B to do.

SMD(D32), total momentum of

drops and flow rate of entrained

air Qa over a cross section

increase with an inlet diameter do

as shown in Fig. 8. This figure

indicates the tendency for K shows

opposite compared with Fig. 5. It is

considered that an orifice diameter

de does not give any effects on B

but an inlet hole diameter do

affect on not only K but also B.

When do becomes smaller as in this

case, K increases and B decreases.

The decrease of K causes the

increase of the ratio of the radial

velocity component to the axial one

and makes a thickness of liquid

film thicker but the decrease of B

acts on them conversely.

Consequently, the relations in

Fig. 8 are obtained.

Figure 9 shows the

distributions of mean velocity, SMD

and volume flux of drops in radial

direction. It is evidence from

Fig. 9 that the mean velocity and

SMD increases with do and that the spray radius is

Therefore, the spray angle vary scarcely with do.

aH

O

>

0 10 20 30

Radial Distance [xlO ^m]

Fig. 9 Effect of Inlet Port Diameter do

on Radial Distribution

not influenced by do,

247
Empirical Formula for Spray Angle a

Figure 10 shows the relation

between the ratio of a spray angle a' to

a theoretical spray cone angle ao and Re

number in the swirl chamber. Reynolds

number is given by using an inlet

velocity into swirl chamber, ui,

calculated from the potential theory as

the representative velocity, a radius of

swirl chamber, and viscosity of liquid.

It may not suitable in this stage for

experimental results to indicate by Re

number because all data are obtained

from sprays of city water and

viscosity, V, is constant. However, Fig

formula by means of the least square meth

i

1

j

vy 1 O SW1F1.02 0 SWIR.03 y

0 SWIR.W 0 Wlfl-05 g $wira.06

1x10" SxlO" IxIO" 2i!0"

Rethoios NuH«t«

Fig. 10 Relation Between Re and a '/a

.13 gives the following empirical

od. a'=ao*e^*(-3900/Re) ;Re=uiMi/2/v.

CONCLUSIONS

The effect of injection pressure and dimensions of parts constructing a

swirl atomizer on the spray characteristics are examined. Size and velocity of

each drop are measured by PDPA and the entrained air into spray is also

measured using the 'porous vessel' method. These results summarized as follows

(1) The volume flux and momentum flux are concentrated near the sheath and this

tendency becomes more remarkable and a thickness of sheath region decreases

with an injection pressure.

(2) The flow factor B and the characteristic factor K affect on a spray angle.

A spray angle increases with the increase of B and with the decrease of K.

(3) The increase of orifice diameter leads the increase of spray angle and the

decrease of thickness -of sheath region. However, the inlet hole diameter

effects scarcely on them. This is because the former affect only on a

characteristic factor K but the latter on K and a flow factor B.

(4) The increase of B and/or the decrease of K and/or the increase of injection

pressure cause the increase of flow rate of entrained air.

(5) Empirical formula for spray angle a' in the researched range is led as

a' = ao*e-^^-(-3900/Re) ; Re = ui*di/2/v

ACKNOWLEDGMENT

The authors would like to express their deep gratitude to Mr.T. Hyodo,

Mr.K. Fukuzato from Seika Sangyo Co. Ltd and staffs of Aerometries Inc. who

offered valuable advice for this study.

REFERENCES

(1) Tanasawa Y., et .al. ; "Design Method For Swirl Injection Nozzle Using

Potential Flow Theory", JSME, Vol 17, No. $8, pp. 107-111 (1951).

(2) Kobayasi K. ; "Atomizing Characteristics of Swirl Injection Nozzle (1st

Report)", JSME, Vol 17, No. 58, pp.112-1U (1951).

(3) Kobayasi K. ; "Atomizing Characteristics of swirl Injection Nozzle (2nd

Report)", JSME, Vol 17, No. 58, pp.1U-118 (1951).

(/i)Tokuoka N. et.al.;"The Spray Structure from Swirl Atomizer (Part 1)",

submitting to 5th ICLASS (1991).

(5)Ricou F.P., Spalding D.B.; "Measurements of entrainment by axisymmertrical

turbulent jets", J. Fluid Mech., Vol 11, No.1, pp.21 -32 (1961).

248
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 23

TWO-COMPONENT PHASE-DOPPLER MEASUREMENTS FROM THE SPRAY

OF A SWIRL-PRESSURE BURNER-NOZZLE IN THE VICINITY OF

THE NOZZLE

K. Lampe, K. Bauckhage and B. Haase

University of Bremen, Verfahrenstechnik

Bremen, Germany

ABSTRACT

A two-component phase-Doppler measuring system (PDM) has been used to measure

simultaneously drop sizes and velocities at a vertically mounted swirl-pressure burner-nozzle

under various atomizing conditions. The measurements from different radial positions in two

horizontal planes, 10 mm and 20 mm below the nozzle, result in a linear dependency of the

main horizontal velocity components on radial distances r. With a second one-component PDM

variations of Sauter-mean-diameter (SMD) were examined in order to describe the dependency

of SMD on the fuel oil viscosity.

INTRODUCTION

Short comings of energy resources and higher requirements for minimizing air-pollutant

emissions call for further optimization of the atomization process in fuel oil burners. Simplex

swirl pressure nozzles producing hollow or solid spray cones are used for instance in househeating systems. Only with the exact knowledge of the spray parameters a further optimization

of the already well doing atomizing process of these burner nozzles is possible. Using the

phase-Doppler technique as described by [1] to [4] sizes and two velocity components of droplets

can be measured simultaneously thus answering questions for local distributions and evolutions

of sprays.

Simplex nozzles of fuel oil burners generate diffusion flames, in which beside the drop

size distribution the fuel oil air mixing ratio plays an important part. Thus operating conditions

as the pressure and the temperature of the preheated fuel oil get influence on the atomization

process and on the properties of the sprays of this type of swirl-pressure burner nozzles.

Therefore our point of interest was to find out the dependency of the local spray cone

parameters, especially of the Sauter-mean-diameter on the oil pressure, the temperature and on

the temperature-dependent physical fuel oil properties (viscosity, density, surface tension)

respectively. All investigations have been done under temperature controlled conditions using

only one type of hollow-cone-nozzle, i. e. Danfoss, type: 60° H, 1,6 kg/h. Another question was

the dependency of the horizontal (radial and swirl) velocity components of drops on radial distance r from the centreline of the spray cone and on the axial distance z from the nozzle. Thus

this dependency had also to be analysed in relation of the temperature dependent properties of

fuel oil to the spray cone parameters. At least one object of our investigations was, to point out

a relationship between radial and swirl velocity components and the radial distance r at different horizontal planes.

249
Experimental Set-up

The experimental apparatus consists of the nozzle mounted centrally in a cylindrical

chamber. The chamber is movable in two directions (x- and y-direction), the nozzle position

can be changed in z-direction (Fig. 1). Without moving the optics of the measuring system thus

each point of spray cone can be reached for measurements without changing the operating conditions of the stationary atomization process. The oil pressure, pj, and its temperature, T,, can

be varied and controlled in a range of 5 to 20 bars and from 20 to 90 "C. The circulated and

rectified gasstream (N2, u ~ 0,2 m/s) is temperature controlled. It is prevented from supplying

the measuring volume with recirculating fuel oil mist by filters. In order to avoid explosion N2

was used instead of air.

(uel-oil-supply

\ fuel-oil-preheater

honeycomb rectifier

window

patternator

OAJ

gas-circulation

I i I i

I M I I

gas-preheater

nozzle

I

I M

oI I I I

i I iTTTT

J] A

"^vj^fuel-oil- separator

Fig. 1 Experimental set-up

Before entering the nozzle the fuel oil temperature and pressure are measured by thermocouples and pressure sensor respectively. A patternator can be installed below the nozzle,

measuring the dependency of mass flow density on radial distance, r, axial distance, z, and on

operating conditions as oil pressure, Pj, and its temperature Tj.

Measurement Procedures

The temperature-dependency of the local drop size distributions and the local distributions of the vertical component, v^, of droplet velocities has been measured separately for

various locations in the spray cone by means of an one-component PDM (Aerometrics). The

main interest was focused on the changes of SMD by varying the temperature dependent physical properties of the fuel oil (and additional of the gas stream).

All two velocity component measurements together with the correlated diameter observations have been done by a four-beam two-colour configuration [5] of a fibre-optic-PDM

(Dantec). For both measuring systems (Aerometrics and Dantec) an argon-ion-laser was used as

light source; photomultipliers, installed under 30° off-axis-angles were used as receivers.

The measuring system of the two-component-PDM is shown in Fig. 2. After passing a

set of optical filters the light is divided into two beams with different wavelengths (blue: 488

nm, green: 514,5 nm) and after that by beam-splitter prisms into the well known dual beam systems. In order to distinguish between the altering directions of the horizontal velocities of very

low values a frequency shift was necessary, realized in the path of the blue light using a bragcell.

250
Fig. 2 Fiber optic phase-Doppler-measuring-system

Signal detections and data acquisition of the two-component PDM were done by a

transient- recorder and a PC-diagnostic-system. The advantage of this system is based on the

ability of logging a whole signal run, which is laid down in a data file. After a great number

(of more than 2000) signal runs has been stored to the buffer of the computer, the quality of

the signals is tested by varying the fault criteria [6]. Test measurements with a small rotating

glas-ball resulted in good accordance with the given values of the rotating speed of the drive

motor and with the particle size given by microphotos. Instead of the transient recorder also a FFT-processor, described by [7], has been used for detecting and analyzing the signals. The

comparison of these acquiring methods resulted also in good accordance.

Discussion of the experimental results

An optical configuration for laser light sheet photography had been used to get first informations about the shape and the extension of the spray cone. Fig. 3 shows a photograph of

the spray cone in the vicinity of the nozzle, received for the same fuel oil and gas temperature

of 30 "C and a fuel oil pressure of 8 bar.

30 -20 -10 0 10 20 30

radial distance r/mm

Fig. 3 Laser-light-sheet photograph

251
It has to be noticed that the extension of the liquid film and of the hollow cone in the

centre of the spray cone is very small. Within less than 5 mm below the nozzle the liquid film

becomes unstable and disintegrates into ligaments and droplets. 10 mm below the nozzle orifice

the hollow cone has disappeared and the shape of the spray cone looks like a solid cone. This is illustrating that the interesting region for drop-size and velocity measurements should be ex- ecuted not far away from the nozzle orifice. It has also to be noticed that changes of the spray

parameters which influence the burning process (with regard to the very short flame) are ex- pected to take place in a distance range of only a few millimeters.

In two horizontal planes, at z = 10 mm and z = 20 mm, and for differing radial positions

r measurements of drop-sizes and of 2-component drop- velocities have been carried out under

constant gas and fuel oil temperature conditions and for an atomizing pressure of 8 bars. The

vertical velocity component (z-direction) is given by v^, the resultant horizontal drop-velocity

by V|, (Fig. 4 and 5). The resultant horizontal velocity v^ is given by the vector addition of the

radial and the swirl velocity components (v^ v^p). For the assumption of a symmetric spray cone

and of a linear dependency of Vj. and V(p on r (which is a weak one), v^ can been calculated at least from the measured velocity component v^,, which is perpendicular to the interference

fringe pattern.

atomizer

Fig. 4 Spray-cone parameters

horizontal Interference

Vtn' measured horizontal velocity component

(vertical to the interference fringe pattern) Vf radial velocity component

V<p« swirl velocity component

Vk • resultant horizontal velocity component "->--»•

Fig, 5 Horizontal velocity components

252
Fig. 6 shows mean values of the horizontal velocities Vj^ versus the radial distance r for

z = 10 mm and z = 20 mm. The best fit of all mean values results in the regression lines. That

means we receive a linear dependency of horizontal mean velocity, v^, on radial distance r.

radial distance/mm

Fig. 6 Axial and horizontal velocity mean values at z = 10 mm and z = 20 mm

For larger radial distances from the centre line the radial droplet- velocity increases

strongly for growing drop-sizes and is directed outwards of the spray cone. For small radial distances a narrow velocity distribution is resulting for the whole range of drop-size distribution

without a correlation between drop-size and -velocity (Fig. 7b). Fig. 7 shows a comparison of

the cross-correlation of the horizontal velocities and the diameters at a) r = 13 mm and b) r = 4

mm in the horizontal plane, z = 20 mm. It is astonishing to get nearly the same diameter distributions at both positions, because we expected smaller droplets in the centre and larger

droplets at the periphery of the spray cone. This fact again demonstrates that the interesting

changes of the spray cone parameters take place near the nozzle orifice. This means that the influences of the quality of atomization on combustion have their domain already a few millimeters below the nozzle.

540.

480.-

540. nam 480.-

420.- 420.-

nil

III)

-f 360.

s 300. 1 240.

mmmm

MT^r-n—

I

nil

~t 360. m

Mti^±J—

1

1113

nil

s 300.-

§ 240.- ^ 180.-

nil

nil

<-> ^ 180.-

III! oZ 120.- o

60

0

-60,

il

veIc III

Mil a nil nil

60- M nil ssaa

0

nil rnnins

-60,

J

nil

0 1 8 36 54 72 9 0 IC 8 12 6 14 4 IE 2 180 0 1 3 3 5 54 7 2 9 0 IC OO 6 14 4 16 2 180 W diometerW 0EL55 (Jv) S:1357 M:25 Ml diameter H OEL52 (_dv) S:1447 M:58

a) (p = 90% r = 13 mm, z = 20 mm b) (p = 90% r = 4 mm, z = 20 mm

Fig. 7a,b Diameter-velocity correlation

253
A comparison of the results of droplet velocities shows a strong decrease with growing

z-values, caused by drag of the droplets executed by the entrainment of the gas into the spray

cone. This can be seen for example from Fig. 6 for the mean horizontal velocities for the two

distances z = 10 and 20 mm. The axial mean velocity, v^, decreases with growing radial distances and fluctuates around a regression curve of nearly constant value for small distances, r. A significant sensitivity of the dropsize-distribution for temperature influences (mainly

via viscosity) can be demonstrated by the SMD of local distributions versus oil temperature as given in Fig. 8. It has been measured by the one-component PDM-configuration. In Figrs. 8

and 9 the SMD is plotted versus temperature and viscosity respectively for three positions at different horizontal planes, i. e. z = 20mm; 35 and 50 mm. The temperature influence via density and surface tension on the drop-size distribution could not be analysed explicitly, because

of the small changes of these physical properties within the temperature ranges varied (Table 1). It had also been tested whether the gas temperature affected the spray parameters; but within

the range of variation the effects could be neglected compared with those resulting from the

fuel oil temperature, because of the small distances of possible heat transfer processes in the disintegration zone before ligament disruption.

60

50H

CO 30-

20-

10-

1 1 1 11

I

1 1 • \

\

1

' i'

1

' V z=20mm,r=0mm"

V \ z=35mm,r=0mm-

* z = 50mm,r=0mm-

—'

\—'\—'—

r

20 30 40 50 50 70

Temperature/°C

80 90

Fig. 8 Dependency of the Sauter-mean-diameter on temperature (T^^ = 30 "C)

60-

50-

w z = 20mm,r=0mm

z = 35mm,r=0m.m

^ z=50mm,r==0mm

0.6

Viscosity/ mPa*s

1.0

Fig. 9 Dependency of the Sauter-mean-diameter on viscosity {T^ = 30 "C)

254
Table 1 Properties of fuel oil

temperature density viscosity surface tension

0 3

C g/cm mPa*3 dyn/cm

30 0.822 3.42 26.15

40 0.813 2.69 25.35

50 0.804 2.16 24.60

60 0.795 1.78 23.90

70 0.786 1.48 23.15

80 0.776 1.28 22.40

Table 1 also shows a strong decrease of the viscosity, /i, of the fuel oil for growing temperatures. Therefore the decrease of viscosity has been postulated to be mainly responsible for

SMD changes for growing temperature. In this context it should be mentioned that one receives

an exponential dependency of SMD on the viscosity. Similar correlations have been found by

Jasuja [10], Jones [11] and Knight [12] in the form

SMD -

but the resulting coefficients for B in our investigations are different from those in the

literature. Wang and Lefebvre [8],[9] did similar investigations using blended diesel oil with

polybutene. They also found an exponential dependency of SMD on the liquid viscosity. A better fit for the regression line is given by

SMD = B(l) . (M//lo)^(2^ + B(3), Mo = 3.83 mPa * s

[JLq is the viscosity of standard state at a pressure of 1 atm and a temperature of 25 "C.

The values of the coefficients B(l), B(2) and B(3) for different spray positions can be taken

from Table 2. Fig. 10 shows, that the SMD depends on viscosity as well as on location in the

spray cone. The local dependency can be interpreted by the different trajectories of drops of

varying sizes and by the changes of drop path's for increasing temperatures of the fuel oil.

Table 2 Coefficients of dependency of SMD on viscosity

axial

distance z

spray cone

angle B(1) B(2) B(3)

mm degree fim ^m

20

35

50

0

0

0

9.19

6.80

3.52

1.19

1.24

1.53

4.12

8.10

14.55

CONCLUSIONS

A new two-component PDM with direct storing of signals and succeeding analysis has

successfully been tested. Measurements of particle size and velocity distributions in the vicinity

of the nozzle with hollow cone spray demonstrate a linear dependency of the horizontal resultant velocity component on the radial distance r. This linear dependency with different

gradients has been found in two horizontal planes. Because of the entrainment of the surrounding gas and the drag of droplets the horizontal resultant velocity decreases for increasing axial

255
distance z. According to this fact the curvature of the particle trajectories vary with increasing

distances z tending more and more to the vertical direction.

Temperature dependencies of the spray-cone parameters have been investigated in three

horizontal planes at different radial distances. The point of interest was the correlation between

SMD and the physical properties of the fuel-oil. Because it has been found, that within the

range of temperature variation there have been only small changes of the surface tension and

the density, these properties have been neglected. Thus only the growing viscosities due to

decreasing temperatures are responsible for SMD-augmentation. The results of these temperature

investigations demonstrate an exponential increase of SMD for higher values of viscosity. Additionally a dependency of the local value of SMD on its position in the spray cone has been

found. Patternator investigations show, that this local dependency may be reduced to the temperature dependency of the curved droplet trajectories. The phenomena of changes of drop path

curvature for higher temperatures has to be investigated in future.

REFERENCES

1. Durst, F. and Zare, M., "Laser Doppler Measurements in Two- Phase Flows", Proc.

LDA-Svmp.. Copenhagen, pp. 403-429 (1975)

2. Bachalo, W. D. and Houser, M. J., "Development of the Phase/Doppler Spray Analyzer

for Liquid Drop Size and Velocity Characterizations", Proc. AIAA/SAE/ASME 20th

Joint Propulsion Conf. Cincinnati . June 1984

3. Bauckhage, K.and Flogel, H.-H., "Simultaneous Measurement of Droplet Size and

Velocity in Nozzle Sprays", Proc. of the 2nd Int. Svmp. on Appl. of Laser Anemometrv

to Fluid Mechanics. 2.-4.7.1984, Lisbon

4. Buchhave, P., Saffman, M. and Tanger, H., "Simultaneous Measurement of Size, Concentration and Velocity of Spherical Particles by Laser Doppler Method", Proc. of the

2nd Int. Svmp. on Laser Anemometrv to Fluids Mechanics. 2.-4.7.1984, Lisbon

5. Bauckhage, K. and Schulte, G., "Lasermethoden in der StromunesmeBtechnik". Editor

B.Ruck, Chap. 7: 'Phasen-Doppler-Anemometrie', AT-Fachverlag Stuttgart (1990)

6. Ruck, B., "Laser-Doppler-Anemometrie". AT-Fachverlag, Stuttgart (1987)

7. Bauckhage, K., Schone, H. and Wriedt, Th., "Using Fast-Fourier-Transform (FFT) for

the Phase-Doppler-Difference Analysis of powder and metal sprays", Proc. of the Int. Conf. on Laser Technologies in Industry. June 1988, Porto

8. Lefebvre, A.H., "Atomization And Sprays". Hemisphere Publ. Corp., New York (1989)

9. Wang, X.F. and Lefebvre A.H., "Mean Drop Sizes from Pressure-Swirl Nozzles", AIAA

J. Propul. Power Vol. 3, No. 1, pp 11-18 (1987)

10. Jasuja, A.K., "Atomization of Cruede and Resichnal Fuel Oils", ASME J. Eng. Power,

Vol. 101, No.2, pp 250-258 (1979)

11. Jones, A.R., "Design Optimization of a Large Pressure Jet Atomizer for Power Plant",

2nd ICLASS. Madison, Wis., pp 181-185, proc. (1982)

12. Knight, B.E., "Communication on the Performance of a Type of Swirl Atomizer", by

A.Radcliffe, Proc. Inst. Mech. Eng., Vol. 169, pp 104 (1955)

256
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper24

CHARACTERIZATION OF INDUSTRIAL AIR ATOMIZED LIQUID SPRAYS

FOR FLUIDIZED BED GRANULATORS

T. Isenschmid

Institute for Process and Cryogenic Engineering

Swiss Federal Institute of Technology

ETH Ziirich Zentrum, Zurich, Switzerland

ABSTRACT

Atomization is a key unit process for the design of spray dryers - in our study a continuous fluidized

bed granulator. A test rig for industrial scale nozzles has been installed to characterize two-fluid atomizers

for various spray operating conditions. Distributions of drop size, drop velocity and mass flow in different

locations of the spray cone were measured with a Phase-Doppler-Particle-Analyzer. The nozzles tested

up to now are of two-fluid type with independent control of air and liquid flow. The atomization of the liquid

jet is performed by compressed air. Air pressures varied from 2 to 7 bar, the liquid flow in a range of 1 15 to 800 l/h. Water, water based solutions, water based suspensions or a mixture of all of these have been

used. The different water based liquids were especially prepared to vary systematically the main

parameters influencing atomization. Feed-suspensions with a very narrow particle size distribution of the

suspended solids were produced by hydrocyclone classification. The viscosity has been increased by

adding polyglycol. Thus viscosities from 1 to 150 mPa s and suspensions with a solids concentration as

high as 30 % could be tested. This allowed to compare a water spray pattern to a spray pattern produced

by liquids with suspended particles and/or with higher viscosity liquids. Monitoring of the viscosity by a newly developed orifice simply measuring a pressure difference allowed an easy on-line control of

viscosity respective viscosity changes during spray operations.

Keywords

Spray characterization, two-fluid nozzles, fluidized bed granulators, pneumatic nozzles for

suspensions, spray dryer conditioned products.

1. INTRODUCTION

Disintegrating a liquid jet by various means into single drops is an increasingly and often used

technology in todays industry. In many operations the atomization process plays a significant role [1]. The

generation of a well defined spray has to be considered as a "key unit operation" especially in following areas:

- atomization of fuels

- production of powders and granules

- heat and mass transfer operations

- application of thin layers to a surface.

Requirements concerning drop size, drop velocity and mass flow distribution in the spray are determined by process conditions and requirements of products. For continuous production of powders

and granules spray dryers have long been standard in process industry [2]. However, recent

developments for dryers have led to different design solutions [3] adapted to more sophisticated

requirements.

257
1.1 Industrial application: Fluidized Bed Granulator

The Fluidized Bed Granulator (FBG), as stiown in Figure 1, is a new drying system [4] for the

production and conditioning of dyes and agricultural products in chemical process industry. The

advantages of this technology are:

- product quality requirements: Conditioned products which are dust free, free flowing and

which show fast redispersion and uniform appearance can be achieved

- safety/hygiene: Hazardous products are easier to handle since the fines are recycled

within the apparatus and a coarse product reduces risk of dust explosion

- building space requirements: The height of FBG installations can be reduced considerably

due to reduced specific drying volumes.

Granulator

Filler

Compressed Air

r-1 ri r"i r M I ! !

1 1 iliii

il

I

I

I I

i

>

I I !

I I ' L_J LJ L-J

oFan

Exhaust

Fluidized Bed

Drying-/ Fluidizing Air

Heater

Figure 1:

Simplified flow scheme of a

fluidized bed granulator

Product

Such a drying installation can be regarded in macro- and microscale. The behaviour of the

macrosystem - which is considered to be the FBG with its in- and outgoing mass flows - is governed by the

corresponding mass and energy balances. Interaction of the nozzle spray cone with the fluidized bed - one calls it the microsystem - is however a critical item of such a dryer. For example, for scale-up purposes,

for changes of feed suspensions or for required product qualities, one needs to know the distributions of

drop size, drop velocity and mass flow for the actual spray cone at different levels. Those define number

and distance of nozzles above the fluidized bed, the drying behaviour of the droplets in the freeboard and

the specific liquid load per unit area of fluidized bed surface.

2. NOZZLE TEST RIG

To investigate spray patterns of nozzles with liquid flow rates up to 1000 l/h, a nozzle test rig of

industrial size has been designed and built (Figure 2). The spray chamber is built from plexiglass in order to have visual control of the spray patterns as well as to avoid any unwanted outside disturbancies of preset

spray and flow conditions. The cross section of the spray chamber is 1 .4 m square. It has been decided to perform the spray pattern measurements in a constant co-current laminar downstream air flow, despite the

gas flow situation in the FBG is opposite. Theoretical considerations about the spray cone behaviour

under the before mentioned differing conditions showed that at least in the distances to be investigated

no significant influence has to be expected.

The liquid to be tested is kept in a closed cycle. A frequency controlled screw pump conveys the

liquids from the stirred liquid tank to the pneumatic nozzle with air supply from a pressurized air tank.

The spray produced by the two-fluid nozzles, after having passed the measurement level, is sucked

to a vane blade droplet separator at the bottom of the test rig and is separated from the gas stream.

258
see

- Water Separator

Compressor

Tank Screw Pump

Spray

Collector Droplet

1 ^ Separator J

Pump p

**** H

" 3.Am

•1 (or Online Viscosity Monitoring C*^VV. •-'..'Vi^-o";; ;t .

• 'i Jo"/.^; s^o' : S':o'i."..;^

3 .V*

5 •

Figure 2: Scheme of industrial size nozzle test rig

The collected liquid is recycled to the storage tank. All flow rates are indicated and controlled. They

can be varied following the demands for measuring the industrial relevant parameters. After careful

evaluation of existing particle sizing measurement systems an Aerometrics Phase-Doppler-ParticleAnalyzer (PDPA) has been applied. The optical system has been strictly separated from the test rig and

mounted on separate stands. In order to measure spray characteristics at all locations in the spray cone the

nozzle is attached to a three dimensional traverse system. This allows to move the nozzle outlet to nearly

any point of the spray chamber and to control its exact position.

3. MEASUREMENT SYSTEMS

3.1 Sorav pattern measurements

The PDPA from Aerometrics was used as a non-intrusive and simultaneous measurement

technique to obtain size, velocity and concentration of the dispersed drops in the spray as further

described in [5, 6]. Since the measured droplets were transparent and their refractive index close to that of

water, the receiver has been positioned at an off-axis angle of 30 degrees. At this angle the refracted light

dominates the reflected light component. The index of refraction was determined with an "ABBErefractometer".

To calibrate the PDPA a commercial generator of monodispersed drops as further described in [7, 8]

was used. This method has been established as a primary standard for instrument calibration. The liquids

tested to produce these droplets of uniform size were identical to the process liquids described in chapter

5 - except no solids could be added due to clogging of the small orifices. The measured drop diameter

d30 by the PDPA compared to the calculated drop size

d =

6 V

7t f

t1/3

(3.1)

where V was measured by a scale and f controlled with an oscilloscope (LeCroy 9600). A mean

deviation of 2 % resulted for the PDPA.

2.59
A comparative study was performed by means of a Zeiss IBAS 2000 image analyzing system (IBAS)

and a Helos-Sympatec laser diffraction meter (Helos). Tfierefore, two sieve fractions of glass beads were

measured by tfiree different sizing systems.Tfie results of these measurements by comparing the volume

mean diameters dao of the spatial distribution are as follows:

Table 3.1 : Comparison of particle diameters by different sizing techniques using glass beads

Sieve fraction 1 Sieve fraction II 160-224 iim 355 - 560 |j,m

Sizing technique d30 [m] d30 [^im]

IBAS 192 518

HELOS 198 497

PDPA 208 599

The rather large diameter measured with the PDPA for the II sieve fraction is believed to be due to asphericity and deviations from the correct index of refraction of the glass beads. Additional information

concerning such sizing technique comparisons can be found in [9, 10].

3.2 On-line viscosity control

The viscosity of the atomized fluid has a decisive influence on the spray characteristics. For that reason it is very important to measure and monitor the viscosity of the liquid on-line immediately before it is sprayed as realized in the test rig. Another question is how to measure a correct viscosity for a liquid with

suspended particles due to a certain non-newtonian behaviour.

Careful evaluation of different possible measurement principles lead to the decision to install a custom manufactured orifice in the liquid pipe leading to the nozzle (Figure 2). From the pressure

difference measured across the orifice an on-line viscosity can be derived. Liquids with a known viscosity are initially used to calibrate the orifice. The opening of the orifice has been chosen of the same size as the

liquid nozzle opening. Therefore, similar shear stresses occur at the location of the viscosity measurement

and at the nozzle. On-line viscosities with our new method are compared with a conventional viscosimeter

(Contraves Rheomat 115):

Table 3.2: Vi scosities of polyglycol water mixtures measured by the two different methods

Viscosity [Pa s] lO'^

Rheomat 23,2 29,3 39,7 48,8 87,0 121,7

Orifice on-line 19,2 32,6 36,9 51,1 86,6 107,1

4. NOZZLES TESTED

The nozzles tested are of two-fluid type with external mixture of the fluids (Figure 3). Such nozzles

are currently applied sucessfully in FBG operation following the arrangement in Figure 1. Air and liquid flow of the nozzle can be varied independently to adjust the spray cone conditions to the liquid distribution required for granulation. Compared to pressure nozzles the liquid exit opening is large - therefore clogging can be avoided even if the liquid contains suspended particles.

A swirt body in the nozzle's liquid path produces a liquid cone shaped lamella outside the nozzle.

This technique - sometimes called pre-atomization - forces the compressed air to impinge on the liquid

lamella in a close to rectangular way to guarantee optimal atomization. The air pressure has to be kept

above the critical pressure ratio. Therefore, the atomizing air reaches velocity of sound in the narrowest

cross section and expands vigorously immediately outside the nozzle. Typical air pressures vary from 3 to 6 bar. The liquid pressure determined by the requested flow rate can be kept generally very low (< 0,5

bar).

To evaluate the measurement data of the symmetric spray cone a coordinate system following

Figure 4 for positioning the PDPA probe volume is used.

260
4- SO Suspension

Air Air

Figure 3: Two-fluid nozzle as used in tests Figure 4: Coordinate system for position of probe

(Schlick 0/5) volume in spray cone

5. COMPOSITION AND SIMULATION OF TEST LIQUIDS

In a chemical production process the final step is often the separation of the product from a

suspension, the so called formulation. To transfer the solids content of such a formulation into a product,

drying in fluidized bed granulators is more and more preferred.

To characterize a formulation however, one has to analyze first the composition of suspension and

to divide it into fractions of pure liquid, disolved and solid phase (Figure 5).

For the atomization process of a formulation it is important to know the influence of its viscosity, of its surface tension, % solids concentration, the solid particulate size in the suspension and the particulate

size of its solids.

Feed

Feed Composition

40% Pure Liquid

40% Dissolved

Substance

20% Solid Matter

60% Solids Content 60% Solids Content

40% Liquid 40% Liquid

Figure 5: Typical Composition of an industrial formulation

261
Since real formulations cannot be used in tests for many reasons (price, safety, handling, ...) these

features had to be simulated. Therefore new components had to be found to synthesize suspensions

meeting the specifications determined by the industrial process as well as by the measurement technique.

For the tests reported here we replaced the real components as follows:

liquid => water

solvent => polyglycol

solids china clay or filter eel (FC).

Viscosity. To vary the viscosity polyglycol has been dissolved in water in changing quantities. For

example, adding 20 % polyglycol by weight increased the viscosity to 150 mPa s (water: ti = 1 m Pa s).

Solids. China clay and filter eel (FC), used as filter aid in filtration processes, were applied as solids in different concentrations. In a hydrocyclone classifying pilot unit suspensions with different size fractions

have been produced to investigate the influence of different particulate size of solids as well.

The model suspensions being sprayed are characterized in the following table:

Table 5.1 : Characteristics of the model suspensions being tested

Suspension Solids material solids d50 d84-dl6

type concentration [nm] [^im]

[% by weight]

B china clay 5 4 7

D filter eel 8 13 29

E filter eel 6 8 14

filter eel various 16 33

Surface Tension . The surface tension (a) of water was measured as 0,072 N/m. Mixtures of

polyglycol and water showed a decrease of surface tension which was independent of the amount of

polyglycol added. The average surface tension for all water polyglycol mixtures was found to be 0,058

N/m.

Index of Refraction . The index of refraction had to be measured for each tested liquid. A classical

"Abbe-refractometer" was used.

6. MEASUREMENT RESULTS

A water spray with the same operating parameters of the nozzle served as a comparison to each of

the other tests conducted with the formulations mentioned in chapter 5. Figure 6.1 shows a typical size, mass and velocity distribution for a water spray at three different

distances z from the nozzle. As a representative mean diameter of the droplets the sauter mean diameter

(d32) has been chosen since it represents an optimal average value for heat and mass transfer

calculations. One can see, that with increasing distance from the nozzle the larger droplets travel to the

edge of the spray. At the same time velocity and mass distribution become more even.

It is interesting to know how the sauter mean diameter of spray changes if only one of the critical

parameters is varied. A higher liquid flow rate causes larger droplets and a narrower spray cone (Figure

6.2). Liquids with viscosities of 49.6 and 82.4 mPa s are compared to water in Figure 6.3. In accordance to existing experience a higher viscosity corresponds to a distinct increase of droplet size. Simultaneously

however, the maximum drop diameter measured is significantly larger than that of a water spray.

Increasing solids concentration has a similar effect (Figure 6.4). With a slightly increased amount of

solids the sauter mean diameter becomes obviously larger, but doesn't increase as much as expected

when the solids concentration is further raised. A variation of the suspensions with similar solids concen262
Distance from Nozzle Sauter Mean Diameter [pirn] Vertical Velocity [m/s] Volume Flow [err? /s err? I

z = 300 mm

2 = 600 mm

z = 1000 mm

Figure 6.1 : Local size, velocity and mass distribution for a water spray at three different distances from

nozzle z = 0,3/0,6/1 ,0 m

(Schlick 854/0, s = - 4mm, V| = 326 l/h, pA = 3,2 bar)

tration (Figure 6.5) results in increased drop size of the suspensions compared to water. An obvious

influence of suspended solid particulate size distribution could not be found, however.

Scattered refractive light signals from drops containing higher solid concentrations show a low

signal-to-noise ratio due to multiple scattering by solid particles within the drops. At the same time, higher

light absorption by these particles reduce also the strength of the refractive light component in relation to the reflective one [11]. Therefore, further research in optical spray measurements should be directed to

liquids that favour reflection in order to reduce signal rejection and to become independent of droplet

content when measuring with the PDPA. A possibility to improve spray characterization could be the

development of appropiate model suspensions that meet optical as well as process requirements.

7. CONCLUSIONS

The presented industrial-scale nozzle test rig with PDPA measuring technique has given valuable

insights into characterization of sprays from two-fluid nozzles. The main results is that by PDPA

measurements it is possible to characterize the generated spray by distributions of droplet size, velocity

and mass flow rate as function of the distance from nozzle, which allow adaption of spray behaviour to different process requirements.

The spray characteristics for two-fluid nozzles as function of on-line controlled viscosity changes

(1...100 mPa s), of varying solid concentrations (0...0,3 kg/I) in sprayed suspensions and of varying solid

particulate size (d5o:4...16M.m) in the suspensions could be verified experimentally.

Despite the automation technique used, the time consumption for performing such measurements

is still extremely high due to calibration, adjustment and handling procedures that cannot be automated

263
264
yet. Therefore, further research efforts should especially take into consideration optical sizing, process

and tinne requirements.

NOMENCLATURE

SYMBOLS AND ABBREVIATIONS

d [m] drop diameter

f [s-"!] frequency

p [bar] pressure

r [m] radius s [m] gap between liquid tip and nozzle exit

V [m3/h] volume flow

x,y,z [m] coordinate system

INDICES

32 sautermean

30 volume mean

50 volume median

84, 1 6 points on the cummulative volume distribution curve A air

I liquid

ACKNOWLEDGEMENTS

The author is glad to acknowledge the financial support provided by SANDOZAG, BASEL, and the

Kommission zur Forderung der wissenschaftlichen Forschung (KWF) , project number 1613. Prof. L. Reh's valuable advice is gratefully appreciated.

REFERENCES

1 ] Walzel, P.: "Zerstauben von Fliissigkeiten", CIT 62 (1990) Nr. 12, S. 983-994.

2 ] Masters, K.: Spray drying handbook. 4th ed., George Godwin London (1985).

3] Gehrmann, D.: "Entwicklungstendenzen der Trocknungstechnik in der chemischen Industrie",

CIT 62 (1990) Nr. 10, S. A512-A520.

4] Uhlemann, H.: "Kontinuierliche Wirbelschicht-Spruhgranulation", CIT 62 (1990) Nr. 10, S. 822-

834.

5] Bachalo, W.D.: "Method for measuring the size and velocity of spheres by dual-beam lightscatter interferometry". Applied Optics, Vol 19, No. 3, pp. 363 (1980).

6] Bachalo, W.D. and Houser, M.J.: "Phase/Doppler spray analyzer for simultaneous

measurements of drop size and velocity distributions". Optical Engineering, Vol 23, No. 5, pp.

583-590 (1984).

7] Berglund, R.N. and Liu, B.Y.H.: Environmental Science and Technology, Vol. 7, (1973).

8 ] MDG-Manual, Aerometrics Inc., 550 Del. Rey Av., Sunnyvale, CA. 94086.

9] Dodge, L.G. and Hirlemann, E.D.: "Performance Comparison of Malvern Instruments Laser

Diffraction Drop Size Analyzers", ICLASS, London, England, (1985).

10] Young, B. W., Bachalo, W.D.: "The direct Comparison of three "In-Flight" droplet sizing

techniques for pesticide spray research"

11 ] Bauckhage, K., Floegel, H., Fritsching,U., Hiller, R.: "The Phase-Doppler-Difference-Method",

Part. Part. Syst. Charact., 5, pp. 66-71 (1988).

265
II i
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 25

CORRELATIONS FOR DIESEL SPRAY PENETRATION

INCLUDING THE EFFECTS OF THE BREAK-UP ZONE

A.J. Yule, M.R. Mirza and I. Filipovic*

Dept. of Mechanical Engineering, UMIST, Manchester, United Kingdom

Faculty of Mechanical Engineering, University of Sarajevo, Sarajevo, Yugoslavia

ABSTRACT

The semi-empirical penetration rate equation, often used to match the fully atomized

downstream spray, has been modified to include the effects of the break-up zone near the

injector nozzle. A single equation is developed, which accurately covers the whole spray flow

field, rather than the two equation, two-zone approach used in previous work. The equation

is compared with data derived using three nozzles, with variation in gas density and

cross-flow velocity. The break-up time is an adjustable constant in the new penetration

equation which is proportional to the time at which the spray penetration distance is the same

as the break-up length. Matching the equation with the measurements provides useful data on

how the break-up time, and thus break-up length, varies.

INTRODUCTION

The Equivalent Gas Jet Assumption and Spray Penetration

The 'equivalent gas jet' concept correlates experimental data on the penetration rates of

the downstream "well atomized" zones of sprays from single hole diesel injector nozzles.

Figure 1 shows the notation for the spray. This gives Upocxp"^ and, integration leads to the

dimensionally correct equation:

0.2 5 0.5 0.5

X = C 7- ' Dn t (1)

p 1 Pg ^

and Ujnj is related to by the nozzle discharge coefficient

2 2 Ulnj = Cd(2^P/pl)0-5 and Mj = 7rCD%.^P/2

Assuming a self-preserving velocity profile:

(2)

Mj = 27rps;U2r dr = 2Trpsb2U2 }f2rjdTj = l-^p^h^li^ (3)

o max o

The density is the "average" local density of the spray and, if, one ignores vaporization,

Ps = ("^L + Pg(2Tb2Un,3x^,- mLPL))/(27rb2Un,3x^, ) (A)

Far downstream the spray should spread at a rate equal to that of a turbulent gas jet, b=C2X,

and Cj should be a "universal" constant. Equation (3) gives:

U

max

Mj.

27r^2Ps

0 . 5 -1 -1 (5) C X
Fi(. 1 Definitioo of spray cbiracurisiics

70

1 60

50 X

t-su

*0

30

1

20

10

0

o o o EXPERIMENTS ( tllRZ* 1990 ) DENT (1971 ) ARAI (198« ) PISCHINGEB (1955 ). EON I 1 ) C. . S. 8

7

Pg • 22. 2 ixr

Tg . 290 tC fg . 26.7 kg /'

Vg • 0. 0 >/•

0^ • 0.265m

0.0 0.2 0.4 a 6 0.6

TIME FRDH START OF INJECTION

Fig. 2 Compiruoii between different

0 1.2 1.4 1.6

I m )

and for large x eq (4) reduces to - pg and, using eq (2),

Umax = clcjyp^'°'\^P/p^)°'\ x/2 (6)

The (rather sweeping) assumption that the penetration velocity of the front edge of the spray

is the same as the peak velocity of a steady spray at the same position x=Xp gives

Ujnax~Up~^^p^^^ and integration of (6) yields eq (1) with

Ci=(C-l Co^r°-')°-V (7)

Values for given by different authors vary considerably. Reasons were given in (1) which

recommended an average value C, = 3.0. For a steady, incompressible gas jet one may

assume -2.P=^pgU2jj^j/Cj)2 and eq (6) is ^max^^m']~^3(^^^id~^ where

2-0.5 "°- 5 °-

5

Cg = (2C2^2)/2 and = 2(2 C^Cj^) . (8)

Measurements in an air jet (2) gave C3 = 5.17, with a virtual origin at x=7Di^. A typical

value Cq = 0.65, for a diesel injector gives, =3.08, when using eq (8) with C3 = 5.17.

Sources of Variation in C,

Break-Up Process. Equation (1) is not applicable near the orifice, where the predicted

penetration rate is too high. For a zone of incomplete atomization near the nozzle, up to x = Xg, one expects a penetration velocity similar to the nozzle exit velocity and (3) assumed,

for Xp <xb.

Xp = 0.39 (2^^P/Pl)° - ^ (9)

A typical value of Cj) is 0.65, rather than 0.39 so that Ug < Uj^j and (4) explains and

confirms this observation. (3) proposed a two-equation correlation and Xp = Xg occurs where

eq (9) intersects eq (1), with =2.95. This gives:

tB = 28.65pL DN/(Pg^P)°-5 (10)

Figure 2 compares this approach to one of the authors' experiments. Eq (9) gives reasonable

fit to the initial penetration but a sharp break-point is not clear in the measurements. Fig 2

shows the curves of (5) and (6) which use = 3.01 and = 3.91 respectively.

Virtual Origin of Spray Steady gas jets have a virtual origin, at x = Xq, so that the

spreading of the jet is represented by b = C2(x - x^,). Modification of penetration equations

268
to model this should be considered and (7) shows that a virtual origin gives an apparent change m the value of C,. & & ff

Pressure Differential (3) used a rig which maintained constant after the needle

valve opened. In general ^P varies during a "real" injection and definitions for ^P include ^PO' " -J ^^•^'iiiy.iKjiij lui IIH^IUUC

based on the needle opening pressure, and ^P, averaged during the injection period- The most physically realistic definition for aP should be used.

Measurement Technique/Accuracy The method of measurement of x„ is a source of

differences between values. Penetration distances, measured along the hozzle centre-line

(by using a shadowgraph method with a photodiode) were 10% less (8) than values measured using photographs and the "largest radial distance from nozzle orifice" definition used here The front edge of a spray pulse is usually irregular and not always sharp. One must define

the edge of the spray: Here the furthest visible extent of either liquid phase or gaseous phase (vaporized) fuel is defined as the spray edge.

Spray Angle For all sprays Ps p as x «> and eq (1) assumes a universal downstream spray angle, tan-^C,. However authors have, paradoxically, provided data on

spray angle as a function of pg and other parameters. As shown by (8), for example the downstream spray angle is 18 - 20', which agrees with that expected for a turbulent gas jet The lower spray angles reported nearer to the nozzle or at low gas pressures represent a

transient stage of the spray due to (a) break up, (b) a region where p, > p^ or (c) very

large droplets (particularly when p is small). A varying spray angle should also ^give variation

of from the ideal 'gas jet' value.

EXPERIMENTS

Experiments at UMIST used three single hole diesel injectors operated with commercial

pumps and Table 1 summarises the operating conditions. Penetration measurements were made m pressurised gas chambers with, for = 0.265mm. a variable cross-flow velocity W All

measurements were made at a gas temperature 290K ± 5K. Penetration rate measurements

were made by high speed movie photography and spark photography. The "zero time" for

all penetration measurements, is defined as the time at which liquid commences to emerge from the nozzle. It is necessary to ensemble average penetration measurements, for a number

of pulses, and a minimum of 10 spray pulses were used. This scatter is due to cvclic

variations, (9).

(mm)

L/Dn Open i ng

Pressure

(bar)* (bar)* (bar)* (bar)+ (kg/m3) (m/s)

Vo 1 . pe r Pul se

(mra-^

)

I nject i on

Durat i on

(ms)

0.265 2 160 210 215 11.8

22.2

22.2

22.2

32.5

14.2

26.7

26.7

26.7

39

13.5

0

8.8

13.5

13.5

11.4 1.36

0.213 3.6 220 312 300 5 25

45

6

30

54

0

0

0

10.1 1.5

0.46 1 .65 120 254 245 5 25

50

6

30

60

0

0

0

53.7 1.9

1 Table 1. Summary of Test Conditions

*Gauge Pressures +Absolute Pressure

269
A MODIFIED PENETRATION EQUATION

Pressure Differential and Basic Constant C,

The penetration data, in Fig 2, are typical in showing that far downstream the

correlations of (5) and (3) underpredict the penetration, whilst (6) gives a small overprediction.

The correlations are evaluated assuming eq (1) with ^P = ^P. The pressure in the middle

and later parts of the injection period cannot influence the penetration rate during the first 0.5msec. There is a rapid increase in injection pressure, with a peak pressure occurring

between 0.5 and 1.0msec after needle opening time. An increasing Pinj> with time, should

principally affect the shape of the penetration curve during the break-up process, with less

effect further downstream and (4) gives a detailed consideration of this. Table 1 includes

values of the mean injection pressure and the injection pressure averaged during the first 0.5msec of injection and values of these quantities agree to within 4%. The mean pressure

differential, ^Pq.s. should be more representative of the initial spray conditions during

penetration. This agreement between averages during the initial period and during the whole

period occurs for similar injector - pump combinations and thus (averaged over the full injection period) may generally be used in eq (1). The value C, =3.8, gives best agreement

for t > 0.5msec. The three injectors used here have a similar value, Cq = 0.67 ± 0.03, and

using eq (8), = 4.71 Cq°-5. Two methods of modifying eq (1) to include a time-dependent parameter are discussed below.

Time-Varying Power Modification

Assuming a time-dependent power so that eq (1) becomes:

xp = 3.8 (^P/Pg)°-25 DjjO.5 tF(t) ^3^)

then F should reduce to the value 0.5 as t increases. This modification occurs for a time tg.

The following empirical form for F was found:

F = 0.5 + (0.032/(7^ + 0.19)) (12)

Figure 3 shows the data of Fig. 2 fitted by eq (11). A value tg = 0.25msec gives best

fit and this procedure provides characteristic break-up times from penetration curves. The

penetration curves obtained for all of the operating conditions in Table 1 gave 0.21msec < tg < 0.27msec. An investigation (4), which includes comparisons with break-up times derived by

other techniques, led to the following general expressions:

UB/Uinj = and xg = 2 . 8 x lO^e-o • ^^Dj^ (13)

The approximation Ug « Xg/tg, provides an expression for tg. Using (rL~0.03N/m,

PL=861kg/m3 and 1*1^=1 0~5m2s~^ gives (SI units):

-0.28-0.05 /i/N

tg = (3.75 X 105)Djj Pg (14)

Unlike eq (10), tg has a very weak dependency on gas density. The effects, on tg or Xg, of

the internal nozzle geometry and turbulence level are secondary, as suggested by conductivity

probe measurements (3).

270
Hyperbolic Modification of Penetration Equation

Break-up is a gradual, chaotic process so that exponential expressions, may be used.

For example:

Xp = 3.8(^P/pg)o-25 DjjO.5 to. 5 tanh[(t/tB)°-M (15)

where is found to have the same values as those used in eq (11) (i.e. tg - 0.25msec) and

can be described approximately by eq (14). Figure 3 includes eq (15) and its more rapid

tendency to reach the "classical" form of eq (1), is more realistic so that eq (15) maybe

preferred to eq (11). Figures 4 and 5 show penetration data (8) for 0.213mm and 0.46mm

diameter nozzles compared with equation (15) and the agreement is excellent.

Penetration Rate with a Cross-Flow of Gas

An empirical correction to the penetration correlation has been proposed (3) for

injection into a swirling cross-flow, with the cross flow at right angles to the injector axis:

'^p = and = (1 + (Wg/Ui^j))"^ ' (16)

o?

where XpQ is the penetration without cross-flow, given by eq (15). (3) proposed that U

should be used in eq (16), rather than Uj^j which was used for convenience. The use

Ujj^j, in eq (16), was unsatisfactory and the following definition of gave good agreement

with the present experiments:

C, = (1 + (Wg/Up))-i

(17)

Where Up is the average penetration velocity up to the time t.

U = 1 t U dt ^ ^ ^^tj dt (18)

*^ t ^ t dt t O O

Correlation of cross-flow effects by the ratio of the cross-flow velocity and the average

penetration velocity appears physically reasonable. Equation (17) and (18) give:

271
(19)

X

— CORRELATION

«0 EXPERIMENT. Jham (1986) c 0 Pg .5 b*- /

50 • P9 •50 bar

40 —

»

JO "9 « 0. 0 >/

20 290 K \ . 0.44 HI

IC • 254 trar

0 .

0.0 0.2 0.4 0.6 TIME FROM START OF INJECTION (ms)

Fig. 5 Comparison betwefin measured penetration

rate and new correlation; nozzle

diameter 0.46min

40

s

€

30 X

IliJ 20 aatn 10

CORRELATION

EXPERIMENTS,

o Pg s11. 6 bar

• Pg .22.2 bar •32.5 bar

Tg • 290 K

Wg - 13.5 «/«

D. = 0. 245 mm

0.0 0.2 0.4 0.6 TIME FROM START OF INJECTION

Fig. 6 Measured peoetration rate ui cross- flow and new correlation; nozzle diameter D.265mm

0.8

(ms)

This modification for the "compression" effect of the cross-flow should be assumed valid only

for the range of cross-flow velocities and penetration distances which have been investigated.

Figure 6 shows measurements of spray penetration for Wg = 13.5m/s with variation of gas

pressure and Fig. 7 shows measurements, for variation or the cross-flow at one pressure.

There is excellent agreement with eq (19). When the spray is not injected normally to the

cross-flow, or when the spray has deviated significantly a component of cross-flow velocity in

the same direction as the direction of penetration becomes important. This produces a "stretching" effect which opposes the compression effect of the orthogonal flow. Experiments

(10) with angled injection into a cross-flow show that further modification of the penetration

correlation is required under these conditions. It has been found that these correlations are in

good agreement with CFD predictions of the flows (7) when a submodel of the break-up

process is included.

CORRELATION o

EXBERI11ENTS. 0 . 40

o Wg > 0. b/s Oqx^

« Wg a 6. 6 «/•

30 - 1 5. 5 >/•

20

Pg • 22. 2 bar

Tg . 290 K

10 fg . 26. 7 kg /'

0. . 0. 265

0

0 0 0.2 0. 4 0. 6 0.8 TIME FROM START OF INJECTION

Fig. 7 Measured penetration rate at constant KU pressure, with variation of cross- How velocity, compared with new

correUtion; nozzle diameter 0.265mm

272
CONCLUSIONS

A modified form of the penetration rate equation, for sprays produced by simple orifice diesel

injectors, has been derived by considering the effects of the break-up zone near the nozzle.

The equation gives excellent agreement with data for three nozzles, with variation in gas

density and cross-flow velocity. Use of the correlation provides information on break-up time,

which shows little variation for the range of conditions which have been examined, and

break-up length, which has a much stronger dependency on both gas density and nozzle

diameter.

NOMENCLATURE

b

C C1 » 2

c.

'D

Dn

F(t)

mg and mr

Mf

'^inj

aP

^o _ and -i^P

rReg

0. 5

t, t U

u

u

B

inj

B

max

Up

We

Wg

Pg,

V

Ml

PL' Ps

\-0. 5

)

half-width of spray or jet

empirical constants

cross-flow parameter, eq 17

nozzle discharge coefficient (Cq = Uijij(2^P/pL)"

equivalent nozzle diameter, Dg = D^{piJpg)°-^

nozzle orifice diameter

variable power in modified penetration equation (11)

gas and liquid phase mass flow rates through spray cross-section

momentum flow rate through nozzle orifice (instantaneous)

gas pressure (absolute)

injection pressure

pressure difference between nozzle inlet and outlet (^P = Pjj^j - Pg)

value at needle valve opening time

average values for whole injection period and first 0.5msec

gas and liquid phase volume flow rates at spray cross-section

radial distance from centre-line

Reynolds number, Reg = PiJJqD^/

time after commencement of injection and break-up time

velocity in nozzle orifice

penetration velocity in break-up zone (« Xg/tg)

centre-line velocity

penetration velocity = dxp/dt

average penetration velocity at time t, = (Xp/t)

Weber number, We = pgUg^Dj^/tTL

gas cross-flow velocity

distance along axis of nozzle

maximum penetration distance, measured radially from nozzle orifice

densities of gas, liquid fuel and spray

dimensionless radial coordinate r/b

surface tension of liquid

viscosity of liquid

REFERENCES

Williams, T.J., "Parameters for correlation of penetration results for diesel fuel

sprays", Proc. Inst. Mech. Engineers, Vol. 187 69.73, pp771-774, London (1973).

2, Wygnanski, I. and Fiedler, H., "Some measurements in the self-preserving jet", J. Fluid Mech., 38, 3, pp557-612 (1969).

273
3. Arai, M., Tabata, M., Hiroyasu, H. and Shimizu, M., "Disintegrating process and

spray characterization of fuel jet injected by a diesel nozzle", SAE Technical Paper

840275 (1984).

4. Yule, A.J., Mirza, M.R. and Filipovic, I., "On the break-up times and lengths of

diesel sprays", Submitted to Sprays and Aerosols' 91, University of Surrey, Sept.

(1991).

5. Dent, J.C., "A basis for the comparison of various experimental methods for studying

penetration", SAE Trans., Vol. 80, Paper No. 710571 (1971).

6. Pischinger, F., "Verfahren zur untersuchung von diesel-einspritzstrahlen", MaschBau

Warmew, 10 (1955).

7. Yule, A.J., Mirza, M.R. and Filipovic, I., "Measurements and correlations of diesel

sprays", submitted to Atomization and Sprays, Jan. (1991).

8. Tham, S.Y., "The study of diesel spray structure and penetration in high pressure and

temperature quiescent gas", PhD Thesis, University of Manchester (UMIST) (1986).

9. Aval, S.M., "A study of pulsed sprays in a high pressure and temperature gas

cross-flow", PhD Thesis, University of Manchester (UMIST) (1990).

10. Mirza, M.R., "Study of diesel sprays interacting with cross-flows and solid boundaries",

PhD Thesis, Submitted to University of Manchester (UMIST) (1991).

274
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 26

BREAK-UP LENGTH OF A LIQUID JET AND

INTERNAL FLOW IN A NOZZLE

H. Hiroyasu*, M. Arai^ and M. Shimizu^

* Dept. of Mechanical Engineering, University of Hiroshima, Hiroshima, Japan

Dept. of Mechanical Systems Engineering, Gunma University, Kiryu, Gunma, Japan

* Dept. of Mechanical Engineering, Kinki University, Osaka, Japan

ABSTRACT

A new break-up phenomenon in a liquid jet, in which was friction inside the nozzle hardly contributed to the

formation of the surface disturbance of the issuing jet, was discovered in a jet of a higher speed than a spray jet

Owing to lower turbulence, the break-up length was very long and the disintegration of the jet was driven

only by an interfacial force between the jet and surroundings. A cavitation fixed at an entrance of the nozzle,

generated strong internal turbulence, and the surface disturbance of the jet was a result of this. However, if the cavitation was too strong to prevent disruption of the cavity, a constricted jet, that was separated from the

wall, issued from the nozzle, similar to super-cavitation flow around a wing. The break-up lengths were

shortened by the cavitation but elongated by the separated flow. They are discussed using photographs of

internal flows in a transparent nozzle.

INTRODUCTION

It has been known recently that the break-up phenomena of liquid jets are controlled by both internal

turbulence in a nozzle and interfacial force between the jet and surroundings. The main cause of

disintegration from a smooth jet to a droplet state is thought to be the surface instability expressed by the Rayleigh's

theory In a wavy flow region, the instability of the jet increases with the turbulence in it If the velocity

of the jet further increases, interfacial forces, such as shear forces and pressure perturbations around the jet,

become more important factors in the turbulence motion of the surface. Also the turbulence in the internal

flow increases with an increase in the velocity. Increased velocity will finally result in a cavitation fixed at the

entrance of the nozzle. This generates strong turbulence in the internal flow of the jet. This strong

turbulence is amplified by the interfacial forces mentioned above and the jet disintegrates quickly as in the spray

region

In the spray flow region, since the intensity of the cavitation is strongly dependent on the nozzle

structure, the break-up length shows different behaviors with different nozzle geometries. This Hnk has already

been suggested in previous researches However there has been little attention paid to the internal flow

behavior after cavitation. Strong cavitation that stretched from the entrance to the exit of the nozzle could

exist in a set condition. The constricted jet, which started from the entrance and continued to the exit, could

sometimes be observed with a further increase of the velocity In such a condition, no strong turbulence was

generated and a high speed smooth jet was obtained. In spite of the strong interfacial force caused by the

high speed jet, the break-up length was elongated, since there was no surface disturbance that could be amplified.

In this paper, the main concentration of the study was on the phcnomenological analysis of the break-up

mechanism of this smooth jet.

275
BREAK-UP LENGTH OF A LIQUID JET

Constricted and Non-constricted Jets

The velocity of a jet from the nozzle is usually lower than the theoretical velocity derived by the Bernoulli's

equation. This is due to pressure drops caused by the wall friction, and by the turbulent wake or cavitation at the

entrance, which are not negligible. The turbulence caused by these pressure drops results in the surface

disturbance and this leads to the disintegration of the jet. The length-to-diameter L/D ratio of the nozzle and the

Reynolds number Re of the flow, control the intensity of the wall friction. The wake and cavitation are affected

by the geometry of the nozzle entrance. Based on this consideration, two kinds of injection nozzles, shown in Fig. 1, were selected to study the general

forms of the break-up length. Water was continuously injected through these nozzles. The injection velocity of

the jet was calculated as the sectional average velocity in the nozzle. The Reynolds number Re of the jet was

calculated with kinematic viscosity of water, nozzle diameter and the above velocity. An electrical circuit was

used to determine the break-up length ^\

The break-up lengths of the jets injected through the nozzles-A and -B are shown in Fig. 2 as functions of the injection velocity. With the nozzle- A, the spray jet was observed in the higher velocity

region than 10 m/s, and the break-up length decreased with an increase of the velocity from this value.

However at a certain value of this velocity, the break-up length discontinuously elongated and a smooth jet was

observed. There was no spray jet formed after this transition. When the injection velocity was reduced after the

transition, the break-up length decreased but was still longer than the break-up length existing before the transition.

The break-up length had the hysteresis in both the wavy and spray regions.

The break-up length of the jet produced by nozzle-B changed smoothly with an increase in the velocity. The

length was shorter than that produced by nozzle-A after the transition, in spite of the jet still resembling the

pre-transition form. However it was much longer than the jet before the transition.

5000

Fig. 1 Nozzle specifications

0-1 10 10 10^

Injection Velocity V; m/s

Fig. 2 Break-up lengths of constrict and non-constricted jets

276
The photographs of the jets formed by nozzle-A are displayed in Fig. 3. The numbers that indicate the jets

correspond to the numbers in Fig. 2, numbers 9 and 10 being the spray jets. The photograph indicated by

11 was taken just after the completion of transition. The surface of the jet was very smooth and initiation

of surface disturbance, that can be seen in the photograph, was related to downstream region of the jet. The

break-up mechanism that was observed in the downstream regions of the jets, indicated by the numbers 1 1 to

16, was almost same as the mechanism observed in the jets 7 to 9. It is concluded that there was no surface

disturbance that could be amplified by the interfacial force produced around the high speed jet The break-up length

after transition was elongated by the related development of the surface disturbance.

The internal flow in nozzle-A observed by the transparent nozzle is displayed in Fig. 4. When the velocity

increased from lOm/s, a fixed cavitation appeared at the entrance of the nozzle. The turbulence caused by the

destruction of the cavity produced the surface disturbance of the jet in co-operation with the wall friction in the

nozzle. This disturbance was amplified by the interfacial force and resulted in the short break-up length of the

jet. The cavity was stretched to the exit with an increase of the velocity in the spray jet region. However, when

the velocity reached the transition value, the reattachment point of the separated flow caused by the cavitation

moved away from the exit. The flow that was separated from the wall was generated at the entrance.

After this transition, no wall firiction affected the flow and a very smooth jet appeared as mentioned before.

1 2 3 4 5 6 7 8 9 10 n 12 13 \U 15 16

V, m/s Q3 OA 0.6 0.9 11 21 95 105 12 13 14 16 18 20 22 25

Nozzle -A P = 3 0mm L/D = i Fb = 01MPa

Fig. 3 Photographs of the jets injected through the nozzle-A

PiMRa 0-10 0-115 0-120 0-125 0130 0-I^A 0-200

Vlm/S 10-5 113 11-5 11-7 12-0 12-6 13 9

Nozzle-A L/D=A D=3 0mnn Pa=01MPa

Fig. 4 Internal flows in the nozzle-A

277
The photographs displayed in Fig. 5 are the jets injected by the nozzle-B. It had a quadrant shape at its

entrance. No cavitation and no constricted flow were produced but the wall friction was always acting on the

jet. As a result, the surface of the jet, close to the exit, had slight disturbance that could be amplified by the

interfacial force and it caused a longer break-up length than the constricted jet injected by the nozzle-A.

123i567 8 9

Vitnte 03 Oi 06 09 11 21 95 105 12

0

20-

40-

60^ mm

I !

20

15 16 22 25

t

n ^

1

i I

No-^zle-B D=3.0nnm L/D = A Fb^O.IMft

Fig. 5 Photographs of the jets injected through the nozzIc-B

General Forms of the Break-up Length

General forms of the break-up length are schematically shown in Fig. 6 as a function of the injection velocity.

By applying Tanasawa's classification of the disintegrating states for a jet, the dripping region was determined

and is indicated by symbol A, the smooth jet B, the transition jet C, the wavy jet D and spray jets which are shown by E, G and F. The numbers 1 and 2 are the points where the break-up length elongated

discontinuously as the injection velocity increased. However 3 and 4 are the points where the break-up length

discontinuously shortened with a velocity decrease. The hysteresis of the break-up length is shown by the

broken lines.

When the surrounding pressure or density was too low to generate the possible interfacial force around the

jet, no spray jet was formed and break-up length increased with the velocity as shown by symbols SH and H.

When the pressure was not below this limiting value, the break-up process, indicated by SF, appeared on the

constricted smooth jet with high velocity.

9i_

3^ r / 1 SF

b/ V

G

F_

a/

Injection Velocity Vi

Fig. 6 General behavior of the break-up length

278
The photographs of the high speed jet indicated by SH, H, SF and F are shown in Fig. 7. The jet shown by

SH exited from the nozzle as constricted flow and no disturbance was imparted to the jet. However, the jet shown

by H was non-constricted flow and had turbulence as a result of the wall friction. The difference in surface

perturbation between SH and H was caused by turbulence, due to the geometrical disparity of the nozzles.

When the ambient pressure was so high that the interfacial force between the jet and surroundings had sufficient

intensity to modify the break-up phenomena, the break-up behavior of SH and H changed to SF and F, respectively.

The spray jet indicated by F was usually observed to resemble a diesel spray injected into pressurized

surroundings.

INTERNAL FLOW AND CAVITATION

Super-cavitation

The schematic appearance of the jets are illustrated in Fig. 8. The jet indicated by (a) exits from an orifice

and has a very smooth surface, since no wall friction is acting on it. The jet (b) is produced by a nozzle that has

a smooth entrance, or is long enough to damp out turbulence generated at the entrance. The surface of the

issuing jet has a slight disturbance that is generated by the wall friction. The jets indicated by (c), (d) and (e)

are formed by the same nozzle but different injection velocities. The jet (c) is observed to be in a relatively low

velocity condition. The entrance of sharp edge type forms a wake in the internal flow of the nozzle. It generates

the flow pattern that results in a turbulent jet, when acting in co-operation with the wall friction. When the

velocity increases beyond a certain value, a cavity fixed at the entrance appears as shown in (d). A disruption of

the cavity generates the intense turbulence which can perturb the jet surface close to the exit This situation

is usually observed in a spray jet and it is the reason why the break-up length of the spray jet decreases with an

increase of the velocity. However if the cavitation is too strong to prevent a separated flow expanding and

reattaching to the wall, a constricted jet, that is separated from the wall, issues from the nozzle. This is similar

to super-cavitation flow around a wing or in a long orifice The terminology of this situation, supercavitation inside a nozzle, is proposed in this paper. Wall friction and turbulence generated by cavitation can not work on the constricted jet. The break-up length in this situation is elongated since there is no surface

disturbance that can be amplified by the strong interfacial force around the high speed jet

Fig. 7 Effects of surrounding pressure "^'8- * Schematic appearances of the jets on the disintegration of high speed jet

279
Cavitation Number

A cavitation number is usually used as a criterion of an onset of a cavitation Figure 9 shows the

schematics of the flow and the pressure distribution in the nozzle.

The static pressure P2 is defined using the Blasius's equation of the wall friction as

P2 = Pa + If ^-L^) 2D Vj^

(1)

f- 0.316

Re ^-^^ ( Blasius's equation ) (2)

The velocity Vj is derived by

Vi = (-^)' Vi

Dc (3)

and the static pressure Pj is expressed by

Pi = P2 + l-pV.M l-( D_)M

2 I Dc 1

Then, the cavitation number k„ is defined as

(4)

iq = -PliPy_

2

-( Dc )4 / p, -p, ^ f(L-Lc) ^ ^ I

J

I 2 I (5)

where, P is the vapor pressure of the liquid.

280
Table 1 is a list of the calculated cavitation numbers using the equation (5) and reasonable assumptions for

Lj, and P^. The intensity of the cavitation increases with a decrease in the cavitation number. The value smaller

than 1.0 usually indicates the existence of the fixed cavitation. The negative value means too strong cavitation

exists. The appearance of the jet could be explained as follows. The cavitation number in the flow with no wake

had a relatively large value. The transition from cavitation to super-cavitation was indicated by a medium value

of the cavitation number. The negative values except Fig. 7-F correspond to the super-cavitation situations.

The spray jet indicated by Fig. 7-F also had a low enough cavitation number to form a constricted jet.

However, it was a special situation where the interfacial force caused by surroundings of high density was

strong enough to dominate over the disintegration process of the jet. Even if the constricted flow started at the

entrance of the nozzle, the surface disturbance developed before it reached the exit of the nozzle and a turbulent jet

resulted. Further study on the cavitation number calculated by correct values of and other variables is needed to explain the transition from cavitation to super-cavitation.

Table 1 Cavitation numbers calculated by the equation (5)

Nozzle D

(mm)

Pa

(MP*)

Condition

-X- Kc Comments

A 3.0 0. 1 Fig.5-No.10 0.632

Transition from cavitation

to super-cavitation

B 3.0 0. 1 Fig.7-No.10 1 1. 6 No wake

B 0.3 0.003 Fig.3-SH 3.28x10"^ No wake

A 03 0.003 Fig.3-H -0.870 Super-cavitation with

turbulence

A 3.0 0.07 Fig.3-SF -0.826 Super-cavitatation

A 03 3. 0 Fig.3-F -0.851 Spray jet

,

High turbulence

^- -0^=0.6 , f^rl-e.= o and R/ = 2.35KR3 were assumed

CONCLUSIONS

From the phenomenological analysis of the liquid jets, the relationship between the break-up length and the

internal flow in the nozzle was explained as follows.

(1) The jet coinciding with the fixed cavitation at the entrance of the nozzle appeared as a spray jet The

break-up length in this situation decreased with an increase in the velocity since turbulence generated by the

cavity increased with the velocity increase.

(2) However, if the cavitation was too strong to be disrupted in the nozzle, a constricted jet issued

from the nozzle. The appearance of this flow field was similar to the super-cavitation around a wing. The

break-up length was elongated since there was no surface disturbance that could be developed.

(3) The cavitation number in the internal flow of the nozzle was introduced to explain the intensity of the

cavitation.

281
REFERENCES

1) Rayleigh, Theory of Sound, 2nd edn., Dover reprint, New York, (1945).

2) Levich, V. G., Phyicochemical Hydrodynamics, Prenetics-Hall, Inc., Englewood Cliffs, N. J. (1962).

3) Tanasawa, Y., High speed diesel engines I, "Fuel systems and spray combustion", Sankaido ( in Japanese ), (1956), 87.

4) Rupe, S. H., "On the dynamic characteristics of fuel liquid jet and a partial correlation with orifice geometry",

JPL Tech. Kept., 32 (1962), 207.

5) Lee, D. W., " The effect of nozzle design and operating conditions on the atomization and distribution of fuel

sprays", NACA Tech. Rept., No.425 (1932), 565.

6) Ruiz, F. and Chigier, N. " Mechanism of high speed atomization", Proc. 3rd ICLASS, Vol. 1 (1985), VIB/3/1.

7) Reitz, R. D. and Bracco, F. V. "Mechanism of atomization of a liquid jet", Phys Ruids Vol.25 No.lO (1980),

1730.

8) Hiroyasu, H., Shimizu, M. and Arai, M.," The break-up of high speed jet in a high pressure gaseous

atmosphere" Proc. 2nd ICLASS (1982), 69.

9) Arai, M., Tabata, M.,Hiroyasu, H. and Shimizu, M., " Disintegrating process and spray characterization of fuel

jet injected by a diesel nozzle", SAE Paper No. 840275 (1984).

10) Arai, M. Shimizu, M. and Hiroyasu, H. " Break-up length and spray angle of a high speed jet" Proc. 3rd

ICLASS, Vol.1 (1985), IB/4/1.

11) Arai, M., Shimizu, M. and Hiroyasu, H. "Break-up length and spray formation mechanisms of a high speed

liquid jet", Proc. 4th ICLASS, (1988), 177.

12) Yamazaki, T., "Cavitation Engineering", Nikkan-Kogyo Press. (1978).

13) Yokota, M., Ito, Y. and Shinohe, T, "High Speed Photographic Observations of Cavitation Arising in The

High-Speed Oil-Flow Through a Very Small Long Orifice", 9th Int. Symp. on Jet Cutting Tech., Sendai Japan

(1988), 13.

14) Daugherty, R. L. and Franzini, J, B., "Fluid Mechanics", McGraw-Hill,Inc., New York, (1965).

282
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 27

AN EXPERIMENTAL STUDY ON INITIAL BEHAVIOR OF

DIESEL FUEL SPRAY CHARACTERISTICS

J.-C. Huh*, G.-Y. Lee^ and O.-Y, Yang*

*Dept. of Marine Engineering, Cheju National University, Cheju, Korea

^Dept. of Mechanical Engineering, Inha Technical Junior College, Inchon, Korea

*Dept. of Mechanical Engineering, Inha University, Inchon, Korea

ABSTRACT

An experimental study has been conducted to investigate the characteristics of

diesel fuel spray injected from a single hole nozzle into a cylindrical steel chamber

with transparent windows. With modified fuel injection system, the spray tip

penetration, the spray angle and the initial behaviour of diesel fuel spray have

been observed by shadowgraph. The effects of preheated fuel at temperatures of room

temperature ~ 150 °C and ambient gas pressures of 0.1, 1.5 and 3 MPa on the

characteristics of diesel fuel spray were also investigated.

The spray angles in the case of preheated fuel supply were larger than those of

non-preheated fuel conditions. Particularly, spray angle in the initial stage of fuel

injection was sharply increased due to the result of change of fuel properties, but

the spray tip penetration was decreased.

With the increase of ambient gas pressure, the spray tip penetration was decreased

and the spray angle was increased. The transition point in the spray tip penetration

was observed at the higher ambient gas pressures considered and it strongly depends upon

the fuel temperature.

The initial stage of injection of diesel fuel spray showed the form of nondisintegrated intact core, but the formation of ligament increasingly was developed

with the increase of injection time.

INTRODUCTION

Combustion of liquid fuels in diesel engines is highly dependent on effective

atoraization to increase the specific surface area of the fuel and thereby achieve

high rates of mixing and evaporating. It is commonly known that the disintegrating

process of fuel spray mainly depends on the properties of injected oil, fuel

temperature, injection pressure, ambient gas pressure and injection nozzle

geometry (1-4). The combustion process is inherently affected by the spatial and temporal

distributions of atomized liquid fuel and/or atomized fuel within the combustion

chamber, which, in turn, are strongly dependent on the details of the initial atomization

process itself.

Reitz (5) reported that the jet appears initially to emerge shaped like a blunt

cylindrical rod, with an exit diameter approximately equal to the nozzle exit hole

diameter. Newman, et al.(6,7) showed the existence of a continuous portion of liquid in

the vicinity of the nozzle in a high speed jet. In addition, Pai and Nijaguna(8)

supported the evidence that the liquid, initially, just after leaving the nozzle is in

283
the form of a continuous body. However, initial behaviour of fuel spray at a location

just behind the nozzle exit has not been well understood.

The purpose of this experimental work is to obtain the experimental data on the

spray tip penetration and spray angle of the diesel fuel spray. The influence of

preheated fuel and ambient gas pressure on the behaviour of diesel spray is studied.

Time - variation process of diesel fuel spray within the injection time of 1 msec

from the nozzle exit has been principally investigated by the shadowgraph by means of a incident light.

EXPERIMENTAL APPARATUS AND PROCEDURE.

The experimental apparatus consists of a fuel injection system, electric heaters,

an electronic cntrol unit, a pressure chamber, a light source and a camera as shown in

Fig.l. One of car plunger pumps mounted on the injection pump test device was used as a pressurizer. The diesel fuel from the plunger (4) flows into an accumulator (6)

which is the inside volume of 5.1 liter and is then pressurized to the required high

pressure. The pressurized fuel is heated up by the first electric heater (7) supplied

to a supplementary accumulator (26). The second heater inserted in the supplementary

accumulator heats up the preheated fuel again and the accumulator reduces the pressure

pulsations during the whole injection period. The nozzle is controlled by electronic

control unit (19,20). The fuel injection is synchronized with the input signal in the

solenoid. The setting time can be varied from l//sec to Isec.

Experiments were made for injecting a single charge of diesel fuel into the

chamber through a nozzle, and for photographing the developing spray by the stroboscope

or nanolite through windows in the cylindrical chamber wall. The pressure chamber was designed for gas pressure up to 5MPa and the highly pressurized N2 gas was supplied to

the chamber. The inner diameter and the length of chamber were 196mm and 500mm

respectively. The chamber was made of steel with transparent rectangular windows of

230 X 130 mm on to opposite faces.

The shadowgraphic method was introduced in this work. Photo- interrupter (11) was

installed at the tip of the spindle connected to the needle. The spindle interrupts the

inspection light beam of the photo-interrupter. At the moment of operating solenoid,

the injection start signal from photo-interrupter trigers the digitalized regarding

circuit (21) during the desired time from 1 //sec to 10 msec. A stroboscope or a nanolite (22) is, then, activated by an output of the retarder and emits the flash.

The flash duration of the stroboscope and the nanolite is about 3 //sec and 17 nsec,

respectively. The camera was at the state of B-shutter and photographed the spray at

the lightening moment.

The experimetal conditions are shown in Table 1.

Table 1. Experimental conditions

Titles Descriptions

Injection Pressure (MPa)

Ambient gas pressure (MPa)

Ambient gas temperature

Preheated fuel temp. ( <>C ) Liquid

Injection Nozzle

Lenses

7, 14, 21

0.1 1.5 3.0 (N2 gas ) Room temperature

Room temp., 50,100,150

Diesel fuel (specific weight: 0.84)

Aspect ratio(L/D):2 D=0. 65mm, Single hole

105mm microlens (f=4)

Tele-micro lens (magnifying ratio: 4. 5)

Film KODAK TRI- X (ASA - 400)

284
RESULTS AND CONSIDERATION

Spray tip penetration and spray angle

Many photographs of sprays by shadowgraph are taken to measure the spray tip

penetrations and spray angles. Photo. 1 shows representative photographs of the

developing process of spray for the preheated fuel. The spray tip penetration and the

spray angle were measured on the basis of those figures.

Fig. 2 shows the variation of spray tip penetration and spray angle for various

fuel temperatures and ambient pressures at the injection pressure of 7 MPa as a function

of time.

It can be seen from Fig. 2(a) that the logarithmic spray tip penetration for all

four fuel temperatures linearly increases up to before transition time with the slope

of 1 and then increases with that of 0.4-0.6. It is considered that the liquid column is

disintegrated before the transition point and the spray is then fully developed after

it. The lower ambient pressure and the higher fuel temperature allow the shorter length

and the more delayed occurrence of transition point for spray tip penetration with

respect to time. This reveals that the fuel temperature is one of the important

parameters governing the disintegrating process of spray.

From Fig. 2(b), the effect of preheated fuel on the spray angle is clearly found at

the initial period of injection with Pa =0.1 and 3 MPa respectively. From 1.5 and 2.0

msec after injection repectively for those ambient pressures, the nearly constant spray

angle with the variation of fuel temperature can be found. On the other hand, the spray

angle is increased with the increase of fuel temperature through the whole duration in

the case of ambient temperature of 1.5 MPa. It is particularly noticed that the spray

angle after transition point at 1.5 MPa is coincident with that at 3 MPa for the

preheated temperature of 150 °C. This effect is attributed to the variation of surface

tension and viscosity of fuel as the fuel temperature increases.

The effect of ambient pressure and fuel temperature on the spray tip penetration

and spray angle at the injection pressure of 14 MPa is shown in Fig. 3.

It is found from Fig. 3(a) that for all three ambient pressures the logarithmic

expression of spray tip penetration and time shows the linear relationship with two

distinct slopes. This result shows good agreement with that of Arai,et al.(9). For the

given conditions of fuel temperature the variation of spray tip penetration shows a similar trend to that obtained for the various ambient pressures. The spray tip

penetration decreases with the increase of fuel temperature at Pa = 0.1 and 1.5 MPa

respectively. It is evident from the figure that the fuel temperature of 150'C

remarkably reduces the spray tip penetration for the ambient pressures considered, while

the other fuel temperatures at the given ambient pressures has no effect on the spray

tip penetration. As same as the observations from Fig. 2, the higher fuel temperature

and the lower ambient pressure show the delayed transition.

The influence of fuel temperature on the spray angle at the ambiemt pressure of 0.1

MPa is not significant. The spray angle with fuel temperature of 150''C at Pa= 1.5 MPa is

considerably increased comparing with the other fuel temperature considered. The spray

angle in this case is higher than that obtained at Pa= 3 Mpa. The effect of fuel

temperature on the spray angle is clearly found at preheated temperature of 100 and 150

°C repectively from the Fig. 2(b).

The effect of the fuel temperature and the ambient pressure on the spray tip

penetration and spray angle at Pi = 21 MPa with respect to time is shown in Fig. 4.

The spray tip penetration and the spray angle at the ambient pressure of O.lMPa is

constant for all preheated temperature of fuel considered. The transition point for

fuel temperature at the ambient pressure of 3 MPa can be found at 1.2 msec for room

temperature, at 1.8 msec for 50°C, and at 2.5 msec for lOO'C and 150''C respectively.

This appears to be due to the ratardation of transition time by the higher fuel

temperature, leading to the increase of penetration.

The spray angle in this ambient pressure is, however, nearly constant regardless of

285
preheated temperature of fuel. It is evident from Fig. 4(b) that the spray angle is

increased with increasing the ambient pressure.

Initial behaviour of spray

Photo. 2 shows the enlarged photographs of spray at two different injection

pressures. In the case of injection pressure of 7 MPa shown in Photo. 2(a), a continuous

sinusoidal wave is formed on the surface of the liquid column at injection time of 0.05

msec. After the injection time elapsed over 0.15 msec, the unstable growth of suface

waves with short wavelength on the liquid column which is caused by the relative motion

of the jet and the ambient air grows larger, and this leads to the formation of

ligaments. With the continuous interaction with the ambient air, the ligaments yield

droplets at the injection time of 0.3 msec. The process of fully developing spray is

clear from the photographs of injection time of 1.0 and 2.0 msec respectively.

In the case of injection pressure of 21 MPa shown in Photo. 2(b), liquid column of

1 mm can be observed at a location just behind the nozzle exit at the injection time of

0. 01 msec and the tip of the liquid column has the shape of rivet, which is considered

to be caused by the interaction between the surface tension of diesel fuel and the

aerodynamic resistence of ambient air. At injection time of 0.05 msec, the formation of

ligament reveals in the circumferential surface of the unbroken portion. As injection

time increases to 0.15 msec, the surface of the unbroken portion has been largely

deformed and the vigorous formation of ligament is evidently detected. At injection time

of 0.4 msec, ligaments are observed to coexist with many droplets disintegrated from the

ligaments. Fully developed spray is shown after injection time of 1 msec and the liquid

column is observed as the form of cone. The enlarged photographs of spray at two different fuel temperatures are shown in

Photo. 3. Comparing with the effect of fuel temperature on the spray characteristics, it

is found that the spray is considerably expanded in the case of fuel temperature of 150

°C. This appears to be due to the decrease of surface tension and viscosity of fuel

resulting from the increase of fuel temperature.

CONCLUSIONS

Important conclusions to emerge from this experimental study on the initial

behaviour of diesel fuel spray characteristics are:

1. The spray angles in the case of preheated fuel supply within given conditions

were larger than those of non-preheated conditions. In addition, the spray angle in the

initial period of fuel injection was rapidly increased owing to the result of variation

of fuel properties, but the spray tip penetration was decreased. With the higher ambient

gas pressure, the transition point in the spray tip penetration strongly depends on the

preheated fuel temperature.

2. The spray penetration was decreased and the spray angle was increased with

incresing the ambient gas pressure considered. The transition point in the spray tip

penetration was not found at the ambient pressure of 0.1 MPa regardless of fuel

temperature.

3. The initial period of injection of diesel fuel spray reveals the shape of

non-disintegrated intact core. However, the formation of ligament was increasingly

developed with respect to injection time.

REFERENCES

1. Schmeitzer , P.H., " Mechanism of Disintegration of Liquid Jets", Journal of

Appplied Physics, Vol.8, pp. 513-521 ( 1937 ) 2. Walsh, G.J. and Cheng, W. K.

, "Effects of Highly-heated Fuel on Diesel Combustion ". SAE Paper , No. 850088, ( 1985 ) 3. Dent , J.C., " A Basis for the Comparison of Various Experimental Methods

of Studying Spray Penetration.", SAE Paper No. 710571 ( 1970 )

286
©0©

l0 0 I

1. Fuel Tank

2. Fuel Filter

3. Motor

4. Fuel Injection Pump

5. Pressure Gauge

6. Accumulator

7,9. Electric Heater

8,18. Thermo-couple

10,15. Pressure Transducer

11. Photo-interrupter

12. Solenoid

13. Nozzle Holder

14. Nozzle

16,17. Temp. Controller

19,20. Nozzle Driving

& Controller

21. Retarder

22. Stroboscope & Nanolite

23. Oscilloscope

24. Pressure Chamber

25. Camera

26. Supplementary

Accumulator

Fig.l Schematic diagram of experimental apparatus

200

100

- 50

Ec

o

«

c

«

Q.

Q. 10 -

><

(0a

CO

0.1

(a)

Pa : 3.0MPa

1.5MPa

g

0.1 MPa

D : 0.65 mm

IVD- 2

Tf o room temp.

A 50 °C 100 »C

T> 1 50 °C

J L I I I I I J L

0.5 1 2 3 4 Time (msec)

60

50

^ 40

30

a

<0 20

10

0.

)

(b)

i

o

o

ir

D : 0.65 mm

UD : 2

Tf o : room temp.

A : 50 °c

: 100 °c

<r : 1 50 °C

Pa : 0.1 MPa

1.5 MPa

3.0 MPa

.0.5 1 Time ( msec I

3 4

Fig. 2 Influence of ambient pressure and fuel temperature on spray

tip penetration(a) and spray angle(b) ( Pi = 7 MPa )

287
200

100

££co

c

Q.

50

a. 10

a. 5

-

0.1

200

100

- 50

£

E

O

(0

0)c

Q.a 10

(0

(a)

Pa : 3.0 MPa

D : 0.65 mm

L/D - 2

Tf o room temp.

A 50 °C 100 »C

ti- 1 50 °C

J L I I I I I I J L

60

50

40

c

30

(0c

a. V) 20

1 0

(b)

D : 0.65 m m

L/D : 2

Tf o : room temp.

A : 50 °C

: 100 °c * : 1 50 °c

Pa 0.1 MPa

1.5 MPa

3.0 MPa

'AA

O

A A^D

• V

0.5 1 Time (msec

I

J I I I 1 I I 0.1 0.5 1

Ti me ( msec i Fig. 3 Influence of ambient pressure and fuel temperature on spray

tip penetration(a) and spray angle(b) ( Pi = 14 MPa )

(a)

Pa: 3.0 MPa

0.1 MPa

a

(0

6

D : 0.65 mm

UD- 2

Tf o room temp.

A 50 °C a 100 °c

1 50 °C

J I

I I I I I

60

50

40

U>c 30

(0a« 20

10

(b)

D : 0.65 mm

UD : 2

Tf o : room temp.

A : 50 °C

: 100 "c

: 1 50 °c

Pa: 0.1 MPa

3.0 MPa

J L I I I I I I 0.1 0.5 3 4 0.1 Time (mseci

0.5 1 Time ( msec

)

Fig. 4 Influence of ambient pressure and fuel temperature on spray

tip penetration(a) and spray angle(b) ( Pi = 21 MPa )

288
01 005 0 15 04 lO(msec)

(b) p- ,21 MPa

Photo. 2 Enlarged photographs of spray ( Pa = 0.1 MPaJf = room temp.)

289
Tf;

Room

Temp.

Tf;

I50°c

0.7 (msec)

Photo. 3 Enlarged photographs of spray ( Pi = 14 MPa,Pa = 0.1 MPa )

4. Nagai,N. Sato,K. and Lee,Ch.W., " Atomization Characteristics of Superheated

Liquid Jets", ICLASS - 85, pp. VB/3/1-VB/3/11 ( 1985 ) 5. Reitz, R.D., "Atomization and Other Breakup Regimes a Liquid Jet", Ph.D. Thesis,

Princeton University, ( 1978 ) 6. Newman, J. A. and Brazustowski,T.A. , " Behavior of Liquid Spray at High

Pressures.", AIAA Journal Vol.8, No.l, pp. 164-165 ( 1970 ) 7. MacCarthy, M.J. and Molly, N.A., " Review of Stability of Liquid Jets and

the Influence of Nozzle Design.", Chem. Eng. J. Vol. 7, pp. 1-20 ( 1974 ) 8. Pai, B.U. and Nijaguna, B.T., " The Characterization of Spray" Proc. 2nd

ICLASS pp. 29-35 ( 1982 ) 9. Aral, M., Tabata, M., Hiroyasu, H. and Shimizu, M., " Disintegrating Process

and Spray Characterization of Fuel Jet Injected by a Diesel Nozzle", SAE Paper

No. 840275 ( 1984 )

f

290
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 28

MODELING DreSEL ENGINE SPRAY VAPORIZATION AND COMBUSTION

M.A. Gonzalez* and R.D. Reitz^

*Intevep S.A., Venezuela

^Engine Research Center, University of Wisconsin, Madison, WI, U.S.A.

ABSTRACT

Diesel engine in-cylinder processes have been studied using computational models with particular

attention to spray development, vaporization, fuel/air mixture formation and combustion in conditions

of high temperature and high pressure. A thermodynamic zero-dimensional cycle analysis program

was used to determine initial conditions for multidimensional calculations. A modified version of the

time-dependent, three-dimensional computational fluid dynamics code KIVA-II, with a detailed

treatment for the spray calculations and a simplified model for ignition, was used to determine details

of the closed cycle events. These calculations were used to obtain an understanding of the potential

predictive capabilities of the models. It was found that there is a strong sensitivity of the spray

calculations to numerical grid resolution. However, if proper grid resolution is used, the spray

calculations were found to reproduce experimental data adequately for non-vaporizing sprays.

However, for vaporizing sprays in high temperature engine environments the computations

underpredicted measured gas phase (vapor) penetration results substantially. This underprediction of

spray penetration reduces the accuracy of combustion predictions greatly. The atomization drop size

was found to be a key parameter influencing spray penetration predictions and this indicates that

improved atomization models are needed for engine conditions.

INTRODUCTION

Detailed models of engine combustion are of interest because they give insight about the combustion

process. This information is helpful to aid engine development efforts. However, due to the complexity

of the controlling phenomena, models are still being formulated and they require detailed validations

before they can be used for performance predictions. This study describes initial attempts to assess the

capabilities of current spray and combustion models for diesel engine combustion.

The computations were performed using a modified version of the KIVA-II computer code [1]. In the

spray atomization submodel, drop parcels were injected with sizes equal to the nozzle exit diameter

following Reitz [2] and the subsequent breakup of the parcels and the resulting drops is computed using a

stability analysis for liquid jets. This model was chosen over earlier models because, in principle, it removes the need to specify an assumed initial dropsize distribution at the nozzle. The kinetics

chemistry submodel used for combustion considers a single step Arrhenius mechanism for the

stoichiometric reaction of the fuel. The pre-exponential factor, activation energy and component rate

constants are those given by Bergeron and Hallctt [3].

291
The engine computations were made for a Cummins NH engine for which extensive experimental data are available with measurements of injection characteristics, cylinder pressure, and flame

temperatures (Yan and Borman [4]). Results using a cycle analysis simulation program are available for

the same engine [5]. Data calculated using the steady state mode were used as input for the initial

conditions in the multidimensional calculations as shown in Table 1. The injection velocity data was

determined based on the injector link load vs. crank angle data assuming a constant nozzle hole

discharge coefficient equal to 0.7. The turbulent kinetic energy was initialized at intake valve closing,

using the equations of Grasso et al. [6].

Table 1 Engine Conditions

Cylinder Bore 139.7 mm

Stroke 152.4 mm

Compression ratio 13.23

Displacement 2.33 liters

Fuel unit injector

Number of spray nozzle orifices 8

Nozzle hole diameter 0.2 mm

Spray axis, angle from head 18 degrees

Combustion chamber Quiescent

Piston crown Mexican hat

Engine speed 1500 rev/

m

Overall equivalence ratio 0.6

Air flow rate 3.53 g/cycle

Fuel flow rate 0.144 g/cycle

Engine temperatures:

Cylinder wall

Cylinder head

Piston surface

Mass average gas temperature at JVC

Cylinder pressure at IVC

Swirl number

Fuel

Injection starts (BTDC)

Injection ends (ATDC)

Number of parcels injected

405 K

486 K

578 K

359 K

157.9 kPa

1.0

Tetradecane

18 deg

11 deg

500

For the engine computations the 3-D computational mesh had 15 x 4 x 18 grid cells (in the radial,

azimuthal and axial directions, respectively) at TDC. The radial spacing was 5 mm with 3 cells in the

squish region and the geometry was based on direct measurements of the piston profile. Other

computations were also made with a finer 31x8x15 grid to test the influence of grid spacing. The

calculations were started at inlet valve closing (150 deg BTDC) and considered a 45 degree sector of the

engine that included one of the eight spray plumes.

RESULTS

Engine Computations

Temperature contours and droplet locations (circles) projected in the plane containing the axis of the

spray at TDC are shown superimposed in Fig. 1. Three main regions can be identified: the first, located

closest to the nozzle, has the lowest temperature because it is where the spray vaporization occurs; the

second is the combustion region with the highest temperatures in the domain (also located close to the

injection nozzle) and; the third is the undisturbed gas phase region. Most of the injected drops reach

the critical temperature of 695 K for the tetradecane fuel soon after leaving the nozzle and then

disappear. Between 9 deg BTDC and TDC, the liquid penetration length remains approximately

constant, and the liquid phase is confined to the region close to the nozzle due to the fast vaporization

of the drops in the spray.

Two major problems are in evidence in these calculations. First, the flame location does not agree

with the experimental results of Yan and Borman [4]. Using a radiation probe mounted in the cylinder

head, they found that the flame reaches the piston bowl outer surface between 6 deg BTDC and 1 deg

ATDC. However, the calculated flame never reaches the field of view of this probe. Second, as shown

in Fig. 2, the calculated cylinder pressure (chain-dashed line) is substantially lower than the measured

cylinder pressure (solid line). The amount of fuel burned in the combustion chamber was too low, and

this explains the small energy release and subsequent small pressure rise. The amount of fuel

evaporated was close to the amount of fuel injected, indicating a fast vaporization of the fuel once it was injected in the combustion chamber. The vaporized fuel creates a rich region near the nozzle which

reaches the rich flammability limits. At first, combustion depletes the oxygen in this region.

Thereafter, the combustion is controlled by turbulent diffusion of oxygen into the flame zone. The

292
18.0

14.0 I I

I I I

. < .

I

I I

I I 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0

radial position (cm)

Figure 1. Cylinder gas temperature contours and drop locations (solid circles) in the plane mid-axis of

the spray at TDC.

diffusion process is comparatively slow and there are regions close to the nozzle where the oxygen

concentration is zero. Similar computations that show incomplete or slow rates of combustion have been presented by

Takenaka et al. [7] and McKinley and Primus [8]. These workers speculated that the problem was

related to errors in the drop drag coefficient or to inadequate numerical resolution. However, the

precise reasons for the shortcoming were not established definitively in those studies. Takenaka et al. artificially reduced the drop drag coefficient by a factor of five in their computations in order to

increase the spray penetration sufficiently to match experiments. McKinley and Primus introduced an

ad hoc correction to the drop-gas coupling terms in the conservation equations to include an effect of

numerical grid size. 30.0 -1

10000.0 1

eooo.o

I

6000.0 «

oa

^ 4000.0

2000.0

-^25.0

c

o

.u 20.0

15.0

P 10.0

Priem

Priem

Kivo

Kivo

0.1 m/s

100.0 m/s

0.1 m/s

100.0 m/s

Q.

O

5.0 -

0.0

-leO.O -120.0 -60.00.0 60.0

crank angl* (d*gr«*s)

120.0 0.0

0.0 5.0

Figure 2. Calculated and experimental cylinder

pressures using the atomization model [2] and

using instead an initial drop radius of 5 mm

(without drop breakup or coalescence)

10.0 15.0 20.0 25.0 30.0

time (ms)

Figure 3. Calculated drop radius using KIVA and

Priem vaporization models with and without

convcctive effects. Hcxadocane fuel. Initial

diametcr=50 |im, Pressure=4.0 MPa, Gas

temperature=800 K.

293
Several numerical experiments were conducted in the present study to try to identify factors which

might increase the spray penetration and improve the fuel distribution. Higher initial turbulent

kinetic energy, higher swirl ratio, reduced aerodynamic droplet drag due to the influence of

vaporization, reduced fuel vapor diffusivity, reduced fuel vapor pressure and higher enthalpy of

vaporization, did not show sufficiently large effects to be able to explain the discrepancies. The

influence of the chemistry model was also studied by introducing an artificial ignition delay, by

varying the kinetic constants, by changing the temperature dependency on the reaction rates and by

introducing the diffusion burning combustion model of Magnussen [9].

Although each of these modifications did influence the predicted rates of pressure rise and the peak

cylinder pressures, they did not consistently change the combustion chamber pressure, which was still underpredicted, or the spray penetration, which still remained low.

Based on these tests it was concluded that three areas needed further study: first, to determine if drop vaporization rates predicted in KIVA were consistent with those from other more comprehensive

models; second, to examine the influence of numerical error on the results and; third, to explore the

influence of the drop breakup process and, in particular, the dropsize predicted from the atomization

model. These aspects are discussed next.

Effect of Vaporization Model. A numerical experiment was conducted in which a single drop was

injected into a high temperature and high pressure environment. A cylindrical computational domain

of 4 cm diameter and 12 cm length was discretized using a grid of 20x1x60 cells with 1 mm radial x 2mm

axial spacing near the nozzle. The diesel engine environment at the time of injection was simulated

using air at 800 K and constant pressure of 4.0 MPa. The initial condition was selected to be quiescent to

remove any uncertainties associated with fluid motions. The 50 mm diameter hexadecane drop

initially was at 298 K and two cases were considered: 1) injection velocity of 0.1 m/s (for practical

purp»oses zero velocity) and 2) injection at 100 m/s.

The evaporation routine in KIVA uses the Ranz-Marshall correlation [1], with a correction for

convective effects. Computations with this model were compared with results obtained using the model

of Priem et al. [10] which includes a variable density term (which is important during the heat-up

period of the drop) and a correction to account for superheating of the diffusing vapor around the drop.

The Priem model was also modified to include the effect of convection. Several combinations of

dropsizes, injection velocities, gas temperatures and pressures were compared in the study and each case reached the same qualitative conclusions as those that are given in Fig. 3. As can be seen in Fig. 3, although the evaporation routine in KIVA does not include the mentioned

corrections, it predicts lower vaporization rates than the more comprehensive Priem model. Thus, one

can conclude that the observed low spray penetration predictions in the engine computations cannot be

explained by any obvious shortcoming of KIVA's vaporization model, since the more comprehensive

model gives even faster vaporization under all conditions.

Effect of Grid Resolution. To assess the effect of numerical error, non-vaporizing sprays were also

studied to remove the coupling between drop breakup and vaporization. The experiments of Hiroyasu

and Kadota [11], with measurements of spray tip penetration using diesel fuel injected in nitrogen gas at 300 K and 3.0 MPa, was used for comparison. A computational domain of 4 cm diameter and 12 cm length

was discretized using a grid of 40x8x24 (0.5 mm radial x 5 mm axial). Diesel fuel was simulated using

tetradecane and the environment gas flow was quiescent. The initial injected drop radius was 150 mm

(equal to the nozzle hole radius) and the injection velocity was held constant at 90.3 m/s.

The results shown in Fig. 4 correspond to 4 ms after the beginning of injection. At this time the

experimental spray tip penetration was 7.5 cm and the spray angle was 13 degrees. As seen in the left- hand-side of Fig. 4, the computations gave a low spray tip penetration (about 6.4 cm) and significant

radial dispersion of the drops with about 40% of the injected mass outside of the measured spray angle.

The dispersion of the drops near the nozzle reflects errors in the calculation of the gas velocity in the

region of large velocity gradients near the nozzle exit.

Further calculations demonstrated that the grid spacing of 5 mm in the axial direction was too large.

Results obtained by increasing the grid resolution in the axial direction by using a grid of 0.5 mm radial x 2 mm axial resolution are shown in the right-hand-side of Fig. 4. Now the spray penetration is

294
10.0

12.0

10.0

I

• I I I -2-10 1 2

radial location (cm)

12.0 I I I I

• I -2-10 1 2

radial location (cm)

Figure 4. Influence of mesh refinement on spray p»enetration. Sprays shown at 4 ms after injection. Left:

0.5 mm radial by 5 mm axial resolution. Right: 0.5 mm radial by 2mm axial resolution.

Mill"'

CONTOUR LEUELS

0. 1D00D

0.30000

0.50000

D. 70000

D. 90000

1. 10000

1.30000

1.50000

1.70000

1.90000

2. 10000

2.30000

2.50000

2.70000

2.90000

3. 10000

3. 30000

3.50000

Figure 5. Predicted equivalence ratio contour on the plane mid-axis of the spray at TDC without

considering combustion.

295
increased to 8.5 cm and the percentage of mass outside of the spray angle is only 4.6%. The calculation

overpredicts the measured spray tip penetration somewhat, but dramatically improves the predicted

dispersion of the drops.

These results demonstrate a strong sensitivity of spray calculations to numerical grid resolution and

suggest the need for grid sensitivity analyses of engine calculations. Accordingly, the engine

computations were repeated using a fine grid with 31x8x15 cells and 8 planes in the squish region at TDC (a radial mesh spacing of 0.78 mm was used in the vicinity of the nozzle). It should be noted that

the use of such a fine grid is prohibitive for routine calculations due to the long computational times.

The complete run took 4.0 hr of cpu time on a Cray Y-MP computer compared to 0.9 hr for the 15x4x18

grid. In order to decouple the contributions, these computations not consider combustion.

Figure 5 shows the equivalence ratio contours in the mid-axis plane at TDC. The fuel vapor tip (e.g.,

equivalence ratio of 0.1 contour) is located about halfway between the nozzle and the piston bowl.

These results again show low fuel vapor penetration, even in the absence of combustion. This indicates

that grid resolution effects cannot explain the observed low penetration in the engine's highly

vaporizing environment.

Effect of Initial Dropsize. The formation of small drops with their high vaporization rates and

high deceleration due to their strong interaction with the gas could be responsible for reduced spray

penetration. To explore this possibility, the experiments of Kamimoto et al. [12] were used. These

consist of n-tridecane injections into nitrogen gas at 875 K and 3.0 MPa. In the experiments gas and liquid

phase penetrations were determined from Schlieren photographs.

A cylindrical computational domain of 4 cm diameter and 12 cm length was discretized using a grid of

20x8x30 cells (0.5 mm radial and 2 mm axial spacing near the nozzle). The environment was quiescent

and the initial injected drop radius was 80 mm (equal to the nozzle hole radius). Three different

experimental conditions were studied with injection pressures of 26, 50 and 80 MPa.

The calculated gas phase penetration seen in Fig. 6 (arbitrarily represented using that axial location

where the fuel vapor contour reached 0.002 kg/m-^) shows a lower gas penetration than the experiment;

and the difference between them increases with increasing injection pressure and consequently higher

injection velocities. These high temperature and high pressure results confirm the low spray

penetration predictions found in the engine computations and suggest the need for a revision of the spray

calculation in order to identify the source of the discrepancy.

0.0 1.0 2.0 3.0

Time (ms)

Figure 6. Calculated and experimental gas (vapor) penetration for different injection pressures.

Experimental data from Kamimoto et al. [12].

296
The spray atomization model predicts the formation of very small drops at the high injection

velocities of interest in engines. For example, for the case with injection pressure of 50 MPa in Fig. 6, the

radius of the injected drops was 80 |im and the calculated wavelength of the fastest growing surface

waves was between 0.45 and 0.80 |im. This results in the formation of a spray of drops with sub-micron

sizes since the initial drop size is assumed to be of the order of the unstable wavelength in the

atomization model [2]. Similar small spray drop sizes are predicted using other atomization models

such as the Taylor Analogy Breakup (TAB) model of O'Rourke and Amsden [13]. These small drop sizes

lead to very fast vaporization rates with corresponding low penetrations.

The marked influence of atomization drop size on penetration was demonstrated by injecting drop

parcels with a fixed initial drop radius of 5 |im and by also suppressing drop breakup, collisions and

coalescences in the calculation. The results given in Fig. 7 compare the predicted spray p>enetrations in

two engine computations made using the atomization model for the initial drop size (top) and the 5 |im

initial drop size case (bottom). As can be seen, greater spray penetration is achieved with the the use

of the larger initial drop size.

The improved penetration also has a beneficial effect on combustion predictions. This can be seen in

the cylinder pressure comparison shown in Fig. 2 for the case of 5 |im radius injected drops (dashedline). There is an increase in the calculated peak pressures and a different rate of pressure rise. The

improvement results from an improved fuel vapor distribution and consequent higher energy release.

From Fig. 2 it can be inferred that a bigger initial dropsize would be able reproduce the experiment even

better. These results indicate the need to improve the prediction of dropsizes from atomization models.

18.0 n

o

X

O 14.0 I 1 1 I I I I I I I

1

I I I I I I I I I

I

I I I I I I I I I

1

I I I I I I I I I 1

0.0 2.0 4.0 6.0 8.0

radial position (cm)

Figure 7. Drop locations in the plane mid-axis of the spray using the atomization model [2] (top) and,

instead, an initial drop radius of 5 |im (without breakup or coalescence in the computation - bottom).

297
CONCLUSIONS

The study found difficulty in reproducing measured diesel combustion trends. Various possible causes were examined and the importance of the atomization model was pinpointed. The results showed:

1) The spray calculations were sensitive to numerical grid resolution. However, if proper grid

resolution was used, the calculations reproduced experimental data adequately for low injection

pressures and non-vaporizing sprays.

2) Spray calculations in high temperature environments were found to underpredicted measured gas

phase penetration. This trend was found in both constant volume bomb and engine computations.

3) The underprediction of spray penetration was found to reduce the accuracy of engine combustion

predictions considerably.

4) The details of the spray vaporization model were unable to explain the low spray penetrations.

5) The sensitivity of the results to the size of the injected drops suggests that improved atomization

models are needed for engine conditions. Experimental data is needed under highly vaporizing

conditions to validate new models (pressure 3.0-5.0 MPa, temperature >700 K, injection velocities (100-

300 m/s), including measurements of fuel vapor concentration, dropsizes and liquid core penetration.

ACKNOWLEDGMENTS

The authors thank Gary Borman for helpful comments. Support for this work was provided by Intevep

S.A. Venezuela, by Army Research Office contract DAAL03-86-K-0174 and by NASA-Lewis grant NAG

3-1087. Funding for the computations was provided by Cray Research, Inc. and the San Diego

Supercomputer Center.

REFERENCES

1. Amsden, A. A., O'Rourke, P.J., and Butler, T.D. "KIVA-II: A computer program for chemically

reactive flows with sprays," Los Alamos National Laboratory Report LA-11560-MS (1989).

2. Reitz, R.D. "Modeling atomization processes in high-pressure vaporizing sprays". Atomisation and

Spray Technology, Vol 3, pp. 309-337 (1987).

3. Bergeron, C.A. and Hallett, W.L.H. "Ignition characteristics of liquid hydrocarbon fuels as single

droplets," The Canadian Journal of Chemical Engineering, Vol 67 (1989).

4. Yan, J. and Borman, G.L. "Analysis and in-cylinder measurement of particulate radiant emissions and

temperature in a direct injection diesel engine," SAE Paper 881315 (1988).

5. Lei, N. "A cycle simulation program for the dynamic operation of a single cylinder direct injection

diesel," MS Thesis, Department of Mechanical Engineering, University of Wisconsin-Madison

(1988).

6. Grasso, F., Wey, M.-J., Bracco, F. V., and Abraham, J. "Three-dimensional computations of flows in a stratified-charge rotary engine," SAE Paper 870409 (1987).

7. Takenaka, Y., Shiozaki, and Joko, I. "Numerical Analysis of the Fuel Spray Formation in DI Diesel

Engine," SAE Paper 902078 (1990).

8. McKinley, T.L. and Primus, R.J. "Three-Dimensional Calculations of Air Motion, Sprays, and

Combustion in a Quiescent Direct-Injection Diesel Engine," ASME Paper 90-ICE-2 (1990).

9. Magnussen, B.F., and Hjertager, B.H. "On mathematical modeling of turbulent combustion with

special emphasis on soot formation and combustion," 16th Symposium on Combustion, The

Combustion Institute, Pittsburgh, PA (1976).

10. Priem, R.J., Borman, G. L., El Wakil, M.M.O., Uyehara, A. and Myers, P.S. "Experimental and

calculated histories of vaporizing fuel drops," NACA TN 3988 (1957).

11. Hiroyasu, H., and Kadota, T. "Fuel droplet size distribution in diesel combustion chamber," SAE

Paper 740715 (1974).

12. Kamimoto, T., Yokota, H., and Kobayashi, H. "Effect of high pressure injection on soot formation

processes in a rapid compression machine to simulate diesel flames," SAE Paper 871610 (1987).

13. O'Rourke, P.J. and Amsden, A.A. "The TAB Method for Numerical Calculation of Spray Droplet

Breakup," SAE Paper 872089 (1987).

298
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 29

MATCHING OF ENGINE CHAMBER AND FUEL SUPPLY

SYSTEM VIA KIVA-H MODELING

K,C. Tsao, Y. Dong and Y. Xu

Dept. of Mechanical Engineering

University of Wisconsin-Milwaukee

Milwaukee, WI, U.S.A.

ABSTRACT

A comprehensive model --KIVA-I I —code based on the numerical calculation

of transient, two- and three-dimensional, chemically reactive fluid flow with

sprays was used to study the spray penetration in three classic types of combustion chambers of direct injection diesel engine. The effects of changes

in several operating parameters such as the speed of the engine, the fuel

mass per cycle and the initial swirl ratio were accounted for, as were the

effects of changes in the clearance and the geometry of the piston bowl. In

addition, the effects of different injection parameters viz. initial injection

velocity at the nozzle, injection rates, nozzle diameters and Sauter mean

radius of the droplets were investigated in relation to spray penetration and

spread and the relationship between the spray and the shape of the chambers.

The initial injection velocity, the Sauter radius and the rate of injected

fuel per cycle all play a significant role in spray penetration and air fuel

mixing. Moreover, the nozzle orifice diameters and duration of injection do

affect the spray penetration, spread and air fuel mixing though they are

governed predominantly by the other parameters such as injection pressure,

mass injection rate, etc. The velocity and distribution of gas motion

described by the initial swirl ratio and clearance height of the piston also

affect the spray characteristics, especially in the outer peripheral region

of the spray cone.

INTRODUCTION

Matching of injection characteristics, air motion and combustion chamber

geometry are of prime importance in the development of more efficient and

cleaner emission direct injection diesel engines. Direct injection diesel

engines offer great promise in terms of both efficiency improvement and pollution control. The important factors which influence the open-chamber engine

performance are the fuel penetration, vaporization and mixing. This means

that the rate and completeness of combustion are controlled by the way the

fuel vapor gains access to the available air, which is in turn determined by

such factors as the trajectory of the sprays, the droplet velocity and the

size distributions, and the nature of the prevailing gas flow field. The

299
spray trajectory is strongly influenced by the gas flow, especially after

injection is terminated.

Injection rate is an important parameter as it has a strong influence on

the rate of pressure rise after ignition. Injection duration is another,

particularly at full load, since it influences engine performance, the fuel

economy and the density of smoke directly. Fine atomization of fuel is not

sufficient in achieving an efficient combustion unless proper propagation of

sprays in the engine combustion chamber is attained. So, both the atomization

and penetration of sprays play a key role to reach a desired pattern of distribution and a shorter delay of the injected fuel. In addition, the penetration of finely atomized fuel is required to match that of air motion and

combustion chamber shape under varying fuel injection characteristics.

Recognition of these problems has prompted the application of multidimensional modeling techniques to the in-cylinder processes in D.I. engines. The

results presented in this paper were obtained using KIVA-II computer code of

Los Alamos National Lab. [!]• The work reported here attempts to introduce

some cohesion into the studies of sprays in engine chambers by carrying out a limited variation of engine parameters. Three different shapes of omega

chambers, prevailing in modern high speed D.I. diesel design, are used as

typical configurations. The effect of engine speeds, initial swirl ratio,

and clearances in the normal range of an operating engine is assessed for a bowl shape. Other fuel-supply parameters such as the size of injector nozzle,

the duration of injection, the velocity at the injector hole exit and the

spray mass per cycle are assessed for a bowl shape. The detailed parameters

incorporated for discussion here are shown in Table 1.

METHOD OF CALCULATION AND CASES CONSIDERED

In the swept and clearance volumes, the grids expand and contract with the

piston motions. And the number of the axial meshs decreases correspondingly

as the piston moves toward TDC. The process of contracting and expanding is

calculated from 90° BTDC to 60" ATDC. All operations of the KIVA-II program

(including compiling, running and data processing) are carried out on a Cray

Y-MP supercomputer.

Simulations of the in-cylinder flow were carried out on various combinations of parameters shown in Table 1. The variations with * mean that they

are a baseline case in calculation. The main structure parameters and operatTABLE 1. Parameter Variations

Variations Levels

Engine speed (r.p.m.)

Clearance (cm)

Geometries of bowl

Diameter of injector nozzle (cm)

Velocity at injector hole (cm/s)

Spray mass (g/cyc .

)

Initial swirl ratio

Sauter mean radius (cm)

Duration of injection (CA°)

1500 2000 2500* 3000 3500 4000

0.12 0.23* 0.38

#1* #2 #3"

0.018 0.021 0.024* 0.027 0.031

4000 16000* 28000 40000

0.0116 0.0232* 0.0348 0.0464

0.0* 1.5 2.0 2.5 3.0

1.5x10-4 5.0x10-4* 1.0x10-3

10 15* 20 24

300
Fig. 1. Spray in bowl #1. Fig. 3. Spray in bowl #3.

ing conditions of the engine used were given in reference 5. The engine is

typical of a modern high speed diesel engine for off-road applications. In

all cases, the calculations are commenced at the inlet valve closure, at which

stage the air is assumed to have zero radial component and the axial component

is assumed to vary linearly between the cylinder head and the piston crown.

All wall temperatures, including the cylinder wall, cylinder head, piston

face, valves and other exposed surfaces, are taken as 400 K.

RESULTS

The evolution of the spray fields will now be presented in the form of

plots at selected crank angles spanning a period between 40° BTDC and 20" ATDC.

The spray will be shown as viewed from the side through the full depth of

field.

Baseline Results

The calculated spray obtained using bowl #1 are depicted in Fig. 1. The

spray field in Fig. 1 shows that during the injection period, spray is developing as the crank angle increases and there is a "solid core" in the spray which

is seldom influenced by the gas flow. However, as soon as the injection ends

(Fig. 1(d)), the velocity of the spray decreases greatly. Afterward, the spray

is influenced by the gas flow, which includes the entrainment and dispersion

of the gas and the spray. Here the fuel spray, under the influences of squish

flow and the presence of a curved bowl surface, is deflected downward into the

piston bowl. At the same time the entrainment of gas causes the fuel liquid

to be drawn up from the bowl near the center line. Only very early in the

injection period does the bulk of the spray move in the injection direction.

Beyond 15°, BTDC the spray is pushed downward by the stronger squish motion,

so that the spray tends to spread out across the piston bowl volume. It could

be presumed that for a bowl with a smaller squish volume the fuel liquid would

tend to spray up to the top of bowl.

301
The Geometry of Chambers viz. Spray

The matching of the injection spray and the geometry of chambers is of

prime importance. Under the conditions that all fuel parameters, gas flow and

operating parameters are kept constant, only the angle of direction of spray

is adjusted, the spray in two different geometric chambers #2 and #3 is illustrated in Fig. 2 and Fig. 3. For chamber #2, comparing with chamber #1 (see

Fig. 1), the d/D keeps constant; however, H/D decreases about 55%. Because

the fuel is injected into a narrow space between the cylinder head and the

bottom of the bowl, the entrainment of gas is not sufficiently strong in the

direction of the core of spray. It is evident that the spray penetration in

chamber #2 is less than that in chamber #1 at the same crank angle. The

situation is more evident in chamber #1 . So, some measures should be taken if

a shallow w chamber is preferred. The measure includes increasing the

initial momentum of spray, or adjusting the angle of spray so that the mixing

and atomization of spray can be performed well.

Effect of Initial Swirl Rat io

The effects of the different swirl levels on the droplet trajectories and

locations are shown in Figs. 4 and 5. The amount of initial swirl level would

affect the gas velocity profiles and the turbulence levels in the bowl. It is

noted that the interaction of the swirl and the squish has a significant effect

on the axial and radial flow patterns and results in a system of two counterrotating toroidal vortices in contrast to the no swirl case [5]. The direction

of velocity in the region between the vortices is upward towards the cylinder

head in the case of stronger initial swirl. In contrast, it is downward

towards the bottom of the bowl in the case of no initial swirl. So, from Fig.

5 it is seen that the location of droplets is on the top of bowl if there is

swirl. On the other hand, the existence of initial swirl would accelerate the

dispersion of spray well, which is favorable to mixing and vaporization of

fuel spray with air in the chamber.

Fig. 4. Influence of initial swirl Fig. 5. Influence of initial swirl

ratio on spray SR=0.0. ratio on spray SR=2.5.

Effe ct of Clearance on Spray

The difference of clearance height means the change of the compression

ratio. The range of the clearance height is from 0.12 cm to 0.38 cm. In

fact, the compression ratio is altered from 19.1 to 14.0 respectively. On the

other hand, the change of the clearance height also results in the difference

of squish volume from 0.272 cm^ to 0.862 cm^.

302
A larger clearance means a lower compression ratio, which results in a lower pressure and temperature at late duration of the compression stroke.

This case is out of favor from the point of view of the evaporation and atomi

zation of spray. By comparing Figs. 6 and 7, this case is shown obviously.

In addition, large a clearance also decreases the squish intensity, which

would have an unfavorable influence on the distribution of spray in the bowl

after injection ends.

Engine Speed Effec t

The engine speed range was chosen to cover from 1500 to 4000 r.p.m. which

is the normal operating speed of open chamber diesels. The engine speeds have

a significant effect on the fuel-supplied system such as fuel mass per cycle,

injection duration, etc. Comparison of Fig. 8 and Fig. 9 indicates that an

increase of engine speeds lowers the penetration distance which appears to be

not in favor of engine designer. The results hint that when a high speed

engine is designed, corresponding measures, such as increasing injection

pressure, selecting a suitable diameter of injector nozzle and so on, should

be redesigned.

29.9" BTDC

'23.9° BTDC

19.1" BTDC

Fig. 6. Influence of clearance height Fig. 7. Influence of clearance height

on spray at CH=0.12 cm. on spray at CH=0.38 cm.

Fig. 8. Fuel penetration at Fig. 9. Fuel penetration at

n=1500 rpm. n=3000 rpm.

Effe c t of Fuel Sys tem Parameters

The penetration and atomization of spray are influenced by many fuel system parameters such as injection pressure, diameter of pump shaft, structure

303
and size of injector, etc. Here, the effect of several essential parameters,

the diameter of injector hole dp, the injection velocity v-j , the fuel mass

per cycle, the Sauter mean radius and the injection rate will be discussed.

For baseline condition and at crank angles of 29.7°, 24.6° and 21.2° BTDC the

penetration of fuel spray column is shown in Fig. 10 versus the diameters of

injector hole. Figure 10 indicates that

during injection, both the larger and

smaller diameters of nozzle would decrease the penetration of spray, when all

other conditions are unchanged. This is

because smaller nozzle diameters would

have the spray more atomized, and hence

shorter penetration.

The injection velocity at the nozzle

has a significant influence on penetration and atomization of the spray. An

increase of injection velocity from 40

m/s. Fig. 11(1), to 80 m/s. Fig. 11(11),

i

z

Fig. 10,

DIAMETER OF INJECTOR NOZZLE (mm) Influence of nozzle dia.

on penetration of spray.

shortens the injection duration, increases penetration

dispersion and hence better overall performance.

and offers greater

29.9° BTDC

— • (b) -

•

24.5" BTDC

19.7" BTDC

Fig. 12, Spray field at Ms

0.0116 g/cyc.

14.2° BTDC

29.9° BTDC —^ (») —

'24.5'' BTDC

V^ (b) —y

19.7° BTDC

Fig 11. Injection velocity at nozzle. Fig. 13. Spray field at Ms

0.0452 g/cyc.

304
The spray mass per cycle, which symbolizes the engine load, is an

important operating parameter. Two classical instances of spray mass are

shown in Figs. 12 and 13 for 0.0116 and 0.0452 g/cycle respectively. It is

seen, obviously, that an increase of spray mass per cycle would result in

greater penetrai:ion and density of the spray. The larger droplets penetrate

farther than the smaller droplets and hence the leading region of a spray

plume is enriched with large droplets. The tip droplets, i.e. those with the

farthest axial penetration, are always moving and following the movement of

gas vortex vector in the bowl.

The initial Sauter mean radius was estimated according to the theory

developed by Taylor [6] who studied the rate of growth of the perturbations of

planar liquid surfaces induced by co-flowing gas. The theory states that the

initial droplet radius is directly proportional to the liquid surface tension

and wavelength of the fastest growing surface wave, and inversely proportional

to the gas density and the square of the relative velocity between the gas and

the liquid. The investigation by Bracco [4] showed that stable droplets are

of the order of 5 to 10 ym depending on the relative velocity between the

gas and the liquid phases. For injection velocities larger than 100 m/s,

droplets larger than 10 ym in radius will breakup into smaller droplets.

The spray penetration viz. the initial Sauter mean radius is shown in Fig.

14. As expected, the penetration increases as the initial Sauter mean radius

increases.

The injection rate is a function of the total injected mass and the

duration of injection. For a fixed amount of fuel injected, the penetration

of spray versus the duration of injection was indicated in Fig. 15. It is

evident that an increase in injection rate would result in an increase in

penetration and vice versa.

Fig. 14

CRANK ANGLE BEFORE TDC

Effect of Sauter mean radius

o

U

2.4-

a.

Fig. 15.

CRANK ANGLE BEFORE TDC

Influence of injection

duration.

CONCLUSIONS

The engine designer is confronted with a formidable task to reconcile and

optimize a set of engine parameters relating to engine performance and exhaust

emission. These include, insofar as gas motion is considered, the induction

and combustion chamber configuration which characterize the in-cylinder gas

motion to facilitate air fuel mixing. On the fuel system, the main consideration is the performance of an injector which would supply a fuel spray of adequate penetration and dispersion to match that of the in-cylinder gas behavior,

and vice versa. This paper has demonstrated the following observations.

305
1. The solid core of a fuel spray plume is seldom influenced by the incylinder gas motion during the injection period. However, after injection

terminates, the spray profile is significantly affected by the gas flow,

especially the engine squish velocity.

2. A shallow chamber has a minimal effect on spray penetration. Additional

measures must be incorporated in the design for proper mixing and dispersion between fuel and air.

3. As anticipated, engine initial swirl and squish velocity would alter the

direction of sprays and subsequent improvement of fuel air mixing.

4. An increase of engine speed reduces the fuel penetration distance.

5. Fuel injection parameters, such as injection pressure and duration, exert

a greater influence on the spray pattern than the in-cylinder gas motion.

ACKNOWLEDGEMENT

The authors wish to thank the National Center for Supercomputing Applications, Champaign, Illinois, for a grant on KIVA-II applications.

REFERENCES

1. Amsden, A. A., O'Rourke, P. J., and Butler, T. 0., "KIVA-II: A Computer

Program for Chemically Reactive Flows with Sprays," Report LA-11560-MS,

Los Alamos National Laboratory (May 1989).

2. Gosman, A. D., "Multidimensional Modeling of Cold Flows and Turbulence in

Reciprocating Engines," SAE 850344 (1985).

3. Watkins, A. P., Gosman, A. D. and Tabrizi, B. S., "Calculation of Three

Dimensional Spray Motion in Engines," SAE 860468 (1986).

4. Bracco, F. V., "Modeling of Engine Sprays," SAE 850394 (1985).

5. Tsao, Keh C, Dong, Yu and Xu, Yong, "Investigation of Flow Field and

Fuel Spray in a Direct-Injection Diesel Engine via KIVA-II Program," SAE

901616 (1990).

6. Reitz, R. D. and Bracco, F. V., "Mechanism of Atomization of a Liquid

Jet," Physics of Fluids, Vol. 25, pp. 1730-1742 (1982).

306
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 30

EXPERIMENTAL INVESTIGATION ON MICRO-EXPLOSION OF

EMULSIFIED DIESEL OIL BY HOLOGRAPHY

X, Hang, S. Yunbiao, Z. Chongji and M. Yuanji

I.e. Engine Research Institute

Dept. of Energy Engineering

Zhejiang University

Hangzhou, P.R. China

ABSTRACT

In this study, appreciating the single-pulsed in-line

Fraunhofer holographic system and the high speed schlieren

photographic technique , we have studied the atomizing and

evaporizing characteristics of emusified diesel spray in a high

temperature combustion bomb.

During the research, we have obtained both the micrographes

of oil droplets in the spray field and the macrographes of the

spray development process. While comparing the pure and water

emulsified diesel oil, we found the emusified oil had better

atomizing characteristics than the neat oil. "Fragmentation"

atomizing phenomenon has been discovered in emulsified spray. The

authors propose, according to these features, the water-oil

uncontinuing internal phase boundary decays the uniform of

liquid. We think the diverging and disturbing ability of the

superheated water addition in the emulsified oil contributes to

the improvement of spray atomization.

In addition, we also found a distinctive spray distribution

of a pintle diesel nozzle. That is at the outer region of the

spray the droplets have larger size than those in the inner

region. This tendency is more stronger in low environmental

temperature than in high one.

INTRODUCTION

Diesel oil emulsified with water has been widely studied as

an effective method to reduce both the fuel consumption and

exhaust emissions of the engine. According to the feature that

the combustible air-fuel mixture of diesel engine is formed in

the combustion chamber, spray atomizing characteristics become

very important to diesel engine performance. So it is necessary

to have a further investigation on the spray atomization of

emul s i f i ed oil.

Formerly, Ivanov(l) found the "micro-explosion" phenomenon

of suspended fuel emulsion droplet by high speed motion

photography. Gal 1 ahal 1 i (2 ) studies the flame structure of

307
unsupported single drop for both neat and emulsified oil.

LasheracO) has also observed the micro-explosion of emulsified

free droplet. Although there have been a number of studies on

the i so 1 a ted . emul s i f i ed droplet, as mentioned above, that is not

enough to guide the application of water emulsified oil in diesel

engines. We say this because the condition of single free droplet

is quite different from that of particles in diesel spray, no

matter in particle size, density or in droplet life time. On

account of this, we have paid more attention to the spray

particle field of emulsified oil in diesel engine.

In our study, the high speed schlieren photography and inline Fraunhofor holography have been adopted to observe the spray

development process macrocop i ca 1 1 y and the particles in spray

field microscopically. We found that emulsified oil had some

different features of spray atomization from the neat oil.

EXPERIMENTAL APPARATUS

High Temperaiure Combuslion Bomb

Figure 1 shows a scheme of the high temperature combustion

bomb used in this study. Two quartz windows are installed in both

sides of bomb for optical access.

Nitrogen was used as' the surrounding gas. This made it

possible to observe the atomizing and vaporizing process in

detail. Fuel was injected by a standard diesel injector with

pintle diesel nozzle(hole diametre 1.5mm, spray angle 15deg,

openning pressure 20.0Mpa). The neat fuel used was No.O diesel

oil with a density of 0.819g/cm3 and cetane number of 55.

1

Fio.l Hioh Teffloerature Boisb Pressure Vessal 2. Intake Valve

3. Thermal couDle

5. Cool Water

Jacket

B. Glass Holder

4. Fuel Injector

b. PresureTransducer

7. Eisission Valve

9. Quarie Window

FiQ.2 Layout of In-Line Fraunhofer

Holooraohic Recordino Svsteui

1. Pulsed Laser 2. Svnchronis/a Delavino

3. The Esoander SvstefB

4. Combustion Bosb 5. Fuel Injector

6. Neddie Lift 7. Isaoe Lens

Transducer 8. Holooraohic Plate

9. N2 bottle 10. Workino Plat

QEiicai Arrangemeni of Recording System

The layout of the holographic system and the high

temperature combustion bomb is shown in Figure 2. An in-line

holographic arrangement using a single-pulsed ruby laser was

employed. The laser with a wavelength of 694. 3nm, an output of

lOOmj and a pulsed width less than 50ns.

308
A set of image lens[7] is placed just before the plate not

only to premagnify the conjugate image of the particles recorded

in the plate, but also to eliminate the aberrations which will

distort the image greatly in mi cro-pho tography , So in our system,

the resolving power is high enough to record a 3um particle.

Hoiographic Reconslruciion and compuier processing sysiem

Fig. 3 shows the devices used in holographic reconstruction.

As mentioned in the Figure, a He-Ne laser beam is used as

reconstruction light. Spray field hologram which has been put on

the three-d i mens i o-n displacement is reconstructed by the optical

system. A vedio camera catches the holographic image. A computer

imaging system is used to process and count the size and

distribution of droplets in spray field. The processing of

droplets is based on the theory of grayscale mathematical

me thod ( 4 ) .

VtJ.o

FiQ.3

1.3.

5.

7.

Hclooraohic Reconstruction

& Processino svsteis

He-Ne Laser 2. Soacial Filter

Holooran Plate 4. Ifflaoe Lens

PC Vision 6. Cofflouler Three Disiension 8. Horn tor Disolacesient

FiQ.4 Kicrooraoh of Internal

Phase of Eaulsion

EXPERIMENTAL RESULTS AND ANALYSIS

We have adopted No . 0 diesel oil. Water/oil emulsion with 15

percent of water content by volume which was provided by adding

surfactants of SPAN 0.1% and MAPE 0.2%(volume fraction). A 400X

enlargement microscopic photograph showed a fine and stable

dispersion of water addition. The particles of internal phase are

as small as 2-7um(see Fig. 4). Both the dillution and electric

resistance tests manifested the emulsion were water-in-oil

s t rue tur e

.

Experirnent ResuHs in Ambieni Condition

In ambient condition, we studid the spray atomizing feature

of both the neat and emulsified oil. High speed schlieren

photographes showed the emulsified spray had a smaller spray

angle than that of the neat oil, which were shown in Fig. 5.

While under the same condition, holographes manifested the

and a smaller S. M. D.(Sauter Mean Diameter) of droplet size than

the pure diesel oil. Table 1 showed the S.M.D. results counted

statistically by computer.

309
(a) Pure Oil Sorav (b) Ediulsif ied Oi 1 Sorav

FiQ.5 High Soeed Schlieren Cineiatooraohv

Siauences of Soav Develoosieni

(Aabient Condi tion,Z000f/s)

Time Delay

From Injection

0. 7ms 0. 9ms

Emul s ion

( S . M . D . ) urn 24,12

Pure Oil

(S.M.D. ) urn 25.27 30. 12

Table 1 Results of S.M.D.

The atomizing features of the emulsion, in author's opinion,

are closely related to its internal phase structure and physical

properties. In ambient condition, the emulsion is more viscous

than the pure oil, and this causes the emulsion has a more narrow

spray angle. But, on the other hand, the water-oil uncontinued

phase boundaries bring the more uniform and smaller emulsified

spray distribution. We can find the uncontinuing property of

internal phase of emulsion from the mi cro-pho t ograph . (Fi g . 4 ) . I

t

is obvious that the sheer stress between these water-oil

uncontinuing surface must be weaker than those of the continuing

phase in pure oil. So fragmentation of droplets occured more

frequently at these micro-regions.

Therefore, when the emulsified spray moves in the air with

high speed, the spray will be disturbed and become instability.

We assume that the boundaries of uncontinued water-oil surfaces

at those internal phase are easier to be broken and cause further

fragmentations of particles and a fine atomization.

310
Aioniizaiign and Va^o rizaiion of Emulsion in High Temneraiure

Environmeni

High speed schlieren photographes in 573K and 0.9Mpa

environment of pure and emulsified oils showed a very interesting

experimental phenomenon,

emulsified spray angle was

certain period as shown in

disruptively and developed

0 i 1 (Frame No . 4 , 5 , 6 ) . In high temperature,

At the beginning of injection, the

smaller than the pure one, but after a

Fig. 6, the emulsified spray expanded

to a larger spray angle than the pure

the

high tempera ture

holography showed microscopically

emulsified particles had two distinctive features.

One was the fragmentations of some of large droplets in the

surrounding parts of spray. The other was the existance of many

small particles in outer region of spray, where used to be

dominated by larger size droplets. Fig. 7 is the pictures of

emulsified droplets taken in different

environments.

We haven't found the same phenomenons

pure 0 i 1 spray

.

In high temperature, the appearance of a lot of small

particles in the emulsified spray has improved the atomizing

quality. The water addition in spray is in overheated state after

having been injected into the high temperature bomb. The puffing

of vaporized water might be strong enough to cause fragmentations

of' large particles directly as "micro-explosion", or if not as

strong as "micro-explosion", might intense the internal

disturbance of droplets which are in high speed. In the mean

time, the micro-structure of uncontinued internal phase between

water addition and oil might also optimize the spray atomization.

mentioned above in

Spray Emul s ion Neat Oil

Middle

S.M.D. (um)

11.4 19.1

Surround i ng

S.M.D. (um)

11,5 24.3

Front Edge

S.M.D. (um)

17.9 2^.5

Surroundino

Table 2 Results of S.M.D.

(T=723K F=0.1Mpa, t=0.4ms) Front Edoe

Table 2 shows the results of S.M.D. of particles in

different regions of a developing spray. According to the

results, we concluded that even in high temperature the droplet

size distribution of emulsified fuel is more uniform and smaller

than that of pure diesel oil. Statistic results of Fig. 7

manifest that in the middle and surrounding regions of emulsified

oil, 65% of total counted droplets are among 3-lOum, only small

quantity of big droplets have diametres of 40-50um, while in the

same condition for pure oil. only 26% of total droplets have 3-

lOum diametres, most of the particles in that region own

diametres of 10-40um. At the front edge, 80% of droplets occupied

the diametres of 3-25um in the emulsion, while in diesel oil, the

same amount droplets occupied the diametres of 15-65um.

311
(a) Puro Oil Sprav (b) EffluUif ied Oil Sorav

FiQ.6 High Speed Schlieren Cineaatooraphv

Seouences of sorav Develooaeiit

(573k,0.9Hoa;2000f/s)

(a) T=623K t=0.4ffls (b)T=723K t=0.4(ss (c) T=723K t=0.2as

Fragnientations of Emulsified Particles P^''^ Diesel Oil Partcles

Fig. 7 Holooraohic Microoraohes of Sorav Particles

of Both Emulsified and Pui-e Oil

in Hioh Temoerature

312
Conclusively, in the analysis of emulsion vaporizing

characteristics, we find the strength of micro-explosion of

emulsified spray is not as effective as it is in the single

droplet test. It also implies that the vaporizing water addition

causes disturbance inside oil droplets. It's the internal

disturbance of the large size and high velocity droplets that

brings out a fine atomizing quality in high temperature.

Droplet Size Distribution of Pintie Diesei Nozzie

Having tested both the pure and emulsified oil, we found the

pintle diesei nozzle always formed a distinctive spray

distribution. Table 2 and Fig. 8 show this trend. At the outer

region of the spray, the droplets have bigger sizes than those in

inner region. This trend is stronger in low environment

temperature than in the high one. That is because the high

temperature cause a faster vaporization.

60

50

E

a. w40.

«;3 0

E

C3 O

20

10

\ 1

O 0. IHPa 450"^c

A l.OHPa 25<^c

A

0 5 10 15 20 25 3 0

Distance froc injector (od)

Fio.8 Sorav Droolets Variations With

Distance fron Noiile

According to Rayleigh's atomization theory(6,7), when the

oil jets out of the nozzle, small disturbance will appear on the

surface of the jet. Filaments of oil that splash out of the spray

surface will reduce its speed rapidly. These Filaments are

difficult to break again in low speed and gradually become the

large circle particles by the force of surface tension.

CONCLUSIONS

By the methods of high speed schlieren photographic and inline Fraunhofor holographic techniques, we studied the atomizing

characteristics of emulsified spray in a high temperature

combustion bomb and obtained the conclusions as follows:

1) With some necessary technical measures, it is possible

to record the diesei spray fields clearly in high temperature

environment by in-line holography.

2) Studies manifest the emusified diesei spray has a more

uniform size distribution and a samller S.M.D. particle size than

313
the pure diesel oil. Especially in a high temperature condition,

the spray angle expands disruptively and fragmentations take

place in some of big emulsified droplets.

3) The author assume that the better atomizing performance

of emulsion is mainly caused by the following two reasons.

The uncontinuing water-oil internal phase boundary is a

micro-structure easier for droplet disperse.

In high temperature, puffing of the vaporizing water

addition brings internal disturbance to big droplets and make

them easier to disrupt during the high speed atomizing process.

4) The particle size distribution of spray of pintle

diesel nozzle is featured by a tendency. The tendency is that the

deoplets in outer edge has bigger S.M.D. than those in inner

region of the spray.

REFERENCES

1. DRYER F.L. "Water Adition to Practical Combustion Systems--

Concepts and Applications" 16th Syinp^ on Conibus tion .

2. GALLAHALLI S.R. "Combustion of Drops and Sprays of No.

2

Diesel Oil and Its Emulsions with Water" 18th^ Sympj^ on

Combus tion.

3. LASHERAS, C.K.LAW. "Combustion Characteristics of Water-inOil Emulsion Droplets " Combustion and Flame^ VOL. 37(1980).

4. Kenneth Sloan "Measureing photoreceptor MosaiciAn

Application of Adaptive Gray Scale Morphology" SPiE Ygi^937

Applications of Artificial In telligence , 1988

.

5. NISHIDA "Holographic Measurement of Evaporating Diesel

Sprays at High Pressure and Temperature" JSME Vol.30. No.

259 (1987).

6. REITZ R.D. "Atomi zation of a Liquid Jet" PhZiL^ Fluid 30(7)

July. 1987.

7. REITZ R.D. "Effect of Drop Break Up on Fuel Spray" SAE

860469 .

8. TSAO.K.C. "Puffing and Micro-Explosion Phenomenon of Water

Emulsion Fuels" SAE 860304.

314
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Pap^3l

THE THREE-PARAMETER LOG-HYPERBOLIC DISTRIBUTION

AND ITS APPLICATION TO PARTICLE SIZING

T.-H. Xu, F. Durst and C. Tropea

Lehrstuhl fiir Stromungsmechanik

Universitat Erlangen

Nuraberg, Germany

ABSTRACT

This paper investigates the possibilities of representing measured size distributions in the form of a

three-parameter log-hyperbolic (3P-LH) distribution. This distribution is shown to be a special case of the

more general four-parameter log-hyperbolic (4P-LH) distribution but which exhibits a more stable behavior

in the fitting of its parameters to empirical data. The problem of parameter estimation stability using the

4P-LH distribution has been identified and explanations are given for its causes. The 3P-LH distribution

circumvents these problems while yielding an equally good representation of the data for a large variety

of applications. The 3P-LH distribution is therefore concluded to be a valuable means of presenting large

amounts of size distribution data in a more comprehensible manner.

1. INTRODUCTION

Many industrial processes deal with particles in which either the speed of reaction or the quality of

the product depend strongly on the particle size distribution. Spray painting, spray drying, oil combustion

and numerous other unit operations concerned with heat and mass transfer from dispersed phases are wellknown examples. For the purpose of the optimization of these processes, a mathematical description of size

distributions is needed. In the field of spray research especially, such a quantitative description is required

for extracting information from the large amount of measured data in order to study the spatial variation or to compare different nozzles. A further necessity lies in the research aim of predicting spray properties

and their dependence on nozzle configuration, physical properties of the liquid, supply pressure, etc. Although the mean diameters Dio, D20, ^30 and Dz2 are widely used to characterize the essential

features of the distributions, they are not a sufficient mathematical description because different distributions

may have the same values of mean diameters while exhibiting significant diffierences in the distribution shape.

To predict the average velocity and heat/mass transfer rates of a particle system, the use of the equations

for a single particle with simple replacement of the particle diameter by a mean diameter for describing a

particle ensemble may cause considerable error. Furthermore, even if particle dependent transfer phenomena

can be correlated with a certain mean diameter, this correlation may not be valid for other applications.

For this purpose and for a better understanding and modeling of the process concerned with particles, a

probability density function (PDF) must be used.

In recent years, attempts have been made to find a suitable PDF for particle systems. Among the

various forms of distributions, the log-hyperbolic distribution, which was proposed by Barndorff"-Nielsen [1] in 1977, has been shown by Durst and Macagno [2] to be suitable for many fields of particle physics. Bhatia

and Durst [3] summarized the application of diffierent distributions to sprays and carried out a comparative

study of four distributions. Rosin-Rammler, log-normal, a distribution proposed by Li and Tankin[4] and the

315
log-hyperbolic distribution. They have demonstrated clearly that the log-hyperbolic distribution provides a more appropriate description of the measurement results for eight sets of data from three different sources. The main difference among the possible analytic descriptions of the measured distribution lies in the

"tails" of the distribution. As is seen clearly by the data presented in references [1,2,3], the tails of the

measured distribution can be approximated well by straight lines when the logarithm of PDF is plotted

against the particle size or the logarithm of particle size. This is exactly the description afforded by the

hyperbolic/log-hyperbolic distribution. The present authors have examined a number of PDA-measured

droplet size distributions and found that this behavior held not only for water sprays, but also for paint

sprays[5] and coffee sprays[6].

The purpose in approximating a measured size distribution with an analytic function is to characterize a large amount of information, the empirical distribution, with a small number of parameters. Changes

in the distribution should therefore be reflected in changes of the fitted parameters, preferably also with

immediate physical interpretations of the changes. Hence small variations in the observed distribution

should correspond to small variations of the fitted parameters and there should not exist two or more sets of parameter estimates that yield equally good fittings. Unfortunately, this stability problem does exist for

the four parameter log-hyperbolic distribution. In the next section, this is demonstrated by a few examples

and the cause is analyzed. To solve the problem, a new probability density function is proposed, which

can be considered a special case of the log-hyperbolic distribution but which only has three parameters.

Hence, it is called "the three-parameter log-hyperbolic (3P-LH) distribution". This distribution, some of

its characteristics, the method to estimate the parameters, and its relation to the usual four parameter

log-hyperbolic distribution are introduced in section 3. Some applications and comparisons with the usual

log-hyperbolic distribution as well as an example of using this distribution to analyze a spray are provided

in section 4.

2. THE LOG-HYPERBOLIC DISTRIBUTION AND THE STABILITY PROBLEM IN PARAMETER ESTIMATION

The probability density function of hyperbolic distribution proposed by Barndoff-Nielsen [1] may be

written in the form

fix; a, /?, 6, fi) = A{a, p, 8) exp{-ai/«2 -|- (i - p^f -\- P{x - n)} , -oo < x < oo . (1)

where a > 0, |/?| < a, ^ > 0 and fi G (—oo, oo) represent the four parameters to be fitted. A is the

normalization constant introduced to insure that

/+00

f{x)dx = 1 (2)

-00 and is given by

A V^^_ . (3)

The slope of the left asymptote is given by a -|- that of the right by —{a — f3) and fi gives the abscissa of

the asymptotes intersection.

According to Barndoff-Nielsen [1], when x follows Eq.(l) the distribution of s = exp(x) is the loghyperbolic distribution.

In applying the 4P-H distribution to approximate the measured droplet distributions in sprays, several

undesirable characteristics of the parameter estimation have been observed. In some cases the variations

of the fitted parameters for distributions measured in close proximity to one another were very irregular.

Significant variations were even registered for repeat measurements at the same location in the spray. In

some extreme cases, for example in paint sprays, the estimated parameters depended strongly on the initial

values used in the mciximizing procedure and on the number of grouping classes. Despite these instabilities

in the parameter estimation, the fitted curves were largely coincident. The exact nature of these instabilities were therefore studied in more detail using a computer simulation.

A computer simulation has the deciding advantage over experiments that the true distribution is always

known. In the present case a random number generator was used to produce 20,000 samples that were log316
hyperbolic distributed in the range -100 to 200 and with the known distribution parameters as were estimated

in the example 2 of [7]. Applying existing fitting procedures to the simulated data resulted in parameter

estimates which were rather stable and accurate. A further simulation introduced random fluctuations in the

sample elements, corresponding to a measurement error in an experiment, which was normally distributed

and with a standard deviation of 5%. Parameters were again estimated from the resulting sample. Although

the parameter variation rose slightly among different simulations, the estimates remained excellent.

A third simulation was performed using the same known parameters but truncating the left side of

the generated data to yield a distribution as illustrated in Fig. 1. As seen from the five sets of parameter

estimates corresponding to five parallel runs, the variation of the parameters is high, although the fitted curves lie close to one another. This is typical of the behavior found in spray experiments. From these

simulations it is concluded that the tail portion of the distribution is essential to a stable estimation of the

four parameters.

0.030 M n I n T 1

1 1 1 1 1 1 1 m 1 1 n 1 1 1 II I n r n 1 1 1 Ti r [ I n 1 1 1 1 1 f1

1 1 1 1 r i T nTTTTri r rr nTTTTi i M T H i M itirrr n TTTi m:

X

Figure 1: 4P-H fittings for simulated distributions

Initially, an improvement of the estimation stability was attempted by first taking the logarithm of

X prior to grouping and fitting the distribution. This has the effect of elongating the left side tail of the

distribution. This approach was also studied by simulation and Fig.2 illustrates several simulations using

actual parameter values found empirically in a paint spray. In fact, using the logarithm of x tends to shorten

the right side tail of the distribution and the table included in Fig.2 indicates that to some extent the stability

problem still exists.

-2.0

Cl,-4,i

-6.0

-8.0' -1

J 1 1 M M 1 1

1 M M 1 1 1 1 1 1

1 1 M 1 1 1 M 1 1 M M 1 1 M 1 1 II II M II 1

1 1 1 1 II 1 1 1 11 1 1 1 1 1 1 1 L

1

1

1

11

1

1

1

1

1

a 6

1

1

1

11

1

1

- ooooo 1.95 -.768 .727 3.21

I 1.75 -.556 .642 3.07 - AAAAA 1.79 -.649 .597 3.12

: OOOOO 1.81 -.646 .656 3.13

: Mill - — 1.84 -.650 .681 3.13 INI

-| 1 1 1 1 r M 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

1

1 1 1 1 1 1 1 1

1

1 1 1 1 1 1 1 1 M 1 1 1 1 1 1 1 M 1 1 1 1 1 1 M 1 1 r 2 3

X

Figure 2: 4P-H fittings for simulated distributions

317
Thus the length of the available "tails" in the distribution is a deciding parameter in determining the

applicability of the 4P-LH distribution. To investigate the influence of the tail portion in more detail Eq.(l) can be written as

\nf = \nA + i-Qy/PT{^r^)+l3{x-ti) . (4)

From this equation it is seen that In / comprises three terms, In ^4, a standard hyperbola and a straight line,

the slope of which is given by f3. To investigate the stability problem the symmetrical case of the distribution can be considered, i.e. fS = 0, allowing Eq.(4) to be written as

\nf = \nA + {-ay/6^+{x-^y) , (5)

indicating that the slopes of the asymptotes are given by ±a while the peak value of In / depends on In A

and a^.

If the measured distribution has long tails then the square root factor is dominated by (x — /i)^ rather

than 6'^ in tail regions. A possible measure of this effect is the extreme values of {x — nY/S"^, shown in Table 1 for two of the simulation cases. Case A represents the non-truncated case mentioned above which yielded

stable parameter estimates and case B is the distribution shown in Fig. 2. The poorer estimates (Case B)

were Jissociated with extreme values of (i — /d"^ being less than about 20.

Case A

Case B

left side

283

22

right side

853

7.64

Table 1: Comparison of tail effects of two simulation cases

If {x — n)^/6'^ takes large extreme values, thus dominating the square root factor, the parameter a can be uniquely determined by the asymptote slopes. If on the other hand the parameter 6 remains considerably

influential as in Case B, then the distribution can be fitted with arbitrary combinations of a and 6, which

results in a less stable parameter estimation. Note that the freedom in choosing a and 6 combinations is only possible because the norming quantity In A will automatically shift the fitted distribution to a correct

vertical position, fulfilUng Eq.(2).

3. THE THREE-PARAMETER LOG-HYPERBOLIC (3P-LH) DISTRIBUTION

In an effort to avoid the estimation stability problem of the 4P-LH distribution outlined above, the 3PLH distribution is introduced. This is introduced on the assumption that the particle distribution, especially

in sprays, is described by a hyperbola when plotted in logarithmic coordinates. The derived distribution will

then be related to the more well known 4P-LH distribution.

A hyperbola is given in its general form as— - — = 1 . (6)

Since a normalization constant will be introduced later which automatically shifts the distribution to the

correct position in the y direction, the constant b could be arbitrarily set to 1, yielding

Y = -y/TTx^ . (7)

An asymmetry must be introduced to correspond to experimental observation, however the addition of a

linear term, as in the 4P-LH distribution is undesirable as the geometric meaning of the parameters becomes

unclear. In the present case a rotation of the coordinates by an amount given by the angle 0 is introduced

instead:

X = X cos 6 — ysm6 (8a)

Y = xsme + ycos9 . (8b)

318
Combining equations (7) and (8) results in

a rrZ ~ . 2 a\ , 2 (a^ + l)sintf cos^ . . a* cos^ 9 — sin 0 ^ cos^ 6 — sin 6

The abscissa of the vertex of Eq. (9) is given by

(a' + l)sin d cos^

/ a^-(a^ + l)^sSin tf cos f

V a2 cos2 - sin^tf

Introducing a location parameter /i, which shifts the curve horizontally, the final probability density function

takes the form of

/(,) ^Ae^^-

^, ;_ ^ V(a» COS' > - .i„' + (. .0 - ,Y - ^'^^^^^^ fj/ (x + . (I.)

This distribution is pictured in Fig. 3 in which the physical meaning of the 3 parameters is apparent. The

parameter a > 0 determines the slope of the asymptotes in the unrotated coordinate system. The parameter

1^1 < min{tan~^ a, (sin~^ 2a/(a^ + l))/2} is the rotation angle. The parameter // is the position of the vertex

(mode), yl is a normahzation constant, as given before by Eq.(2).

Figure 3: The three-parameter hyperbolic distribution

The distribution f(x) given by Eq.(ll) is termed "the three-parameter hyperbolic (3P-H) distribution".

When X follows this distribution, the distribution of s = exp(x) is "the three-parameter log-hyperbolic (3PLH) distribution".

This distribution can be related to the four parameter distribution using the following relations:

a =

S

a2 cos2 e - sin^ 6

(g' + 1) sin (9 cos g ~ a2 cos2 9 - sin^ 6

(a'cos^tf-sin'tf)'/'

(12a)

(12b)

(12c)

(12d)

319
The norming constant A of the three-parameter hyperbolic distribution can also be found by combining

Eqs.(12) and (3):

^ ^

x/a'-(a' + l)'sin^g co82g

V cos^ 9 — sin

In addition, all the statistical quantities such as moments and mean diameters can be obtained immediately

by using Eqs.(12) and the available formulas for the hyperbolic distribution in [8]. Several methods of parameter estimation for the 3P-LH distribution have been investigated. The

meiximum likeness estimation, as proposed in [1], has been found to be superior to the maximum likelihood

approach and to the method of least squares. This conclusion is based on comparisons of the computed

mean diameters and moments and held also for the agreement of the estimated parameters with the true

values in fitting the simulated data, especially in the cases in which the tails of the measured distribution are incomplete.

Since the three parameters are related simply to the shape of the distribution, they can be roughly

estimated directly by referring to Fig. 4 and using the relations

. ,\. -\ Xm — t\ 1 _i 12 — s l-tA \ a = tan (-tan h - tan ) (14a)

8 =z -(tan tan ) (14b)

H = Xm (14c)

These relations are useful for obtaining initial values for the minimizing procedure in the maximum likeness

estimation.

Figure 4: Rough estimation of the parameters

The parameter estimation of the 3P-LH distribution is found to be well behaved and stable, even for

samples with very short distribution tails. This is illustrated in more detail in the following section.

4. APPLICATIONS AND EVALUATION

The stability of the parameter estimation for the 3P-LH distribution is evaluated by repeating the

simulation presented in Fig.l, using a truncated input sample. Fig. 5 illustrates that the three estimated

parameters, expressed in this figure as the conventional four parameters according to Eqs.(12), vary little

for different simulations, reflecting the similarity of the different fitted curves. The improvement over the

4P-H distribution is clearly evident.

320
0.030 u I ri ri 1 1 T

) 1 1 1 1 1 1 1 M 1 1 1 T r rT) I f 1 1 1 T 1 1 1 II

1

1 1 1» tttt irym rrrTji ri m i r i r 1 1 n 1 1 1 1 mj 1 1 ii 1 1 1 t i

f

i

Figure 5: 3P-H fittings for simulated distributions

Durst and Macagno[2] have applied the log-hyperbolic distribution to seven different cases of particle

distributions, namely the mass-size distribution of a sand sample, the size distribution of diamonds, the

size distribution of river bed sediments, the size distribution of glacio-fluvial sediments, the measurement

of aerosol distributions, the particle size distribution in a spray and a molecular weight distribution. The

results show that the log-hyperbolic distribution successfully represents the particle size distributions of all of these particle systems. In the present study, the three-parameter log-hyperbolic distribution was also

applied to these seven cases using the same input data.

Fig. 6 illustrates a direct comparison between the 3P-LH and 4P-LH fitted distributions and indicates

that the former provides an equally good representation in all cases except perhaps for the two sediment

examples.

Further examples of the application of the 3P-LH distribution are presented using spray data. These

data are taken from a low pressure water spray, having also been analyzed by Bhatia and Durst and cor- responding to Example 2 in [7] and data sets 4,5 and 6 in [3] respectively. The results of the fitting are

illustrated in Fig. 7. From Fig. 7 it is seen that, for the most parts, the three-parameter curve is almost

indistinguishable from the four-parameter curve. A more detailed comparison is possible by examining the computed mean diameters and the moments

presented in Table 2. Both fitted curves result in almost equally excellent estimates of the mean diameters

and moments, therefore the 3P-H distribution can be considered equivalent to the 4P-H for these applications.

Despite these similarities, it is interesting to note from Table 3 that the close agreement between the two

fitted curves is achieved with significantly different parameter values. In Table 3 Eqs.(12) has been used

to compute the four parameters corresponding to the three parameter fit. Accordingly, the fit results in a

different set of hyperbolic shape triangle parameter x and ^ (see Ref.[8]).

A final application example illustrates how the spatial variation of the three estimated parameters in a spray can be used to interpret the physical changes in the size distribution. Fig. 8 shows the variations of

the three parameters and the mean diameters along the axis of a water spray, the data having been collected

using a phase Doppler anemometer.

In the region z — Smm to 14mm parameter a increases slightly, parameter 0 decreases and fi remains

constant. The rotation indicates a broader distribution especially towards larger droplets, although there is no shift in the mode (vertex). In the region z = 14mm to 16mm there is a steep descent in 9 and fi, while a decreases only very slowly. This indicates a sudden increase in the percentage of small droplets and a rapid

decrease of large droplets, indicating that a violent break-up takes place in this region. Beyond z = 26mm,

9 and /z begin to increase while a continues to decrease slowly. Here the percentage of smaller droplets is decreasing and the distributions are becoming narrower. The interpretation given above can be verified by

examining the behavior of the various mean diameters along the spray center line, also given in Fig. 8.

321
;l 1 1 1 1 M 1 1 1 1 1 1 1 1 1 1 1 1 1 1 M 1 1 1 n 1 1 1 1 1 n 1

1 II 1 1 1 1 1

r Oq^

llllllllll

i o distribution of molecular weight

=1 1 1 1 1 1 1 1 1 1 1 1 1 M 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1- 9 10 11 )2 13 )4

ooooo observed

/?-paramfitRr log--hyperbolic _ _ 4—parameter log--hyperbolic

»

Figure 6: Comparison of 3P-LH distribution with 4P-LH distribution in seven applications

322
- I I [ I I I I I I

I I I I 1 I I I I I I

I I I I I I I I I

I

I M I I I I M I

I M I I I I I I j -3 aXCO observed

fitted 3P-H

- — filled *P-H

Data set 5 [3]

I t I I I I I I 1 I I I I I I I t I I t I 1 I - 20 40 60 80 particle diameter (/itm)

100

I I I I I I I I I

I

I I I I I M I r

I

r I I I I I r I I I

I I I I I I I I I I

I I I I I I I I I .

OOCCX) obBerved

fitted 3P-H

- — fitted 4P-H -.

Data set 6 [3]

•' ' ' I I 1 I 1 I I I I I I I ) I I I t 1 I I r 20 40 60 80 particle diameter (/^m)

Figure 7: Application of the three-parameter hyperbolic distribution to spray

Dio D20 D30 D32 1712 ^3 me

Ex. 2

of [7]

Summ.

3P-H

4P-H

26.93 33.70 39.50 54.25

26.92 33.67 39.45 54.17

26.92 33.72 39.56 54.46

411.1 8904 621562 2.62E7 1.69E9

409.0 8870 624360 2.90E7 1.73E9

412.6 9105 639223 2.99E7 1.78E9

Data

No. 4

of [3]

Summ.

3P-H

4P-H

25.61 29.79 33.36 41.83

25.60 29.79 33.39 41.94

25.60 29.79 33.37 41.87

231.8 2535 1.78E5 5.03E6 2.50E8

232.5 2603 1.83E5 5.25E6 2.61E8

231.9 2568 1.81E5 5.15E6 2.56E8

Data

No. 5

of [3]

Summ.

3P-H

4P-H

23.51 27.96 32.09 42.28

23.50 27.83 31.82 41.60

23.50 27.91 31.99 42.03

229.4 3895 2.13E5 7.67E6 3.58E8

222.1 3577 1.99E5 7.12E6 3.36E8

226.6 3785 2.10E5 7.62E6 3.61E8

Data

No. 6

of [3]

Summ.

3P-H

4P-H

31.90 39.27 45.36 60.52

31.90 38.93 44.80 59.33

32.22 39.57 45.72 61.06

524.5 1.07E4 8.03E5 3.39E7 1.99E9

497.8 9.81E3 7.53E5 3.16E7 1.88E9

509.2 1.03E4 7.84E5 3.34E8 1.98E9

(Summ. represents summation over grouped data)

Table 2: Comparison of the mean diameters and the moments

fitted with 4P-H corresponding to 3P-H

a .06454 0.1119

.02184 0.0672

6 6.5187 8.9568

9.6533 4.5379

X .286338 0.4472

.846385 0.7450

Table 3: Comparison of the parameters for Example 2 from [7]

323
z{nim) z(mm)

Figure 8: Variation of size distribution along the axis of a spray

5. CONCLUSIONS

In this paper the stability problem in estimating parameters of the four-parameter log-hyperbolic distribution was examined by computer simulation and the importance of the sufficiently long tails was identified.

The three-parameter log-hyperbolic distribution was then introduced and shown to be a special case of the

4P-LH distribution. Further simulations indicated that this new distribution did not suffer from the estimation stability problem and gave an equally good representation of measured data for a wide variety of

size distribution applications. Detailed examples were given for spray data and for this application the three

parameter log-hyperbolic distribution can be recommended for future use. The financial support of the Deutsche Forschungsgesellschaft through grant Du. 101/14 is gratefully

acknowledged, as is a study scholarship from PR China for the first author.

References

Barndorff-Nielsen.O., "Exponentially Decreasing Distributions for the Logarithm of Particle Size", Proc.

R. Soc. Lond. A.353, pp. 401-419 (1977).

Durst, F. and Macagno,M., "Experimental Particle Size Distributions and Their Representation by LogHyperbolic Functions", Power Tech., 45, pp. 223-244 (1986).

Bhatia,J.C. and Durst,F., "Comparative Study of Some Probability Distribution Applied to Liquid

Sprays", Part. Part Syst. Charact., 6, pp. 151-162 (1989).

Li,X. and Tankin,R.S., "Droplet Size Distribution: A Derivation of a Nukiyama- Tanasawa Type Distribution Function", Combust. Sci. Technol., 56, pp. 65-76 (1987).

Domnick,J, Dorfner,V, Durst, F, Tropea,C. and Xu,T.-H., "Bestimmung der Zerstaubungseigenschaften von Lacken: Voruntersuchung" , Report No. LSTM 282/1/90, Univ. Erlangen-Niirnberg (1990).

Domnick,J, Dorfner,V, Tropea,C. and Xu,T.-H., "Phasen-Doppler-Messungen an einem Spriihtrockner

der Firma LINOR", Report No. LSTM 278/1/90, Univ. Erlangen-Niirnberg (1990).

Bhatia,J.C. and Durst,F., "LHPDF - A PC Package for Estimating Parameters of the Log-Hyperbolic

Distribution, Moments and Mean Diameters", Report No. LSTM 230/T/88, Univ. Erlangen-Niirnberg

(1988).

Bhatia,J.C., Domnick,J, Durst,F. and Tropea,C., "Phase-Doppler-Anemometry and the Log-Hyperbolic

Distribution Applied to Liquid Sprays", Part. Part Syst. Charact. 5, pp. 153-164 (1988).

324
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper32

A DROP SIZE DISTRffiUTION DATA BASE

FOR FOREST AND AGRICULTURAL SPRAYING:

POTENTIAL FOR EXTENDED APPLICATION

M.E. Teske', P.J. Skyler^ and J.W. Barry^

*Continuum Dynamics, Inc., Princeton, NJ, U.S.A.

''USDA Forest Service, Forest Pest Management, Davis, CA, U.S.A.

ABSTRACT

Over the last several years the USDA Forest Service, and other agencies

and companies, have contracted for wind tunnel tests to determine drop size

distributions of pesticides and simulant spray material atomized by nozzles.

These test results have been assembled into a data base presently containing

224 atomization entries. Nearly all spray materials and nozzle types used by

the Forest Service are represented. Access to the data base is through an

interface program DROPSIZE, developed specifically for the data base and

freely available from the Forest Service. Preliminary examination of some of

the data indicates that spray drop size distributions may be correlated based

upon physical properties, but that unique details in the distributions dictate

the continued use of wind tunnel tests.

INTRODUCTION

The drop size distribution of spray material atomized by nozzles

influences the magnitude of evaporation, spray deposition and drift. Several

factors affecting the atomization process and the resulting spray distribution

include nozzle type, hydraulic line pressure, flow rate, shear across the

nozzle orifice and specific properties of the formulation. The details of the

resulting distribution are critical, especially to forest and agricultural

applications, where specific levels of spray material must be deposited to

achieve success and avoid environmental insult. Concerns over aerial spraying

have escalated in the past several years, as has the cost of pesticides. It

is becoming increasingly clear that the physical processes generating the

spray distribution must be better understood. Risk of potential damage from

driftable material is becoming too high.

Under the direction of the USDA Forest Service, wind tunnel tests of

pesticides and simulant tank mixes have been conducted to:

1. Determine the atomization of tank mixes as influenced by hydraulic

line pressure, flow rate, air velocity and shear across the atomizer,

components including adjuvants (chemical, physical and biological) in the tank

mix, viscosity, specific gravity, surface tension and other atmospheric

conditions

.

2. Evaluate the mixing and handling of the tank mix.

3. Develop recommendations for field use of the tank mix.

Atomization results from these tests are then used in selecting the

optimum combination of factors that have a high probability of meeting field

project objectives. Wind tunnel facilities for testing pesticide sprays have

been established and are maintained at University of California. Davis.

California (the pioneering work of Wes Yates and Norm Akesson) ; Cranf ield

325
Institute of Technology, England: University of New Brunswick, Canada; USDA

Agricultural Research Service, College Station, Texas: and New Mexico State

University, Las Cruces, New Mexico.

Recently, the entire complement of spray distributions generated for the

Forest Service (224 entries) has been compiled [1] and assembled into an

interactive data base [2] . This paper summarizes the contents of this data

base, reviews the historical approaches suggested for collapsing spray

distribution data, and implements three such approaches as examples of the

usefulness of the data base.

DATA COLLECTION

Numerous researchers have investigated spray distributions that are

essentially water-based [3,4], oil-based [5] and generated from standard tank

mixes [6] . Many other studies have been conducted, but these results are

unpublished or held company proprietary. Generic results from these wind

tunnel tests indicate the following:

1. Shear across hydraulic nozzles caused by nozzle orientation, and

shear across the rotating basket cages of rotary nozzles caused by rotation,

are major factors in break up that produce small drops.

2. Increase in surface tension increases drop size.

3. Viscosity is a minor influence on atomization.

4. Rotary atomizers can become overloaded, resulting in larger drop size

atomization even when the rotation rate is held constant.

5. Most nozzles produce a large number of small drops (less than 56

micrometers in diameter): however, these drop sizes generally represent less

than one percent of the total volume in the spray.

6. Slight changes in some chemical, physical and biological properties

of a tank mix can significantly alter the atomization.

It is this last observation that guarantees the continued use of wind

tunnel tests. Although the development of atomization models (invoking

similarity or dimensional analysis) is currently underway, these models are

envisioned to enhance rather than replace the need to characterize tank mixes

and spray systems. An examination of spray distributions in the above

references indicates a strong dependence upon variables that may not be easily

quantified.

Spray drop size distributions are typically developed by using laser

probes to scan the spray downstream of the nozzle. A flow- through wind tunnel

provides an ambient air stream at the anticipated flight speed of the

aircraft. The typical output from a single nozzle study at a single air

stream speed, and at other specified conditions depending on the nozzle type,

is shown in Figure 1. Here the volume fraction is shown, although the same

experiment will recover the number density as well. The typical gross

features of interest are:

1. Dq -]_: The diameter below which 10 percent of the spray material is

found.

2. Sauter Mean Diameter SMD: the diameter whose ratio of volume to

surface area is the same as that of the entire spray.

3. Volume Median Diameter VMD : the diameter below which 50 percent of

the spray material is found.

4. Dq gi The diameter below which 90 percent of the spray material is

found.

326
The ctimulative volume fraction CVF is generally shown, since it tends to

smooth the distribution, de-emphasize details of the peak, and offers a result

that goes between a value of zero and unity. Further operations on CVF offer

the prospects of representing its behavior mathematically. To this point

previous researchers have dealt with limited data. Here, the extensive data

base generated for the Forest Service offers a new look at some old

suggestions for collapsing data.

DATA BASE CONTENTS

Collected from published and unpublished reports, a spray drop size

distribution data base has been compiled [l] . This data base covers 40 spray

materials and 27 nozzle types in 224 unique distributions. Most applications

for these distributions include a variation in air stream speed, angle of the

nozzle to the air stream (where applicable) or hydraulic line pressure or

atomizer rotation rate. The data base is accessed with an interactive program

(called DROPSIZE) that enables the user to extend the data base under linear

interpolation (and extrapolation) between known results for flow rate, air

stream speed, atomizer rotation rate and nozzle exit angle, where applicable

[2]. The code and data base are free from the Forest Service. Table 1

summarizes the spray materials, while Table 2 summarizes the nozzle types

available in the data base.

DATA CORRELATION

Trying to find a way to correlate spray drop size distributions, where

numerous variables can influence the results, has attracted attention to areas

beyond aerial spraying, and has generated niunerous insights into the problem.

Goering and Smith [7] summarize the various techniques that have been

attempted to essentially curve fit CVF. They ended up suggesting a threeparameter technique (upper limit logarithmic normal [8]) that extends the

standard logarithmic normal approach [9] . The most redeeming feature of the

logarithmic normal representation is the expectation of a straight- line fit

through the data plotted on these particular axes.

Lefebvre [10] offers reasons for finding a suitable mathematical

representation for the CVF distribution:

1. Provide a satisfactory fit to the drop size data.

2. Allow extrapolation outside the range of measured values.

3. Permit easy calculation of mean and representative drop diameters and

other parameters of interest.

4. Provide a means of consolidating large amounts of data.

5. Furnish insight into the basic mechanisms involved in atomization.

The existence of the present data base should continue a process that

began over 55 years ago. What is driving the problem now is the requirement

by the U. S. Environmental Protection Agency for drift data to support

registration and reregist ration of certain pesticides. To provide drift data

the driftable part of the atomization must be known; however, wind tunnel

tests are costly. Therefore, it becomes prudent to seek approaches that might

reduce the need for extensive wind tunnel testing.

DATA BASE EXAMINATION

Preliminary correlation of the data in the data base includes three

examples. These examples correlate the data following the techniques

327
suggested previously, and demonstrate how to extend the data base to

conditions not contained in the data base.

1. STRAIGHTFORWARD INTERPRETATION OF LOGARITHMIC NORMAL DISTRIBUTIONS

An examination was made for water released from flat fan nozzles at 90

degrees (normal) to a 100 mph air stream. Four flat fan nozzles (8001. 8004,

8010 and 8020) are contained in the data base with these characteristics.

Their logarithmic normal distributions are shown in Figure 2.

The data is correlated by straight lines on these axes, assuming:

y = a + b X (1)

where x = log ( D ): y = log ( - log ( 1 - CVF ) ) and D = drop diameter (in

micrometers). A least squares fit for all four spray distributions gives;

a = -14.164

b = 2.293 FR'^-^^^-^

where FR is the flow rate through the nozzle (for the 8001 nozzle the flow

rate is 0.01 gal/min) . This correlation collapses the four spray

distributions, and, when compared with the original data, shows a standard

deviation in volume fraction of 0.024.

2. ROSIN- RAMMLER LOGARITHMIC NORMAL DISTRIBUTIONS

An examination was made for water released from a Micronair AU5000.

Seven distributions are contained in the data base: 3750 and 4100 RPMs at 50

mph air stream: 4200, 8000 and 9100 RPMs at 100 mph: and 10850 and 11700 RPMs

at 130 mph. The technique in Example 1 collapsed the data by correlating the

slope of all of the logarithmic normal curves: here, the drop diameters are

first divided by their respective diameters Dq 532* "^^^ diameter below which

63.2 percent of the spray material is found, (as suggested in [10]) before

transformation. The result is shown in Figure 3.

The data is seen to nearly collapse to a straight line, discounting the

leveling off at larger drop diameters. This line may be determined by

Equation 1 where now x = log ( D / D632 ) and y = log (

- log ( 1 - CVF ) )

.

The least squares fit for all seven spray distributions gives:

a = 0.211

b = 2.206

With this normalization, however, the sensitivity of the correlation is moved

to the normalizing diameter Dq 532* Comparison with the original volume

fraction data for all seven spray distributions yields a standard deviation of

0.090, a significant variation, due principally to the curve fit in Dq g32

(which has a standard deviation of 33.9 micrometers). With these simple fits,

significant errors can result in spray drop size distribution, as noted in [6]

for VMD correlations.

3. MUGELE- EVANS ERROR FUNCTION NORMAL DISTRIBUTIONS

An examination was made for the biological pesticide Foray 48B (Bacillus

thurlnglensis) combined with water in several dilutions, for an 8004 flat fan

nozzle at 90 degrees (normal) to a 100 mph air stream. Foray 48B is contained

328
in the data base in five dilution factors (0 percent, or undiluted; 5, 10. 25

and 50 percent water added by volume). In this example, the error function

technique [8] , strongly endorsed by [7] , is used to correlate the spray

distributions

.

Goering and Smith [7] give approximate formulae for generating the three

parameters needed to represent each spray distribution:

n = VMD ( Dq.I + Do. 9 ) - 2 Dp.i Dp.

9

^m ; VMD - Do.l Do. 9 / VMD

^ = Dm - VMD

VMD

5 = 0.394

logic ( ) Dm - Do. 9 (2)

to give

CVF = 0.5 ( 1 + erf ( 5z ) )

z = log ( A D ) Dm - D (3)

and erf is the commonly defined error function. Thus, this approach involves

a best fit to an error function for each spray distribution. For this example

the data was fit with a single curve using the average values for these three

parameters, not including the 50 percent data. The transformed distributions

are shown in Figure 4. It may be seen that the error function fit is quite

good at the lower drop diameters, and becomes progressively worse for larger

drop diameters.

Considerably more work with the data base is needed before deciding on a recommended curve fit technique. The three examples, given above, suggest

that some data collapse is possible. Sufficient wind tunnel tests will still

be required to define spray distribution limits.

CONCLUSIONS

A preliminary examination of an extended USDA Forest Service spray drop

size distribution data base indicates that data correlation and collapse is

possible. The sensitivity of the curve fits to slight changes in the volume

fraction determines the accuracy of the substituted expressions. Compression

of the spray distribution data may result in some inaccuracies; thus, wind

tunnel tests will continue to be important. Correlation techniques should

however be able to reduce the number of wind tunnel tests required.

REFERENCES

1. Skyler, P. J. and Barry, J. W. , "Compendium of Drop Size Spectra

Compiled from Wind Tunnel Tests,** USDA Forest Service Forest Pest

Management Report No. FPM-90-9 (1990).

2. Teske, M. E. , "DROFSIZE User Manual," Continuum Dynamics. Inc. Technical

Note No. 90-10 (1990)

.

329
3. Yates. V. E. , Akesson. N. B. and Cowden, R. E. , "Measurement of Drop

Size Frequency from Nozzles Used for Aerial Applications of Pesticides

in Forests," USDA Forest Service Equipment Development Center Report No.

8434-2803 (1984).

4. Bouse, L. F. , Carlton. J. B. and Jank, P. C. , "Effect of Water Soluble

Polymers on Spray Droplet Size," Transactions of the ASAE, Vol. 31. No.

6. pp. 1633-1641, 1648 (1988).

5. Bouse, L. F. and Carlton, J. B., "Factors Affecting Size Distribution of

Vegetable Oil Spray Droplets," Transactions of the ASAE, Vol. 28. No. 4.

pp. 1068-1073 (1985).

6. Picot, J. J. C, van Vliet. M. W. and Payne, N. J., "Droplet Size

Characteristics for Insecticide and Herbicide Spray Atomizers." Canadian

Journal of Chemical Engineering. Vol. 67. pp. 752-761 (1989).

7. Goering. C. E. and Smith, D. B., "Equations for Droplet Size

Distributions in Sprays," Transactions of the ASAE, Vol. 21, No. 2, pp.

209-216 (1978).

8. Mugele, R. and Evans, H. D.. "Droplet Size Distributions in Sprays,"

Industrial and Engineering Chemistry. Vol. 43, No. 6, pp. 1317-1324

(1951).

9. Rosin, P. and Rammler, E., "The Laws Governing the Fineness of Powdered

Coal," Journal of the Institute of Fuel, Vol. 7. No. 31, pp. 29-36

(1933)

.

10. Lefebvre. A. H. . Atomization and Spravs . Hemisphere. New York. pp. 421.

See pgs. 79-103 (1989)

.

Table 1 Spray materials in the USDA Forest Service data base,

Dipel 6L

Dipel 6L with Triton

Dipel 8L

Esteron

Foray 48B

Garlon

Glycerine

Roundup

San 415 SC 32LV

Thuricide 32LV. 48LV. 64LV

Water

Water with Aatrex

Water with Esteron

Water with Manganese Sulfate

Water with Manganese Sulfate and Nalco Trol

Water with Nalco Trol

Water with Velpar

Table 2 Nozzle types in the USDA Forest Service data base.

Flat fan: 8001. 8002. 8003. 8004. 8006, 8010, 8020

Hollow cone: D2-23, D2-25, D2-45. D3-45, D4-45, D4-46, D8-45. D8-46,

DlO-45

Jet: D8. DIO

Raindrop: RD-7. RD-10

Rotary: Airbi. Beecomist. Micronair AU5000. Micronair AU7000,

Spinner, Unimizer

330
0.15

g 0.10

0.05

>

0.00

200 400 600

Drop Diameter (micrometers)

800

Figure 1 Spray drop size distribution volume fraction for a D4-45 hollow cone

nozzle, spraying water, oriented 90 degrees (normal) to a 50 mph air

stream. Pertinent drop diameters: Dq = 131 micrometers; SMD = 223

micrometers: VMD = 263 micrometers: Dq g = 398 micrometers.

too O

toJQ 3

4 5 6 7

Log ( Drop Diameter )

Figure 2 Logarithmic normal distributions for water released from flat fan

nozzles at 90 degrees (normal) to a 100 mph air stream. 8001: open

circles: 8004: closed circles: 8010: open rectangles; 8020: open

triangles

.

331
OX)

o

-1 0 1 Log ( Drop Diameter / D632 )

Figure 3 Rosin- Rammler logarithmic normal distributions for water released

from seven Micronairs rotating at different RPMs. Data are open

circles. Line is least squares fit to all of the data.

-4 -2 0 2 4 6

Log ( Transformed Diameter)

Figure 4 Mugele -Evans transformed distributions for Foray 48B combined with

water in five dilutions, and released from an 8004 flat fan nozzle at

90 degrees (normal) to a 100 mph air stream. Open circles are 0. 5,

10 and 25 percent dilution data: closed circles are 50 percent

dilution data. Error function curve fit to open circle data is line.

332
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper33

EFFECTS OF INITIAL DROPLET SIZE DISTRIBUTIONS ON

POLYDISPERSE SPRAY JET FAR-FIELD DIFFUSION FLAMES

Y. Tambour, I. Silverman and J.B. Greenberg

Faculty of Aerospace Engineering

Technion - Israel Institute of Technology

Haifa, Israel

ABSTRACT

A theoretical study of polydisperse spray jet far-field diffusion flames is presented. In order to analyze the influence of initial droplet size distribution on the flame, a comparison of four kinds of

initial droplet size distributions is carried out. These are: (A) a spray of "small" droplets of

diameters initially within the range of 20-25.2 p,m; (B) a "bimodal" size distribution with spray

droplets initially in two size groups: "small" 20-25.2 ^m, and "large" 80-100.8 |J.m; (C) all droplets of

the spray are initially of an "average" size of 40-50.4 jxm; (D) all droplets are initially "large" and

are within the limits of the size group of 80-100.8 )xm. All other parameters such as: the flow field,

the total amount of injected fuel, the evaporation model and the chemical reaction model are kept the

same for each of the four cases. The theoretical analysis includes solutions for the spatial evolution

in local droplet size distributions which determine the amounts of fuel vapors which are locally

released via the evaporation process. Thus, solutions of the present study encompass the mutual

relations between local droplet size distributions, local fuel vapor concentrations, and the spatial

temperature-field distribution resulting from the combustion process. The spatial variations of these

physical properties are explained in the paper, emphasizing the differences between the four kinds of

initial droplet size distributions. Significant differences are found in the concentration of droplets

and fuel vapors in the far-field and in relatively cold areas in the jet's outer envelope. These

differences strongly depend on the initial size distribution and may also be important in future

studies of emission of pollutants. Two other important conclusions of the present study are: first,

that a spray of an initial "bimodal" size distribution cannot be replaced by and initial "average" size

distribution, and second, that the presence of "small" droplets in the initial size distribution

significantly affects the behavior of the "large" droplets in the spray flame.

INTRODUCTION

There is a series of questions that spray flame researchers often raise. For instance, can a polydisperse spray be represented by a group of droplets of an "average" size in the theoretical

analysis of spray flames? How does a spray flame resulting from a "bimodal" initial size distribution

of droplets initially in two size groups: "small" and "large" (and nothing in between) differ from a

spray flame of droplets of only an "average" size? What are the changes that can be expected in a flame resulting form a spray of only "large" droplets, if "small" droplets are added to the initial

size distribution? In the light of the latter question, what should be the accuracy required by

modellers in measurements at upstream stations (which may serve as initial conditions for computations)

in terms of the boundaries of the spectrum of droplet sizes in pointwise spray histograms?

These are a few of the important questions that spray flame modelers face nowadays, especially in

333
light of the rapid development in measurement techniques [1-12] which supply plenty of experimental

data on pointwise droplet size distributions of polydisperse sprays in combusting systems [13,14].

In recent years, comprehensive experimental data on downstream changes in spray characteristics of

fuel droplets produced by a twin-fluid (air-kerosene) jet atomizer have been reported by a group of

researchers: Yule, Ah Seng, Ereaut, Felton, Ungut and Chigier [3,15-19]. TTiese experimental data

indicate the existence of 3 zones in the twin fluid atomizer flow field:

(i) The "atomization zone", (up to some 50 nozzle diameters downstream [19]). Within this zone

the break up of the liquid fuel jet into droplets occurs.

(ii) The "near-field" is defined here as the zone that starts immediately after full atomization

has been accomplished. The "near-field" is characterized by large differences in momentum

between the two phases (the droplets and the surrounding atomizing air).

(iii) The "far-field" is defined here as the downstream zone that follows the "near-field" where

experimental data essentially show that most of the droplets travel at a velocity which is close to the host fluid velocity field.

On the basis of these experimental data, theoretical simulations of the far-field downstream

changes in the integral droplet size distributions and in pointwise spray histograms, have been

presented [20, 21], for "cold" and "heated" sprays (i.e. sprays that are introduced to a preheated

environment so that vaporization rates are increased, but with combustion occurring).

As to combusting spray jets, effects oiF far-field coalescence effects on flame height, on maximum

flame temperatures, and on the flux percentage of droplets which may "survive" the flame were analyzed

by Greenberg and Tambour [22]. They also analyzed [23] the effects of total fuel mass fractions on

flame behavior by comparing various flames of the same initial polydisperse size distribution, but of

different fuel consumptions.

These two studies serve as a basis for the present study regarding the solution procedure for the

flow field equations and the spray model (although coalescence of droplets will not be considered

here). Since the purpose of the present study is to shed light on the issues raised in the beginning

of the introduction, we focus here on the comparison between four initial droplet size distributions.

These are: (A) "small" droplets (20-25.2 ^m); (B) "bimodal" distributions: "small" and "large" (20-25.2

^m and 80-100.8 ^m); (C) "average" (40-50.4 \im) and (D) "large" (80-100.8 \im).

The evaporation model and the chemical reaction model are kept the same for each of the four

cases. Also kept the same are the flow field parameters and the total amoimts of injected fuel (a mass

fraction of 20%).

As previously indicated, the present study is limited to the far-field flow. A more advanced model

for jet spray flames which includes the near-field is present being developed by the authors and will

be reported.

To analyze the effects of initial droplet size distributions on spray jet diffusion flames, one

has to rely on pointwise polydisperse spray histograms at a given up-stream station, which serve as

initial conditions for the computations. Thanks to advanced measurement techniques, such pointwise

spray histograms have been reported by experimentalists [12-14]. They are usually of a "bar-graph"

shape, where each "bar" corresponds to the concentration of droplets within a defined range of droplet

diameters, i.e. each "bar" corresponds to a size group of droplets, or a size section.

334
Thus, in a similar way, in the present study the local pointwise, polydisperse size distributions are divided into ten size-groups (sections) according to the following droplet diameters: 0-20;

20-25.2; 25.2-31.8; 31.8-40; 40-50.4; 50.4-63.5; 63.5-80; 80-100.8; 100.8-127; 127-160 ^m.

Conservation equations for each section are used to represent the spray. Derivation of these equations

has been described in detail in previous publications [20-26]. Relevant data is also given in Ref.

[23].

RESULTS AND DISCUSSION

In order to analyze the influence of initial droplet size distribution on the flame, a comparison

between four kinds of initial droplet size distributions is carried out. These are: (A) a spray of

"small" droplets of diameters initially within the range of 20-25.2 jxm. (B) a "bimodal" size

distribution with spray droplets initially in two size groups: "small" 20-25.2 jim, and "large" 80-100.8

|im. (C) all droplets of the spray are initially of an "average" size of 40-50.4 ^im. (D) all droplets

are initially "large" and are within the limits of the size group of 80-100.8 [im. For each of the

above cases, the initial mass fraction of the spray is 15%, whereas in the "bimodal" distribution,

these 15% are equally divided between the small droplets and the large ones.

In Figs. la,b, variation of integral fluxes of the liquid phase (i.e. droplets) and the

corresponding fluxes of fiiel vapors along the axial direction is shown. The integral values at each

downstream axial station were obtained by integration of the relevant fluxes over all radial distances.

The spatial contours of spray mass fractions and spray pointwise histograms are presented in Figs.

2,a,b,c,d, whereas the corresponding contoiu-s of the fuel vapors are shown in Fig. 3.

At the initial upstream station, where the spray is introduced, we assume that the flow already

contains a certain amount of fiiel vapors (5% for all cases "A" to "D"). Thus, the fuel vapor fluxes in

Fig. la all start from the same value. For all cases the initial fuel vapor flux decreases as vapors

are consumed by the flame. However, immediately afterwards a significant increase in fuel vapor flux

is noticed in case "A" (when only small droplets are present) due to the rapid evaporation of small

droplets. This is consistent with the corresponding sharp decrease of the liquid fuel flux for this case in Fig. lb. Eventually, all fuel vapors are consumed by the flame further downstream.

The "average" size spray, case "C", exhibits a similar pattern, although the evaporation rate is smaller. In case "D", of only "large" droplets, the evaporation rate is even smaller, so that vapors

are consumed by the flame before having a chance to build up a high local concentration.

It is interesting to compare the behavior of the "bimodal" size distribution, case "B", with all other cases. First, there is a build up of fiiel vapors due to the evaporation of the small droplets,

(similar to what happens in case "A"), and then, the large droplets manage to survive deeper into the

flame and are noticeable further downstream (see Fig. lb).

The influence of "large" droplets (80-100.8 ]im) on the flame in case "B" is different from that of

case "D". When all the injected droplets are "large", case "D", droplets and fuel vapors are found

further downstream and at larger radial distances at the edges of the jet (see Figs. 1, 2d and 3d) and

in relatively "cold" areas (see Fig. 4). In case "B", "small" droplets (20-25.2 ^m) are present in the

initial size distribution beside the "large" ones and significantly affect the behavior of their

"large" neighbors. The small droplets evaporate faster and their fuel vapors feed the flame. As a

result, high temperatures develop at shorter downstream distances ^(compare cases "B" and "D" in Fig. 4,

note that the "D" flame does not reach the temperature of 2100 K) and thus enhance the evaporation of

the "large" droplets. These effects are most pronounced at downstream stations x=0.1 m and x=0.2 m, as can be seen in Figs. 2b and 2d by comparing the corresponding spray pointwise histograms of case "B"

335
with those of case "D". (Note that the change in scales as one marches downstream. As the spray

-1 -2. -2 -3 -4

evaporates, scales change from 10 (or 10 ) to 10 and further downstream to 10 , (or 10 )).

Another interesting observation regarding the downstream pointwise histograms is their

staircase-shape, ending in a precipice. Such a sh^ is visible not only for a "single-bar" initial size distribution (cases "C" and "D"), but also for the initially "bimodal** size distribution (case

••B").

CONCLUSIONS

To sununarize, the influence of initial droplet size distributions on polydisperse spray jet

far-Held diffusion flames has been analyzed. A comparison of four different kinds of initial droplet

size distributions has demonstrated the direct link between these initial conditions and the spatial

spreading of the fuel droplets and vapors in the jet's far-fleld, emphasizing the importance of the

small droplets when they are present in the initial size distribution beside the large ones. Finally, a "bimodal" size spray and an "average" size spray have been compared showing the differences in fuel

vapor distributions and in other flame features.

1

o.e

o.-r -I

o.e

o.s

0.4

0.3

0.2 -i

o. 1

o.o

o.e

O.T

o.e -

o.e

0.4 -

0.3

o.a -

o. 1

o.o

(a.) Kxael Vapor Fl\j.>c

(t>) Liqiaid. F'Tj.el Flxax

o.oo o.oo O. lO O. IS

X Cm)

In.it.la.1 sizes: — 20 — 2S.2

- -4-0 — 50.4

o.so 20 — ss.s.eoeo— loo.s

0.30

lOO.S

Fig, 1: Variation of fiiel-vapor flux and liquid-fuel flux along the axial direction, (each point on the

graph represents integral values of the fluxes over all radial distances of the jet).

336
1"

i"

i"

I

•

Spray pointwise histograms

J

utA.

I"

i.i

i'

TiI.J

I"

1

J

• L

I" J

i

i.i

•

t-Mft-i.iT«-i a.t

J"

1 i J

I"

r..

•.fl

l>J* B>«.tTt-4

Contours o/ spray mass /racKons

>-Oo e.oa o. le o.»« o.ao X

A. Inilial sizes 20-25.2 ym

Spray pointwise histograms

r

i:

I

l«0 >J

I—I J • *• I" •I—«.l>-l -J J I - — c

a.

•.i»fl-a.iM->

•J

3

Contours of spray mass fractions

B. Initial sizes 20-25.2,80- 10O.B/ini

Fig. 2: Spatial distribution of spray mass-fractions and spray pointwise histograms.

337
CDa

CD 2aouIIH

tab

o

P3 1

(D0 1 0H1 0

(Dn i ' " N

•HMt

g.0u0 1

'. // «

4*HCd

.... .i »>r*rTror.->rfTT

00

-000

X 00

oc

oB

CO

ill

338
TABLE 1

Sectional Vaporization Coefficients

>pr 1 1 on

nn

1 Droplet Diameter |

1 1

1 (/-m) 1

Number of

Monomers '

1 (tAm K^m. ^ 1 ) 1

J

1 0.0 - 20.0 1 0.011250 0.003749

2 20.0 - 25.2 2 0.006672 0.002340

3 25.2 - 31.8 3 - 4 0.004178 0.001498

4 31.8 - 40.0 5 - 8 0.002657 0.000937

5 40.0 - 50.4 9 - 16 0.001668 0.000591

6 50.4 - 63.5 17 - 32 0.001051 0.000372

7 63.5 - 80.0 33 - 64 0.000662 0.000234

8 80.0 - 100.8 65 - 128 0.000417 0.000148

9 100.8 - 127.0 129 - 256 0.000263 0.000093

10 127.0 - 160.0 257 - 512 0.000166

Monomer

E(T) = Diameter:

980.0(1.0+7.

20.0(/Am)

423E-7(T--300. 0)**2.7548) ^mVs)

REFERENCES

1. Chigier, N.A., "Drop Size and Velocity Instrumentation", Progress in Energy and Combustion

Science, 9, (1983), pp. 155-177.

2. Chigier, N.A., "Group Combustion Models and Laser Diagnostic Methods in Sprays (A Review)",

Combustion and Flame, 51, (1983), pp. 127-139.

3. Chigier, N.A., Ungut, A. and Yule, AJ., "Particle Size and Velocity Measurements in Flames by

Laser Anemometer", Seventeenth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, (1979), pp. 315-324.

4. Beretta, F., Covaliere, A. and D'Alessio, A., "Drop Size and Concentration in a Spray by Sideward

Laser Light Scattering Measurements", Combustion Science and Technology, 36, (1984), pp. 19-37.

5. Levy, Y. and Timnat, Y.M., "Two Phase Flow Measurements Using a Modified Laser Doppler Anemometry

System", Single and Multiphase Flows in Electromagnetic Fields: Energy, Metallurgical and Solar

Applications, (H. Branover, M. Monds and P. Lykudis, eds.), Progress in Astronautics and

Aeronautics Series, AIAA, New York, 1985.

6. Stevenson, W.H., "Spray and Particulate Diagnostics in Combustion Systems. A Review of Optical

Methods", Central States Section Meeting of the Combustion Institute, West Lafayette, 1978.

7. Jones, A.R., "A Review of Drop Size Measurements - The Application of Techniques to Dense Fuel

Sprays", Progress in Energy and Combustion Science, 3, (1977), pp. 225-234.

8. Trolinger, J.D., "Laser Diagnostic Methods - A Summary", AIAA 16th Fluid and Plasma Dynamics

Conference, AIAA Paper No. 83-1683, 1983.

9. Holve, DJ., "In situ Optical Particle Sizing Technique", J. Energy, 4, (1980), p. 176.

10. Swithenbank, J., Bedr, J.M., Taylor, D.S., Abbot, D. and McCreath, G.C., "A Laser Diagnostic

Technique for the Measurement of Droplet and particulate Size Distribution", in Experimental

Diagnostics in Gas Phase Combustion Systems, Prog, in Astronautics and Aeronautics, 53, (1977), p. 421.

339
11. Presser, C, Gupta, A.K., Santoro, RJ. and Semerjian, H.G., "Droplet Size Measurements in a

Swirling Kerosene Spray Flame by Laser Light Scattering", ICLASS-85 Meeting, 2, (1985).

12. Bachalo, W.D., Brena de la Rosa, A. and Rudoff, R.C., "Advances in Diagnostics for Complete Spray

Characterization", ICLASS-88, (1988), pp. 287-295.

13. Edwards, C.F. and Rudoff, R.C., "Structure of a Swirl-Stabilized Spray Flame by Imaging, Laser

Doppler Velocimetry, and phase Doppler Anemometry", Paper presented at the Twenty-Third Symposium

(International) on Combustion, Orleans, France, (1990).

14. Presser, C, Gupta, A.K., Avedisian, C.T. and Semerjian, H.G., "Fuel Property Effects on the

Structure of Spray Flames", Paper presented at the Twenty-Third Symposium (International) on

Combustion, Orleans, France, (1990).

15. Chigier, N.A., Ungut, A. and Yule, AJ., "Particle size and Velocity Measurement in Flames by

Laser Anemometer", Seventeenth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, (1979), pp. 315-324.

16. Yule, AJ., Ah Seng, C, Felton, P.G., Ungut, A. and Chigier, N.A., "Laser Tomographic

Investigation of Liquid Fuel Sprays", Eighteenth Symposium (International) on Combustion, The

Combustion Institute, (1981), pp. 1501-1510.

17. Yule, AJ., Ah Seng, C, Felton, P.G., Ungut, A., and Chigier, N.A., "Sprays, Drops, Ehjsts,

Particles: A Study of Vaporizing Fuel Sprays by Laser Techniques", Combustion and Flame, 44,

(1982), pp. 71-84.

18. Yule, AJ., Ereaut, P.R. and Ungut, A., "Droplet Sizes and Velocities in Vaporizing Sprays",

Combustion and Flame, 54, (1983), pp. 15-22.

19. Chigier, N.A., Energy Combustion and Environment, McGraw-Hill, 1981.

20. Tambour, Y., "A Lagrangian Sectional Approach for Simulating Droplet Size Distribution of

Vaporizing Fuel Sprays in a Turbulent Jet", Combustion and Flame, 60, (1985), pp. 15-28.

21. Tambour, Y., "Coalescence of Vaporizing Kerosene Fuel Sprays in a Turbulent Jet", Atomisation and

Spray Technology, 1, (1985), pp. 125-146. See also: "Simulation of Coalescence and Vaporization

of Kerosene Fuel Sprays in a Turbulent Jet: A Sectional Approach", AIAA/ASME/ASEE 21st Joint

Propulsion Conference, Monterey, 1985, AIAA Paper No. 85-1315.

22. Greenberg, J.B. and Tambour, Y., "Far-Field Coalescence Effects in Polydisperse Spray Jet Diffusion Flames", Twenty-First Symposium (International) on Combustion, The Combustion Institute,

(1986), pp. 655-663.

23. Tambour, Y. and Greenberg, J.B., "Fuel-Injector Far-Field Temperature Distributions Resulting From

Coalescence, Vaporization and Combustion of a Polydisperse Spray", ICLASS-88, (1988), pp. 431-438.

24. Tambour, Y., "A Sectional Model for Evaporation and Combustion of Sprays of Liquid Fuels", Israel

Journal of Technology, 18, (1980), pp. 47-56.

25. Tambour, Y., "Transient Mass and Heat Transfer From a Cloud of Vaporizing Droplets of Various Size

Distributions: A Sectional Approach", Chem. Eng. Commun. 44, (1986), pp. 183-196.

26. Tambour, Y., "Vaporization of Polydisperse Fuel Sprays in a Laminar Boundary Layer: A Sectional

Approach", Combustion and Flame, 58, (1984), pp. 103-114.

27. Spalding, D.B., GENMIX: A General Computer Program for Two Dimensional Parabolic Phenomena,

Pergamon Press, 1977.

340
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper34

APPLICATION OF THE SINGULAR VALUE DECOMPOSITION TO THE

INVERSE FRAUNHOFER DIFFRACTION PROBLEM

S. Huzarewicz, G.W. Stewart' and C. Presser

Chem. Sci. and Tech. Lab, National Institute of Standards and Technology, Gaithersburg, MD, U.S.A.

*Department of Computer Science, University of Maryland, College Park, MD, U.S.A.

ABSTRACT

A method for solving the inverse Fraunhofer diffraction problem is presented. The singular value

decomposition is used to obtain a solution for the linear system of equations relating an unknown size

distribution to a measured diffraction pattern. By using only the larger singular values, an approximate

solution is calculated which solves the problem to within the accuracy warranted by uncertainties in the

data. The method is stable in the presence of noise, and it is capable of resolving multimodal distributions.

Results for both numerical simulations and experimental data are presented.

INTRODUCTION

The size distribution of an ensemble of particles can, in principle, be determined from light scattering

measurements. When the particles are larger than a few microns, the angular distribution of scattered light

is readily calculated from Fraunhofer diffraction theory. The inverse problem, which is the determination

of the size distribution from the diffraction pattern, is more difficult. With varying degrees of success, an

iterative method [1], integral transforms [2], and a regularization method [3] have all been used to solve the

problem. In this paper a quick, easy-to-use procedure, which relies on the singular value decomposition, is discussed.

Swithenbank et at. [4] describe an apparatus, represented schematically in Fig. 1, which measures the

angular distribution of scattered light in the near-forward direction (the diffraction pattern). The first lens

to the right of the spatial filter coUimates the expanded incident beam. The detector, which is a twodimensional array of concentric, semi-circular photodiodes, is one focal length from the transform lens, so it

is effectively an infinite distance from the particle cloud. The two lenses provide adequate, but not the only,

conditions for Fraunhofer diffraction; they ensure that the incident light and the diffracted light are parallel.

Additionally, the particles must be larger than the wavelength of the light. For the sake of simplicity, the

particles are assumed to be opaque spheres in the formulation of the scattering problem.

THEORETICAL BACKGROUND

Highlighting the mathematical development of Swithenbank ei al.[4], the intensity of light diffracted

by a sphere of radius a is:

I = Io kau (1)

where k = 2ir/\ is the wavenumber, A is the wavelength of the light source, u « r/f is the scattering angle,

r is the radial distance at the plane of detection, / is the focal length of the transform lens, Iq is the intensity

at the center of the pattern, and J\ is a Bessel function of the first kind of order one. This function is shown

341
in Fig. 2. The fraction of the total incident energy, L{u)o), contained within a circle of radius in the image

plane, centered on the geometrical image, is obtained by integrating the right hand side of Eq. (1) over the

enclosed area to give

L{uo) = 1 - J^{kauo) - J^{kaLOo) , (2)

where Jo is a Bessel function of the first kind of order zero. The energy, E(ui , W2), contained within concentric

circles of radius wi and W2 is

^(wi,W2) = Crra^ [{j^{kaui) + J^(kaui)) - {J^{kau2) + Jl{kau2))] , (3)

where a;2 > wi and C is a constant that depends on the laser power.

Photodiode Array

Laser

Particle

Field

-/

Beam Expander

Spatial Filter Transform

Lens

Photodiode

Array

Incident Light

Figure 1 : Schematic of the Fraunhofer diffraction

apparatus.

Figure 2: Fraunhofer diffraction for a sphere.

1

2

The function I = Iq

2Ji(kc

kaw

In the absence of multiple scattering, Eq. (3) can be written for a polydispersion as m

E{uii,U2) = C-K^Niof [{JlikaiUx) + Jl{kaiUJi)) - {jUkaiU^) + Jl{kaiUJ2))] , (4)

i=l

where the sum is over m sizes and Ni is the number of spheres of radius a,-. Equation (4) may be easily recast

in a form relating the energy to the weight rather than to the number of particles since AT,- = 3Wi/4Trafp , where Wi is the total weight of particles of radius a,- and p is the density of the particles, which is assumed

to be size independent. Upon substitution, the result is

(5)

1=1

where C = 3C/4p. Since the incident beam is not of uniform intensity, it is necessary to assume that

the particles are moving through the beam without a correlation between size and velocity and that the

measurement time is sufficiently long so that a representative portion of the sample passes through the

beam.

If there are n photodiodes, one can write n linear equations in the form of Eq. (5)

Tw = e , (6)

where e is a vector with n components e,-, which is the light energy falling on diode i in the detector array,

T is an n X n matrix with components i,j , which is proportional to the light which has been diffracted by a

particle of diameter dj falling on the i"* diode, and lo is a vector with n components wj , which is the weight

fraction of particles of size dj . The expression for tij is

" dj

j2 f

T^djr2i-i \ . ,2 (

Trdjr2i-i W / 2 / T^djr2i \ 2 ( ^•^2A A

(7)

342
where r2i-i and r2i are the inner and outer radii of the i*'* diode. The constant C need not be included

since the results are to be normalized, that is, Wi = l^£/^"=i

Physically, Eq. (6) means that any diode receives scattered light from the entire range of particle sizes.

However, the maximum energy diffracted by a sphere of diameter dj falls on the perimeter of a circle (at the

detector) with radius r, which is determined from the solution of

0,

(which is the derivative of Eq. (2)) where J2 is the Bessel function of order two. The solution, calculated

numerically with Newton's method, is irdjr/Xf = 1.356602 , or

dj = 1.356602— . (8)

7rr

Using this result, a mean particle diameter can be associated with each diode. Likewise, the lower and upper

limits of each size class are determined from the outer and the inner radii of the diode, respectively. Any

size classes may be chosen as long as the minimum and maximum particle diameters present in the sample

are encompassed and the smallest size is larger than the wavelength.

This work was motivated by the need to rapidly obtain a size distribution from the measured diffraction

pattern for the purpose of process control. Using the singular value decomposition, the size distribution can be determined directly, without the need to invoke an iterative procedure.

THE SINGULAR VALUE DECOMPOSITION

The singular value decomposition of a matrix T is a factorization

T = USV'^

of T into the product of orthogonal matrices U and V and a diagonal matrix 17 = diag(<Ti, cr2, . .

. , cr„),

where T indicates the transpose. The numbers <Ti, which are called the singular values of T, are nonnegative

and are traditionally ordered so that <ti > (T2 > • • > (Tn. There are efficient, stable algorithms for computing

the singular value decomposition [5, 6].

The solution of the system Tw = e, may be expressed in terms of the singular value decomposition as

follows. Let U and V be partitioned by columns in the form

U = {uiU2 ... Un) V = (Vi V2 ...

If T is nonsingular, its singular values are all nonzero, and by the orthogonality of U and V we have

" vJe

w = T-^e = VS-^U'^e = V -^vj . (9)

When some of the singular values are small compared to crj , the solution is ill-conditioned; that is, it

is very sensitive to perturbations in e and T. Stable approximate solutions may sometimes be obtained by

truncating the sum, Eq. (9):

^(fc) = ^_J_„^.. (10)

In ill-posed problems, such as the one under consideration here, the truncated solutions behave in the

following manner. As k increases, the accuracy of w^''^ increases up to a point, after which the solutions

become increasingly inaccurate. To see why this happens, note that from Eq. (9)

,. 1,2 ^^l"7«l'

343
Solutions obtained by application of Eq. (10) to the scattering signature calculated from Eq. (11) with

superimposed noise for a log-normal distribution with <Tg = 0.5 and Dg = 45 /xm are presented in Fig. 3. The

limits of integration were dmin = 6 fim to dmax = 500 fim. The input distribution was obtained through

quadrature of the probability density function over size classes corresponding to those of the diffraction

instrument. In the absence of noise, the reconstructed distribution was indistinguishable from the input

distribution.

The Sauter mean diameter, D32, is rPiD)dD

^32 = TOO , , f^PiD)D-^dD (SLi«;.dr')

where P{D) is on a weight basis. For the original distribution, D32 = 39.7 fim while for the distributions

determined from the noisy data D32 = 39.6 fim (0.3% noise) and D32 = 39.4 fim (2% noise).

The solutions for a Rosin-Ranunler distribution with a = 50 fim and fS = 3.3, are shown in Fig. 4. For the original distribution, D32 = 38.6 fim while for the distributions determined from the noisy data

D32 = 38.7 fim (0.3% noise) and D32 = 38.6 fim (2% noise).

15

c

o

QJZ

10-

7. 0 0)

-5

- Input Distribution

0 0.3% Noise

2% Noise

10° 10' 10'

Diameter (/im)

Figure 3: Reconstructed log-normal distribution {<Tg = 0.5, Dg = 45 fim).

20

-n 15-

c

a>o

10-

S 5H

CJ) 5 0

10'

-5

10

- input Distribution

0 0.3X Noise

A 2X Noise

-If I r Ti ry T' ' t T I I I T I 10' 10' 1 Diameter (/xm)

Figure 4: Reconstructed Rosin-Rammler distri

bution {a = 50,/? = 3.3).

0'

For both cases, in the presence of 0.3% noise, the first twelve singular values {k = 12) were used, while

for 2% noise only the first nine singular values {k = 9) could be used. This suggests that the solution can be determined directly, without the need to calculate the residual at each step in building the solution, once

the magnitude of the noise has been determined.

The solutions for the 50 fim monodispersion shown in Fig. 5 are less satisfactory. Although the principal

mode is located correctly, several small, fictitious modes appear. Even the solution corresponding to the

noiseless data indicates this. Part of the problem lies in the derivation of T. None of the dj from Eqn. (8) is 50 fim, so a meaningful, exact solution of Tw = e does not exist. This suggests that a choice of size bands

other than those dictated by Eq. (8) may improve the results.

The solutions for the polydispersion made up of the sum of three log-normal distributions are shown in

Fig. 6. The probability density function is simply P{D) = Pi{D) + P2{D) -\- P3(D) where the parameters

associated with the subscripts 1,2, and 3 are ((Tgi,Dgi) = (0.3,25 fim), {(Tg2,Dg2) = (0.15,63 fim), and

{(Tg3, Dgs) = (0.2, 125 fim). For noiseless data, the original distribution was recovered exactly. With added

noise, the modes are correctly resolved although some of the weight fractions are off the mark.

In all of these simulations, the noise added to a particular e,- was in many cases on the order of 20 — 30%

of that e,- which is unrealistically large. In order to get a better estimate of the noise, some experimental

data were gathered.

344
From physical considerations we know that w is bounded. Consequently, the inner products uje must

decrease at least as feist as the aj. (Actually, since the small singular values correspond to high frequency

components in the solution, the inner products tend to decrease more swiftly than the singular values.)

Now this decrease in the inner products would cause no harm if we could calculate them accurately.

However, because e is inevitably contaminated with error, small inner products cannot be computed accurately. To see this, let us assume that what we actually measure is the vector e + g, where the components

of g are uncorrelated, with mean zero and standard deviation s. Then the inner products we calculate are vje + vjg. Since \\vj\\ = 1, the standard deviation ofvjg is s, which is independent of j. For small values

of j, the inner products are generally quite large and are computed accurately. But as their value decreases

toward s, they become increasingly inaccurate, until finally they are overwhelmed by the error term vjg.

This error appears in the solution as the vector

whose size is approximately s/aj. Since the aj are decreasing, the error in the solution grows.

This informal analysis suggests a heuristic for bounding k. Estimate the errors in e and scale the rows

of the equation Tw = e so that all the errors are roughly equal to a single value, say s (n.b., the scaling

must be done before the singular value decomposition is computed). Then k should be chosen so that the

inner products uJe remain larger than s. Note that our heuristic does not give a predetermined value of k —just an upper bound. A more

delicate analysis would take into account the size of the error compared to the size of the solution and the

errors introduced in the derivation of T. However, formulas based on such analyses are very sensitive to the

underlying assumptions, and it is better to base the choice of k on simulations and experiments with real

data, to which we now turn.

RESULTS AND DISCUSSION

Simulations

To gain a sense of the accuracy of solutions found through the singular value decomposition, numerical

experiments were carried out. Light energy distributions were determined from the log-normal and RosinRammler (or WeibuU) distribution functions. The log-normal probability density function by weight is

PiD) = ^

exp [-{\nD- InD.f/ia'^] , \/ZTrLJ(Tg

where D is a random variable representing the diameter, > 0 is a scale parameter equal to the geometric

mean diameter, and erg > 0 is a shape parameter equal to the geometric standard deviation. For the

Rosin-Rammler distribution, the probability density function by weight is

PiD) = -^D''-'exp[-iD/af],

where a > 0 is a scale parameter, and ^5 > 0 is a shape parameter. The mode of this distribution is

q;(1 — l/pyl^ if > 1- The constant j3 is & measure of the width of the distribution and a essentially

determines the mode. The theoretical light energy distribution was calculated by numerically integrating

the product of the density distribution function and Eq. (7) over the range of sizes in the .•^..miple

where dmin and dmax are the diameters of the smallest and largest size particles present. To determine

the resolution capability of the method, the theoretical scattering signature for a monodispersion of spheres

with diameter dj was obtained from Eq. (7). The same procedure was carried out for a polydispersion

composed of the sum of three log-normal distributions. Uniform noise levels on the order of 0.3% and

2% of the maximum diode value were superimposed onto the theoretical scattering signatures and the size

distributions were recalculated.

dD , (11)

345
1 00

c75H

a>o

50 ^ 0) Q_

^25

5 OH

-25

10

- Input Distribution

OX Noise

0 0.3X Noise

A 2% Noise

-I—I' TTnj- 1 T I I T I I 10' 10'

Diameter (miti)

15

o

a> ^ 5H

OH

-5

10^ 10

- Input Distribution

0 0.3X Noise

A 2X Noise

10' 10'

Diameter (/xm)

-T—I' r I I IT 0^

Figure 6: Reconstructed polydispersion of the

sum of three log-normal distributions {agi,Dgi) =

(0.3,25 fim), {(Tg2,Dg2) = (0.15,63 /xm), and

l'Tg3,Dg3) = (0.2, 125 /im).

Figure 5: Reconstructed monodispersion of

50 fim spheres.

Experimental data

In the derivation of T it was assumed that the particles were opaque spheres. The derivation is also

valid for circular apertures or opaque discs. A calibration reticle [7], consisting of a random two-dimensional

array of chromium discs on a quartz substrate, was used to obtain experimental scattering data from an

instrument like the one described by Swithenbank et al. [4]. The size distribution of the discs is approximately

Rosin-Rammler with a = 53 //m and /? = 3.2. The average deviations of the signals from the diodes were

experimentally found to be less than 0.3% of the maximum diode signal. The distributions of the signals were not Gaussian due to fluctuations in the intensity of the light.

Results for the reticle are shown in Fig. 7. The circles correspond to the weight fractions of the discs on the substrate, to which a correction has been applied to account for the Gaussian intensity profile of the

incident beam [7]. Calculated from the known distribution, D32 = 40.3 /im [7]; calculated from the inverted

data, Dz2 = 40.2; /im. Despite the good agreement, the residual was somewhat larger than those for the

simulated two-parameter distributions. This is because about 10% of the discs overlap — the diffraction

model does not account for overlap, just how these should be counted is uncertain.

Figure 7: Reconstructed size distribution for the calibration reticle.

346
CONCLUSIONS

An easy-to-use method of directly solving the inverse Fraunhofer diffraction problem has been demonstrated. This method works well for polydispersions even in the presence of noise. When applied to monodispersions or multimodal distributions, it accurately determines the modes although it tends to spread them

out over more than one size class. No technique exists which does not suffer from this deficiency. Because the method is direct, it requires less computational time than iterative procedures, making the use of

diffraction-based particle sizing instruments for process control feasible. A singular value analysis is often

useful when dealing with ill-posed problems. Its application to the system of equations produced by quadrature of Fredholm integral equations of the first kind, which arises from an alternative formulation of the

inverse diffraction problem, will be the subject of a future publication.

ACKNOWLEDGEMENTS

We acknowledge the partial support of the OflBce of Nondestructive Evaluation, Materials Science and Engineering Laboratory, N.I.S.T.

NOMENCLATURE

a, ai radius of sphere

c ratio of consecutive mean diode radii

C,C' constants dependent on the laser power

D, geometric mean diameter, parameter of log-normal distribution

D32 Sauter mean diameter

dj diameter of sphere

rf ff "'max ) "mtn mQVi mi 1 m n mi n imii m ct^ n^Pf^ H 1 a m i^pc XiicLA.iiiiU.iix diiu. iiiiiiiiiiLiiii o^iicic: uiciiiictc:io E energy falling on the image or detection plane

energy falling on i^^ photodiode

/ focal length of transform lens

9 noise vector

I inteniity 01 difiracted light in the image plane

h intensity at the center of the diffraction pattern

Bess ;1 function of the first kind of order 1/ k wavenumber

L fraction of diffracted energy

Ni number of particles n nurriber of terms in summations, matrix dimension

P{D) probabilty density function in the random variable D,

representing particle diameter in this application

r, Vi radial coordinate in the plane of detection or diode radius

T, tij scattering matrix of which tij is an element

T transpose operation

s standard deviation of the noise

U, Uij orthogonal matrix of the singular value

analysis, is an element of U

V, Vij orthogonal matrix of the singular value

analysis, Vij is an element of V

Wi , Wi total weight of spheres, Wi is the cooresponding weight fraction

a scale parameter of Rosin-Rammler distribution function

13 shape parameter of l osin-Rammler distribution function

A wavelength

P

density of a particle

347
diagonal matrix of the singular values, (T,-, of T

geometric standard deviation, parameter of the log-normal distribution function

dimensionless radial distance in the image plane

REFERENCES

[1] Sangeorzan, B.P., Uyehara, O.A. and Myers, P.S., "Time-Resolved Drop Size Measurements in an Inter- mittent High-Pressure Fuel Spray," SAE Paper No. 841361 , Warrendale, PA (1985).

[2] Koo, J.H. and Hirleman, E.D., "A Synthesis of Integral Transform Methods for Reconstruction of Particle

Size Distributions from Forward-Scattered Light," Second Int'l Congress on Optical Particle Sizing, pp.

189-198, Arizona State University, Tempe, AZ (1990)^

[3] Hirleman, E.D., "Optimal Scaling of the Inverse Fraunhofer Diffraction Particle Sizing Problem: The

Linear System Produced by Quadrature" in Optical Particle Sizing: Theory and Practice (G. Gousebet

and G. Grehan, Eds.), pp. 135-146, Plenum Press, New York, NY (1988).

[4] Swithenbank, J., Beer, J.M., Taylor, D.S., Abbot, D. and McCreath, G.C., "A Laser Diagnostic

Technique for the Measurement of Droplet and Particle Size Distribution," Experimental

Diagnostics in Gas Phase Combustion Systems (B.T. Zinn, Ed.), Progress in Astronautics and Aeronautics, Vol. 53, pp. 421-447, AIAA, New York, NY (1977).

[5] Golub, G. H. and Van Loan, C, Matrix Computations, Johns Hopkins, Baltimore, MD (1989).

[6] Dongarra, J.J., Moler, C.B., Bunch, J.R. and Stewart, G.W.. LINPACK Users' Guide. S.I.A.M., Philadelphia, PA (1980).

[7] Hirleman, E.D. and Dodge, L.G., "Performance Comparison of Malvern Instruments Laser Diffraction

Drop Size Analyzers," Third Int'l Conf. on Liquid Atomization and Sprays, pp. 8564/1-8564/14, The

Institute of Energy, London, England (1985).

348
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 35

THE ANGULAR PATTERN OF THE SCATTERED LIGHT AND THE

LORENZ-MIE THEORY FOR THE STUDY OF OH. SPRAYS

F. Ossler*, F. Beretta^ A. D'Alessio' and A. Tregrossi'

*Dipartimento di Ingegneria Chimica, University "Federico 11", Napoli, Italy

^Istituto di Richerche sulla Combustione, CNR, Napoli, Italy

Abstract

The angular pattern of the polarization ratio 7(^)=Chh(^)/Cw(^) of

light scattered by droplets has been computed according to the Lorenz-Mie

theory in the size range from 0.02-^50/xm. It was found that the scattering due

to the smaller droplets with typical size around l/zm is dominated by surface

waves effects and the angular pattern presents relevant polarization peak in

the backscattering region. This scattering feature has been successfully applied

to the characterization of aerosols of light and heavy oils with a Junge size

distribution function. The angular pattern of light scattered by sprays of

heavy oil atomized by air assisted nozzle presents a similar maximum of the

polarization ratio in the backscattering region, thus showing that a large

concentration of micron-sized droplets are present also in this condition. A

quantitative comparison of the data with the Lorenz-Mie Theory indicates that

the overall size distribution of the sprays is bimodal and that the atomization

efficiency has large influence on the ratio of the number concentration between

the smaller particles and those larger than 20/im.

Introduction

The determination of the size distribution of the droplets is certainly

one of the most important problem for studying atomization and spray evolution.

The greatest part of the equations proposed in the current literature derives

directly or indirectly from measurements method based on visualization of the

droplets (1) . Now in practice it is almost impossible to visualize a droplet

smaller than 3-4/im employing radiation in the visible range, according to the

Rayleigh criterium (2); by consequence the lower part of the size distribution

is neglected or improperly taken into account.

Light scattering does not suffer this limitation and allows, in line of

principle, to determine the size and the optical properties of droplets of any

size. In fact when the droplets can be considered as spherical objects the

Lorenz-Mie theory provides an exact framework for the interpretation of their

experimental angular distribution of the scattered light and their extinction

coefficient. The systematic measurements of the angular pattern of the polarized components of the scattered light and their comparison with the theory

have been carried out by our group in the last years in single droplets furnaces, sprays and flames (3-6). It was found that the polarization ratio 7(^)

between the horizontally polarized scattered light and the vertically polarized

one is a quite sensitive function of the size and optical properties of the

349
scatterers. More particularly, the polarization ratio in the side scattering

region around ^=90° allowed to estimate the presence in the spray of droplets

much smaller than those normally reported in the literature. However these

studies were based on the hypothesis that the size distribution function was

monomodal; an attempt to obtain the droplets size distribution of a light oil

spray from the inversion of the scattering spectrum in the u.v. , across the

threshold of the absorption band, did not give conclusive results for the

numerical instability of the inversion method (7).

This paper has the purpose of measuring and interpreting not only the

side scattering region, but the whole polarization pattern, with particular

emphasis to the backscattering regions, in order to asses the presence and role

of droplets smaller than 3-4/im. A preliminary theoretical section contains some

numerical results of the Lorenz-Mie theory and summarizes the different regimes

of scattering from the limit of very small particles where they behave as

Rayleigh dipoles to the limit of very large spheres where the scattering is

described by ray optics. The experimental section discusses firstly the results

obtained on aerosols produced from both light and heavy oils in comparison with

the prediction of the theory. The same type of analysis is then extended to

sprays of heavy oils, produced by a home designed internal air assisted nozzle.

The final part of the communication discusses potentialities and limits

of the scattering diagnostics and point out the relevances of the results on

the size distribution of the smaller droplets for a more deep insight on the

processes of atomization and spray evolution.

The angular patterns of the polarization ratio computed by the Lorenz-Mie

theory

The Lorenz-Mie theory provides expressions for the angular scattering

cross sections in terms of infinite series (2,8):

C^(^) = [2^) ? { ^n^n(cos^)+bnrn(coso}

| (1)

^Hh(^> = [2^] ? {

bn^n(cos^)+anrn(cos^)|

| (2)

where Cw(^) and Chh(^) the vertically and horizontally polarized scattering cross sections respectively and 6 is the scattering angle. The dependence

of the scattering on the size parameter a=7rD/A and on the complex refractive

index m=n-ik is contained in the functions a^ and b^, wich represent the

strength of the electric and magnetic multipoles of the scattered field, while

its angular distribution is determined exclusively by the functions

7rn(cos^)=Pn(^) (cos5)/sintf and r^Ccos^ )=d[Pn(^) (cos^ ) ]/d^ , where Pn^^^ are

associated Legendre functions of the first kind and order n. A discussion of

the properties of the Lorenz-Mie series can be carried out only numerically and

fig.l illustrates how the ratio )=Chh(^)/Cw(^) > defined as the polarization

ratio, varies with the sizes of the droplets in the range 0.02-5-50/im at the

reference wavelength of 500nm, for two kinds of hydrocarbon mixtures. The right

side of fig.l contains the results computed for a typical light oil mixture

transparent in the visible and the left side refers to a heavy oil mixture

partially absorbing at A=500nm. When the size of the droplets is very small

respect to the incident light wavelength, only the electric dipole term of the

series of eq. 1-2 contributes to the cross sections and the sphere scatters

according to the Rayleigh law with 'y(0)=cos'^ 9 . If the size is slightly increased the magnetic dipole term and electric quadrupole terms are relevant and the

350
10

10

10

10.00 f

r 5.00 o

q 0.00 o

_No 4.00 o

Q.

2.00

0.00

1.20

0.60

0.00

Fig.

30

I "

I

I

60 90

I 1

I 1 I

1 ' 120 150

' D > 0.5 Itm

' D - IJ) iifn D - 1.5 /un

10'

10

10

10.00 .

0.00

I r I—I—I—1—r- 30 60

I—r-" 90 120 150 180

Scattering Angle (deg)

1 Angular patterns of the polarization

ratio computed by the Lorenz-Mie

theory for heavy oil (left) and light

oil (right).

minimum of the polarization

becomes shallower and shifts in

the backward direction, as^ it

appears for the first curves of

the upper part of fig.l. As the

diameter of the droplets

increases , more a^ and b^

functions contribute appreciably to the series. It can be

shown that the infinite series

5.00 can be truncated approximately

after n* terms with n*=^

a+(a)^/^ and the terms included

in the interval a-(a)^/^ and

a+(a)^/^ defines the "edge

domain" (9) . The a^ and b^

functions in this range present

strong resonances with the size

reaching values much higher

than the other non resonant

terms of the series (10) . A

1^1.20 physical interpretation of this

behaviour is that when the

i0.6Q sphere is rather small the

incident radiation interacts

0.00 more strongly with a limited

region above and below the

surface of the particle than

with its internal part. Thus

the scattered light has mainly

the character of a series of

surface waves travelling near

the cross sections in this size

:4.00

2.00

the boundary of the sphere (11). Furthermore

range shows also strong variations with the scattering angle due to the oscillatory behaviour of the first associated Legendre functions which define their

angular distribution. The angular pattern of polarization becomes a very

sensitive function of the size of the scatterers but many of its details are

confusing and not relevant when an ensemble polydispersed spheres is considered. Therefore the computations for the size range from D=0.5/im to D=5/im presented in the central part of fig.l are obtained by averaging the monodisperse

case over a AD interval of ±0.254/im and a A^ interval of ±1.6°. It appears that

the resonances peaks are smeared out in the greatest part of the angular

pattern, where the polarization ratio is close to one, with the evident exception of a relevant peak of positive polarization in the backward region. The

angular position and the values of the polarization ratio at the peak changes

noticeably with the size and the real part of the refractive index while the

absorption index has a lesser influence, as is expected for particles with very

large surface to volume ratio. Physically the constructive interference of

surface waves in the backscattering is related to complex, and still not

understood, mechanisms which are responsible of the glory for larger spheres

(9). For droplets larger than 5/Ltm the number of a^ and b^ functions which

contribute significantly to the series becomes larger and consequently the non

resonant terms outside the edge domain are progressively more relevant. Following the van de Hulst (12) correspondence principle these terms are associated to rays with different impact parameters and are diffracted, refracted or

351
reflected, according to the laws of ray optics.

The difference between transparent and absorbing droplets becomes quite

important in this scattering regime, as it is illustrated by the lower part of

fig.l where the angular pattern of polarization ratio are reported for transparent and absorbing droplets ranging from 10/im to 50 /zm. For the transparent

case, shown in the right side of the figure, the 7-ratio in the forward follow

quite closely that predicted by the geometrical optics considering the contribution of refraction and external reflection. The minimum of the polarization

ratio in the backward around tf=160* is instead due to the developing of a

strong rainbow which is preferentially polarized in the vertical plane. Scattering effects due to surface waves are still recognizable in the side scattering region where the polarization ratio is higher than that predicted by

external reflection, and in the backward, beyond the rainbow, where strong

oscillations due to glory effects are present.

The angular patterns of the polarization ratio for partially absorbing

droplets show quite different trends, as it appears from the lower left side of

fig.l. There is no maximum of the polarization ratio in the forward because the

refracted rays are more and more absorbed inside the droplets as its size

increases . For the same reason the minimiam of polarization at the rainbow angle

barely is present for lO^xm droplets and disappears for larger spheres

.

Also surface waves are heavily damped by absorption and no glory characteristics develop in the backward. Therefore 50/im droplets have an angular

polarization pattern identical to that of a purely reflecting sphere with a

deep polarization minimum at the Brewster angle.

Experimental Apparatus and Procedures

Experiments were carried out on clouds of droplets of commercial oils: a

light oil with m=1.48-i-1.38-10-5 and a heavy oil with ra=1.55-i-l. 00-10-2 ^t

A=488 nm. The complex refractive indexes of the fuels were determined in the

following ways: the real parts were determined measuring the Brewster angle

while the imaginary parts were determined using a spectrophotometer.

The clouds were generated by two methods: the finer distribution were

obtained employing

a commercial

pneumatic aerosol

generator (the

heavy oil was

diluted with 90%

dichloromethane in

order to reduce

its viscosity)

;

the sprays were

produced using an

internal mixing

air assisted

nozzle, whose

scheme is reported

in the left side

Fig. 2 Air assisted nozzle (left) and optical setup of fig. 2. The fuel

(right) . passes through the

central duct and

interacts with the swirled atomization air, that passes trough four tangential

small ducts, in a small chamber located at the top of the nozzle. The spray

then comes out from an cylindrical orifice of 0.9 mm ID.

A sketch of the optical setup used for the scattering measurements is

352
reported in the right side of the fig. 2. The light source used was the blue

line of an Ar-ion laser (A=488 nm) plane polarized at 45° in order to obtain

two polarized components, vertical and horizontal, with same intensity in the

scattering volume. The collecting optics can rotate around the scattering

volume in order to continuously measure the angular profiles of the scattered

light. A polarizing cube, inside this optic, discriminates the two polarized

components that are simultaneously measured by two lock- in amplifiers and sent

to a micro computer. A check of the experimental setup was performed verifying

the agreement between the polarization ratio angular pattern of a stream of

monodisperse droplets and the corresponding calculated one.

Experimental Results

Typical angular distributions of the polarization ratio of the light

scattered by aerosols of light and heavy oils are illustrated in fig. 3. The

polarization is almost uniform in the forward and side scattering regions with

horizontal polarization is found in the

backward region near ^=155° for the

heavy oil aerosol. A qualitative

comparison of the experimental results

with the theoretical predictions of

fig.l evidences that the scattering is

mainly determined by surface waves

effects, typical of droplets with

diameter between 0.3/zm and 5/zm.

For a quantitative comparison,

the polydispersion of the droplets

should be taken into account because

the measured polarization ratio is

given by the expression:

/f(D) CHH(D.^) dD

liB)

Jf(D) Cw(D.^) dD

where F(D) is the number size distribution function of the droplets.

Junge has proposed for size

distribution of aerosols an inverse

power law F(D)a:D"P and found p=3 for

atmosphere aerosols. In the present

case the Junge distribution was modified by neglecting the contribution of

droplets larger than lO/zm and by considering constant the number concentration

of the droplets smaller than a cut-off diameter Dq:

F(D) = c .02 /im < D < Do

F(D) = c (Do/D)P Do < D < 10 /xm (4)

F(D) =0 D > 10 jum

The integrals of eq.3 were computed with integration steps of AD=0.02 /xm

and A0=O.2° (angular interval of ±1.6°) and the parameters Dq and p were

obtained by the best fitting with the experimental angular patterns. It was

found that both the heavy and light oils aerosols patterns were described by

the same size distribution with p-3.0 and Do-O.S/xm. The same value of the

exponent p was obtained from the fitting, when the lower part of the size

distribution was modelled assuming a decrease of the number concentration for

7 close to one while a marked peak of

o

QL

0.0 I—I—1—I—1—I—I—I—I—1—I—'—I—'—'—I—I—I

0 30 60 90 120 150 180

Scattering Angle (deg)

Fig. 3 Measured (noisy curves) and

computed angular patterns of

the polarization ratio for

aerosols of heavy and light

oils

.

353
sizes smaller than Dq , thus showing

that droplets smaller than 0 . 2/im do not

contribute to the integrals . Having

demonstrated the sensitivity of the

polarization pattern for the characterization of aerosols, other experiments

were devoted to the characterization of

sprays of heavy oil produced through

atomization. Typical angular patterns

of the polarization ratio of the light

scattered by sprays atomized at different pressure of the atomization air

are reported in fig. 4. The patterns are

qualitatively similar to those exhibited by the heavy oil aerosol with a

relevant peak of polarization in the

backward scattering region. However

differences are evident in the side

scattering region where the 7 -ratio is

lower than in the previous case and

depends on the quality of atomization.

In fact a minimum with 7=0.5 is found

around ^=80° in the case of the worst

atomization with j^j-=0 , IMPa , thus

showing that also larger droplets contribute to the polarization pattern, as it

is expected.

However, any attempt to fit theoretically the experimental patterns with

a single distribution function did not give satisfactory results. In fact Junge

distributions with different p exponents were able to model the polarization

peaks in the backward but gave too high 7-values in the side scattering.

Viceversa F- functions with a modal value of the size around 5-10/im were able to

model the side region but gave values of 7 too low in the backscattering for

reproducing the maximum observed experimentally.

Therefore it was concluded that the size distribution function of droplets produced by atomization is bimodal and for its modelling it was assumed

that the lower tail of the distribution is described by a Junge function Fi(D),

similar to that employed in the aerosol case, and the upper tail of the larger

droplets was approximated by a monodispersion with a typical size D2 . So that:

aerosol

F(D) = NiFi(D) + N2 5(D-D2)

Then the comparison between the experimental angular patterns and the theoretical ones was carried out in dependence of the size of the larger droplets D2

and of the concentration ratio between the number concentration of the small

droplets and the larger ones N1/N2

.

A preliminary series of numerical integrations of eq.3 showed that it was

not possible to reproduce the experimental 7-ratio in the side scattering

region with D2 smaller than 20/im while it was always possible to find a couple

of D2 and N1/N2 which fits the experimental curves for D2^20/im.

This result means that the larger droplets of the spray scatter the light

according the geometrical optics limit, where the Cyy a^rid C^H cross sections

are both proportional to their surface area and thus their polarization ratio 7

is constant. Therefore it is not possible to obtain more information on their

size from polarization measurements but only their surface area per unit volume

N27rD2/4 can be obtained from these experiments.

Figure 4 illustrates the comparison between the experimental data and the

0.00

n=1.55 k=1.00-10"

-I—I—

r

->—

r

-1—I—I—t- 0 30 60 90 120 150 180

Scattering Angle (deg)

Fig. h Measured (noisy curves) and

computed angular patterns of

the polarization ratio for

heavy oil sprays at different

pressure of atomization air.

354
theoretical bimodal model when D2=20/im: the number concentration ratio N1/N2

between the very small droplets and the large ones is around 400 where the

atomization is carried out very efficiently (Pair=0-^MPa) and drops to 100 for

the worst atomization condition (Pair=0 • iMPa)

.

Discussion

The measurement of the complete angular patterns of the polarized components of the scattered light together with the systematic comparison with the

Lorenz-Mie theory allows to conclude that a very large number of droplets with

diameter smaller than one micrometer are produced during atomization of heavy

oil. In fact the analysis shows that the patterns are dominated by non geometrical optics contribution when both light oils and hevy oils are dispersed as

aerosols and that the same is true for the light scattered in the backward

direction, in the case of heavy oil sprays produced by air assisted atomization. This lower part of the size distribution function is not detected when

optical sizing methods based on geometrical optics or, in equivalent ways, when

visualization techniques are employed. On the contrary the angular pattern of

the polarized components of the scattered light does not offer a great sensibility to the characterization of the lower tail of the size distribution,

composed by droplets larger than 10 /zm, whose scattering properties are mainly

determined by geometrical optics.

Scattering due to surface waves has been employed before by the same

group in the diagnostics of sprays by measuring the polarization ratio at 9=

90" employing a monomodal size distribution function. The measurement of the

complete pattern of polarization shows that a bimodal size distribution,

including both very small and large droplets, is the only choice to interpret

theoretically the scattering both in the backward and in the side scattering

region.

The presence of a bimodal size distribution contributes to a better

understanding of the physics of the atomization and of the convection and the

vaporization processes in sprays. The production of a great number of droplets

smaller than l/xm is not satisfactory taken into accountct by the current models

of atomization, mainly based on linear instability analysis (1) . The great

number of very small droplets suggests that the local efficiency of the momentum transfer between the gase phase and the liquid phase is high, otherwise it

should be impossible to justify the large value of the energy necessary to

overcome the surface tension effects. With regard to spray dynamics it is

worthwhile to remark that in the spray atomized in the best condition of our

experiments (see fig. 4) the ratio of the total perimeter per unit areas between

the small and large droplets is in the order of 10 and the ratios of the total

surface area and of the total volume are in the order of 1 and 0.1 respectively. For the worst atomization case the ratios are 2, 0.1 and 0.01 respectively.

By consequence in the first case 10% of the fuel is in the form of droplets so

small that they are directly convected by the air and have a vaporization

lifetime orders of magnitude lower than the larger droplets. This part of the

fuel is readily available for oxidation and pyrolysis reactions in the gas

phase and they play an important role for the ignition and stabilization of

flames. On the contrary for the worst atomized spray there is practically no

fuel available for this purpose.

From an optical point of view it is worthwhile to remark that the fact

that heavy oil is partially absorbing in the visible allowed to find the non

geometrical optics nature of the polarization in the backward scattering

region, and consequently the high concentration of the very small droplets.

This surface wave effect is not so evident in the angular polarization pattern

of transparent sprays because peaks of preferentially vertically polarized

355
light due to rainbow tend to reduce the overall polarization ratio. In fact

fig. 4 shows that transparent aerosols have a less prominent polarization peaks

in the backward than absorbing aerosols with the same size distribution. The

difference between the transparent and absorbing cases becomes clearer for

sprays where the much larger droplets, produced by atomization, damp almost

completely the refracted rays.

Final Comments

In summary this communication gives a contribution on the following

points

:

i) Analysis of the Lorenz-Mie solution for particles around l/xm shows that

light scattering is dominated by surface wave effects; the polarization peak in

the backscattering is an useful tool to identify the size of spherical objects

in this diameter range

.

ii) This method has been applied to aerosols formed by light and heavy oil

droplets with good agreement with the predictions of the aerosol literature.

iii) The same method applied to a sprays of heavy oil droplets allowed to

detect the lower part of the size distribution function which is not detectable

with other optical techniques. Furthermore the bimodal characteristics of

sprays produced by air assisted atomization has been proved.

References

1. Lefevbre, A.H. , Atomization and Sprays . Hemisphere, New York, NY (1989).

2. Born,M. and Wolf , E. , Principles of Optics . Pergamon, Oxford, UK (1970).

3. Massoli, P., Beretta, F. and D'Alessio, A., "Single Droplet Size, Velocity and Optical Characteristics by the Polarization Properties of Scattered Light", Appl. Opt., Vol 28, p. 1200 (1989).

4. Massoli, P., Beretta, F. and D'Alessio, A., "Pyrolysis in the Liquid

Phase Inside Single Droplets of Light Oil Studied with Laser Light

Scattering Methods", Comb. Sci. Tech., Vol 72, p. 271 (1990).

5. Beretta, F. , Cavaliere, A. and D'Alessio, A. "Ensemble Laser Light

Scattering Diagnostics for the Study of Fuel Sprays in Isothermal and

Burning Conditions" Twentieth Symp. (Int'l) on Combustion p. 1249, The

Combustion Institute, Pittsburgh, PA (1984).

6. Beretta, F. , D'Alessio, A. and C. Noviello, "Spray Vaporization and Soot

Formation in Flames Generated by Light Oil with Different Chemical

Composition", Twenty- first Symp. (Int'l) on Combustion p. 1133, The

Combustion Institute, Pittsburgh, PA (1986).

7. Beretta, F. , Cavaliere, A., D'Alessio, A., Massoli, P. and Ragucci, R. , "A Spectral Scattering Method for Determining Size Distribution Functions

and Optical Characteristics of Droplets Ensembles in Fuel Sprays",

Twenty- first Symp. (Int'l) on Combustion p. 675, The Combustion Institute,

1986.

8. Kerker, M. , The Scattering of Light . Academic Press, New York, NY (1969).

9. Nussenwveig,H.M. , "Complex angular momentum theory of the rainbow and the

glory", J. Opt. Soc. Am. Vol.69, p. 1068 (1979).

10. Pluchino, A.B., "Surface wave and the radiative properties of micron-sized particles" , Appl. Opt. Vol.20, p. 2986 (1981).

11. Ashkin A. and Dziedzic, M. , "Observation of Optical Resonances of Dielectric Spheres by Light Scattering", Appl. Opt. Vol.20, p. 1803 (1981).

12. van de Hulst, H.C., Li ght Scattering by Small Particles . Dover, New York,

NY (1981).

356
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 36

EXPERIMENTAL AND NUMERICAL STUDIES ON DROPLET

DISTRIBUTION IN DIESEL OIL SPRAYS

K.-J. Choi, J. Xin and Y.B. Shen*

Department of Mechanical Engineering, University of Illinois at Chicago, Chicago, IL, U.S.A.

Department of Energy Engineering, Zhejiang University, Hangzhou, P.R. China

ABSTRACT

In this paper, the spray atomization processes of pure diesel oil through a pintle

diesel injector were investigated both experimentally and theoretically. The experiments

were conducted in a combustion chamber applying an in—line holographic measurement

system. An axisymmetric turbulent model for theoretical study was introduced by

assuming the spray atomization as the pre—mixing injection of vapor and droplets; the

eddy viscosity was simulated with a sub—grid scale method. The experimental results

indicated that clear holographic images can be obtained by means of an in—line holographic

system. It was observed that droplet size in the core region is smaller than that in the

outer region. The numerical results of the droplet distributions and the droplet trajectories

were in good agreement with the experimental observation. From this, it was indicated

that the atomization processes in a diesel engine can be simulated to a certain degree using

the present numerical model.

NOMENCLATURE

drag coefficient

resistant coefficient defined in equation (7)

g

acceleration of gravity

m

M

P

momentum exchange term between gas and droplets

pressure

mass flux of spray

radial coordinate or radius of droplet

droplet Reynolds number

Q

rRed

tAt

tot

time

time interval between each group of injected droplets

total time of spray injection

-» U instantaneous velocity, =

time mean velocity

fluctation velocity

u + u'

357
U,u,u' velocity component in radial direction

v,v,v velocity componeni in axiai uireciion

TT rnr\Hnliic di t^Iq + itto \TcAr\f*\^v — 1 TT TT I11UUU1U.9 ui icidiivc vciudiyj — | Ug— pk

X spatial vector

- particle location vector

z axial coordinate

B gaseous volume fraction

P density

u kinematic viscosity

effective viscosity, = v

t eddy viscosity defined in Equation (3)

T dimensionless time, = t/t^Qj

or 6

P particle

k kth computational group of drops

coordinate directions

1. INTRODUCTION

Atomization and spray processes in combustion chambers are of considerable

importance to the design of diesel engines. The basic process involved is the disintegration

of the liquid fuel to produce millions of small droplets to increase the total surface area. To a certain extent, the performance of diesel engines depends on the quality of spray

atomization. Up to now, many fundamental studies have been done in this area. In the

past, much focus was placed on single droplets or macroscopic characteristics of sprays,

such as spray penetration and spray angle [1—3]. With the development of new measuring

techniques and instruments, it has become possible to study droplet size distribution,

droplet evaporation and combustion within sprays in high temperature and pressure

environments [4—6].

The diesel spray atomization is a transient, unsteady, phase—changing process

accompanied with disintegration, collision and coalescence of liquid droplets. This complex

process is influenced by many parameters, such as the properties of liquid, the nozzle

specifications, and environmental conditions. Due to the complexity and random nature of

spray atomization, it is time—consuming, expensive, and sometimes almost impossible to

study the spray characteristics and effects of various parameters on the atomization

processes by means of experiments only. Rather, comprehensive understanding of

microscopic mechanisms of spray formation and development can be obtained by means of

combination of experiments and numerical simulations.

In this paper, the authors intended to develop an experimental method measuring

droplet size distribution in a diesel oil spray by using a single pulse Fraunhofer in—line

holographic system. Then, a computational model for the droplet distribution and drop

trajectory was developed.

2. EXPERIMENTAL APPARATUS AND PROCEDURE

An experimental system similar to the one used in our previous study [7] was

applied. A constant volume combustion chamber was filled with nitrogen to avoid burning

of fuel so that the atomization and evaporation processes without chemical reaction could

be observed. A pintle diesel injector in a cooling water jacket for insulation was installed

358
inside the combustion chamber. The geometric angle of the pintle nozzle was 15°. The

combustion chamber was 100 mm in diameter and 140 mm in height. The size of the two

parallel quartz viewing windows was 50 mm in width and 100 mm in height. The windows'

surfaces were normal to the laser beam.

An in—line holographic system was chosen to study the atomization mechanisms due

to its simpler optical arrangement and less requirement for the resolution of holograms.

Concerning the atomization features of diesel sprays, a very high resolving power of

measurement was required. In order to solve this problem, a pre—magnifying system was

adopted to a conventional in—line holographic system. The present system consisted of 8

lenses making the magnification of the objects constant regardless of the distance between

recorded objects and the holographic plate. A single pulse of 100 Mj at 50 ns pulse width

was generated with a ruby laser. The light rays were expanded and coUimated to 90 mm in

diameter. A He—Ne laser was employed for reconstruction of the holograms. The optical

systems for recording and reconstruction are shown in Fig. 1. The recording region is

shown in Fig. 2.

Injector

Holographic Plate Preamplifying Lens

(B)

Fig 1. In-line holograph arrangement: (A) for recording, (B) for reconstructing

At ambient temperature, diesel spray was so dense that most laser beams cannot

pass through. Therefore, a slit made from two knives was installed right beneath the

injector to block a portion of the spray. At a high temperature in the measuring field,

however, sprays have become dilute because of the evaporation of droplets. It was not

necessary to block any portion of the sprays at media temperatures above 573 K.

359
Spra:

Lazer

Beam

Pintle Injector

T

60 mm

Curly

Motion

Fig 2. Schematic diagram of an oil

spray near the pintle injector

1 . 5 rrm 1.0 ram

140 inra>

Chaiber

100 mm

Fig 3. Coordinates and schematic of

computed sub-grids in a chamber

3. THEORETICAL MODELING

There are few models which can simulate the main steps of the atomization process,

such as breakup, collision, coalescence of droplets and interaction between liquid droplets

and the surrounding gaseous medium. Previous research was mainly emphasized on spatial

fuel vapor concentration using a gaseous jet model [8,9].

When liquid fuel is injected into the combustion chamber under a high pressure

difference, there are many tiny droplets in the initial stage of spray atomization. In this

study, however, it was assumed that tiny drops vaporize instantly right after injection, and

otherwise no vaporization occurs in the spray. Therefore, the spray atomization process

was simplified by considering the spray as a mixture of liquid fuel droplets and fuel vapor.

In addition, no breakup or collision of droplets was assumed. The droplet phase was

treated in the Lagrangian approach, and the gas phase in the Eulerian approach. In order

to simulate an actual spray which has broad ranges of droplet size, velocity, and droplet

numbers, a certain number of groups of droplets was selected. Each group of droplets

injected at different time has Npj^ droplets of the same diameter and velocity. Coordinates

and schematic diagram of the system are shown in Fig. 3. For unsteady and axisymmetric

two—phase flow, the fundamental equations were written in cylindrical coordinates as the

follows.

Gas Phase

Continuity equation:

Momentum equation:

v)ft;g=si-|-vp+v. (fc,vj^)+|^ (2)

360
A sub—grid scale turbulence model [10] was used to evaluate the eddy viscosity

given as

where

S -1

^ij - ^

rdH: dU.i

J ^

(4)

Cg is a constant and A^^ a local grid length.

Droplet Phase

The velocity and acceleration of a droplet belonging to the kth group of droplets are expressed by the following

-3r=^pk (5)

-ar = g +

is;; - upk) (6)

The resistant drag coefficient, Dj^, for the specific droplet is given as

Dk = ^^f^g^k + I '^kPg^^m (7)

The first term in the right hand side of Eq. (7) is Stokes drag coefficient for Re^ < 1, and

the second term is for large Re^j.

Nnmerical Methods and Initial Conditions

The above differential equations are discretized into finite difference equations using

a variable spacing grid svstemvs based on a time step. The detailed numerical procedure can

be referred in literature [11]

The gaseous volume fraction in a grid is presented as

m

^ = 1 - X '^kNpk'^CJ - x;k)] /AV (8)

where, AV is the volume of a grid, 5{x) Dirac delta function, m the group number of liquid

drops in the grid (0-2500), and Np^ the number of drops represented by the kth group.

Assuming that each group of droplets enters the chamber at a different time, Npj^ can be

determined by the following equations:

4

pk i ^k ^p = Q^t (9)

361
The time interval between groaps of injected droplets, At, was 6 x 10'^ seconds and the

total injection time was 1.5 x 10 seconds. Therefore, the maximum number of groups

used in this study was 2500. Momentum exchange between gas and droplets is calculated

as

M = -^ [Dk(Ug-u;j5(J-^,)]/AV (10)

k

Prior to injection of the mixture of fuel drops and fuel vapor, the environmental gas

medium is assumed to be quiescent. The gas temperature and pressure are 450 K, 0.9

MPa, respectively. According to the results of our previous experiments [12], the initial

velocity of spray was selected as 170 m/s. The drop diameters were taken at random

within a range between 6 /mi and 60 /mi, and the initial location of droplets at the exit are determined randomly between 1 mm and 1.5 mm in the radial direction from the centerline

of the chamber. The liquid flow rate used in this calculation was 13.3 ml/sec.

4. RESULTS AND DISCUSSION

Experimental Results of Drop Distribution

Fig. 4 shows the droplet size distribution of diesel oil spray in the core and the outer

layer of the spray under ambient conditions. By intuition, it is assumed that the droplet

SMD in the inner space of the spray is much larger than that in the outer region, because

the shear force between the liquid drop and gas medium was small. To the contrary,

however, the experimental results indicate that the droplets in the outer layer are larger

than those in the inner as shown in Fig. 4. It can be found that there are dense droplets

near the core and sparse droplets near the outer layers of the sprays. At ambient

temperature, the size range near the core is about 3 to 30 /an, elsewhere diameters are up

to 60 fim in the outer region.

(A) outer region (B) Core region

Fig 4. Droplet distributions in a diesel oil spray, Pg = 1,2, MP^, Tg = 300 K, and t = 0.9 ms

362
When gas in the chamber is heated to 450 K, 540 K, 650 K and 750 K respectively,

the same phenomena are observed, i.e., large drops are distributed in the outer layer of the

spray. However, due to the strong heat and mass transfer processes between droplets and

gas in the outer layers, droplets in the outer region vaporize faster than those near the core. Consequently, the spatial droplet size distribution in the outer region tends to be more

uniform than that in the core region with increasing temperature, as shown in Fig. 5.

(A) T_ = 280 K (B) T_ = 520 K (C) T = 740 K

Fig 5. Droplet distributions in an outer region of a diesel oil spray at different gas temperatures,

p = 0.5 MPg and t = 0.4 ms

It was also observed that large droplets existing in outer layers form an umbrella

shape spray. It is because when liquid fuel is injected, the perturbation motion is vigorous

in the interface between the unbroken hollow cone and gas medium. As a result, the edges

of the spray is torn first, and then the whole spray. Droplets torn out at the edge are smaller than those in the cone surface of umbrella spray. On the other hand, because of

the effect of strong curly motion, small droplets in the periphery are drawn into the core of

the spray while the trajectory of large droplets changes slowly due to their greater inertia.

Nnmerical Studies on Droplet Distribution and Spray Development

Fig. 6 and 7 show the trajectories of two different groups of droplets injected at

different stage of spray development. The parameters of typical groups of droplets are summarized in Table 1. Dimensionless time r was used to describe the stages of spray

development, which was defined as the ratio of spray developing time to the total period of

injection. Since the atomization process is unsteady, the flow field of gas varies with time.

The numerical results indicate that small droplets are influenced greatly by turbulence.

For small droplets, velocities decelerate quite rapidly, resulting in conformity to the gas

velocity in a short time. Since the gas medium was quiescent before injection, later gas

motion was developed along with the droplet motion and spray development. When fuel is

injected into the chamber, the gas medium is entrained into the spray as shown in Fig. 2

.

The small droplets diverge their initial direction quickly, and are blown into the inner

region of spray due to the curly motion near the spray. The computational results agree

well with experimental results.

363
o Group 1

A Group 313

15

20 40 60 60 100 Axial Direction Imm)

Fig 6. Droplet trajectories of two different

groups injected at the early stage

(t = 0 ~ 0.12)

Group 833

Group 938

20 40 60 80 100 Axial Direction [mm]

120 140

Fig 7. Droplet trajectories of two different

groups injected at the middle stage

(t = 0.3 ~ 0.4)

Table l.Parameters of each group of droplets injected at different time

Group No. of Dimensionless Injection Location Dia. H^^

Droplets Time of Droplets C/^''^)

r (mm) z (mm)

1 0 Til oTo sTg i5o,285

2 4 X 10"^ 1.23 0.0 7.0 313,562

313 0.1248 1.08 0.0 51.4 786

833 0.3328 1.15 0.0 59.2 515

938 0.3748 1.10 0.0 7.3 272,344

Furthermore, the computational results account for the interaction of droplets

through their momentum exchange with the gaseous medium. Table 2 shows the spatial

location of two groups of droplets with time, which have the similar parameters. Droplets

groups 1 and 2 are injected consecutively, with At = 6 x 10 The droplets in group 1 were injected prior to the droplets in group 2. The latter moves faster than the former,

because the droplets of group 2 move in the wake of the droplets of group 1, therefore

reducing the drag force. The droplets in group 2, therefore, maintain a higher velocity, and

surpass the droplet 1. Fig. 9 is the droplet distribution at r = 0.24. As a spray develops, large droplets

are mainly distributed in the front and outer region of the spray due to their larger inertia.

It is observed that the results are in good agreement with experimental results.

364
Table 2.Spatial locations of two groups of droplets

Dimensionless

Time ( y )

Group 1 Group 2

r (mm) z (m.m) r (mm) z (mm)

0.0 1.21 0.0 1.23 0.0

0.075 4.18 11.57 7.1 24.0

0.25 5.06 19.53 4.12 48.3

0.333 3.1 47.7 3.74 84.31

0.41 2.2 116.1 3.54 136.5

25

£

E

c

o

o

a> b

CO

20-

15-

10

5

0

o

o

o •

o o "o

• f® O O o O O

O = • Dd =

Dd =

6-20

20- 30

30- 60

20 40 60 80 100

Axial Direction [mm]

120 140

Fig 8. Predicted droplet distributions at t = 0.24

5. CONCLUSIONS

Atomization features of pure diesel oil sprays have been studied. The following

conclusions were obtained:

1. For the pintle diesel injector, most of the small droplets are in the inner core of the

spray and droplet sizes are proportionally larger with an increasing radial distance

from the spray axis in the tested temperature range.

2. With increasing temperature, droplets in the outer layers will vaporize and their size

decreases because of stronger heat and mass transfer processes in the outer layers,

and in turn, the droplet size difference between in the boundary region and in the

core region becomes smaller.

3. The numerical results indicate that small droplets tend to follow the gas flow in a short period after injection, and the distribution of small droplets depends upon the

gas velocity field.

4. The computational results agree well with the experimental results, i.e., the droplet

size increases with the radial distance from the spray axis. In the spray, the large

droplets distribute in the front and outer region of sprays.

365
REFERENCES

1. William, A., "Combustion of Droplets of Liquid Fuels: A Review," Combustion k

Flame, Vol 21, pp. 1-31, (1973).

2. Varde, K.S., Popa, D.M. and Varde, L.K., "Spray Angle and Atomization in Diesel

Spray," SAE 841055, (1984).

3. Hikaru, K. Hideaki, T., Takeshi, S. and Fujimoto, H., "Investigation on the

Characteristics of Diesel Fuel Sprays," SAE 800968, (1980).

4. McDonell, V.G. and Samuelsen, G.S., "Application of Two—Component Phase

Doppler Interferometry to the Measurement of Particle Size, Mass Flux, and

Velocities in Two—Phase Flows," Twenty—Second Symposium (International) on

Combustion, pp. 1961-1971, (1988).

5. Keiya, N. Naoki, M. and Hiroyuki, H., "Holographic Measurement of Evaporating

Diesel Sprays under High Pressure and Temperature, Nippon Kikai Gakkai

Ronbunshu B Hen Vol. 52, No. 478, 2457-2463 (1986).

6. Wu, K.J., Coghe, A., Santavicca, D.A. and Bracco, F.V., "LDV Measurements of

Drop Velocity in Diesel-Type Sprays," AIAA Journal, Vol. 22, No. 9, pp.

1263-1270 (1984).

7. Ma, Y., Xin, J. and Choi, K.J., "Experimental Studies of Diesel Oil Spray and

Microexplosion of Water—in—Oil Emulsions by Using Optical Imaging Techniques,"

ILASS-AMERICAS, 4th Annual Conference, Hartford, CT (1990).

8. Arcoumanis, C, Whitelaw, J.H., and Wong, K.Y., Gaseous Simulation of

Diesel-Type Sprays in a Motored Engine, SAE 890793, 1989.

9. McMillian, O.J. and Ferziger, J.H., "Direct Testing of Subgrid Scale Models,"

AIAA Journal, Vol. 17, No. 12, pp. 1340-1346 (1979).

10. Travis, J.R., Harlow, F.H. and Amsden, A. A., Numerical Calculation of Two—Phase

Flows, Nucl. Sci. Engineering, Vol. 64, pp. 1-10 (1976).

11. Fan, Z.—Q., Zhang, B.—Z. and Ding, J.—M., "Prediction of Gas—Particle

Two—Phase Flows in Combustion," Vol. 3, Proceedings of International Symposium

Workshop on Particulate and Multiphase Processes and the 16th Annual Meeting of

the Fine Particle Society, Florida, (1985).

12. Xin, J., "Studies on Atomization and Combustion of Diesel Oil and Emulsion

Sprays," Master Thesis, Zhejiang University, China (1987).

366
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper37

A NEW ATOMIZATION MECHANISM FOR AIRLESS SPRAYING:

THE SUPERCRITICAL FLUID SPRAY PROCESS

K.A. Nielsen, C.W. Glancy, K.L. Hoy and K.M. Perry

Union Carbide Chemicals and Plastics Company, Inc. South Charleston, WV, U.S.A.

ABSTRACT

The supercritical fluid spray process uses supercritical fluids such as carbon dioxide and

nitrous oxide to replace the volatile organic solvent content in coating formulations. This not only

reduces solvent emissions to the atmosphere by 50 to 80 percent, while continuing to use existing

coating systems, it can also greatly improve airless spray technology by producing a new

atomization mechanism. Besides being an effective viscosity reducer, the dissolved supercritical

fluid creates an expansive force that overcomes liquid cohesion and surface tension forces that

oppose atomization. This produces vigorous atomization that remedies the deficiencies of the

airless spray process so that high quality coatings can be applied. It produces a fine droplet size

and a "feathered" spray that resembles an air spray system, but without having the high volume of

compressed air that deposits coating inefficiently. The coarse droplet size and "fishtail" spray

pattern characteristic of airless sprays are eliminated. The feathered spray fan can be produced in any width from narrow sprays to sprays that are much wider than current spray tip technology can

produce. This new technology can be applied to most coating systems and can be retrofitted into

existing spray operations. It is expected to be widely used industrially.

INTRODUCTION

Each year, American industry sprays 400 million gallons of coatings such as paints,

varnishes, lacquers, and enamels. They are applied by spray techniques such as air spray, airless

spray, air-assisted airless spray, and rotary atomizers, each of which has advantages and

disadvantages in how they apply energy to atomize the coating. Air spray uses a large volume of

compressed air, which gives the fine droplet size and uniform spray needed for high-quality

coatings but deposits coating inefficiently, so much coating is wasted. Conventional airless spray uses high fluid pressure, but produces coarse droplets and a non-uniform spray, which has limited it to applying low-quality coatings even though it is much more efficient than air spray. Air-assisted

airless spray is a combination of both, but uses less air volume and lower pressure, so it has

intermediate performance. Rotary atomizers can produce a fine droplet size and deposit coating

efficiently, but the spray is unfocused and the apparatus is expensive.

367
These spray techniques share the fundamental problem that coating formulations must have

a very low viscosity to produce the fine droplet size needed for high-quality coatings. This requires

a large amount of organic solvent, which is a major source of air pollution. Therefore, the

government has mandated that solvent emissions be reduced to low levels. Coating formulators

have responded by developing water-borne coatings and high-solids coatings, but these have

created new problems for applicators and have had limited success. Water-borne coatings must

contain enough organic coupling solvent to be compatible with water, so they often do not meet

regulatory limits. Quality has suffered from problems in evaporating the high water

content. Electrostatic spraying is difficult because water is highly conductive. High-solids coatings

retain many advantages of solvent-borne coatings, but they use reactive polymers with low

molecular weight, so they suffer from new problems such as running and sagging, cratering, poor

pot life, and curing that is slow or requires high temperatures.

The requirement of very low spray viscosity means that most of the solvent in coating

formulations is used just for viscosity reduction. Film formation actually requires a high coating

viscosity to prevent running and sagging on the substrate. Therefore, the solvents used for

viscosity reduction are fast-evaporating solvents that evaporate in the spray so that the viscosity of

the deposited coating is much higher. The coating requires a much smaller proportion of slowevaporating solvents to obtain proper film coalescence, leveling, and reflow to form a smooth

coating after application and during curing. This large viscosity increase that must occur during

spraying has so far been a spray technology barrier to reducing solvent usage while retaining good

coating performance. The previous methods of using less solvent to obtain low spray viscosity

have invariably reduced coating performance on the substrate as well.

We have developed new spray technology that solves this dilemma by using

environmentally compatible supercritical fluids such as carbon dioxide and nitrous oxide to replace

the volatile organic solvents used to obtain low atomization viscosity. Only the much smaller

amount of slow-evaporating solvent needed for film coalescence and leveling is used. This reduces

solvent emissions by up to 80 percent while using conventional polymers with high molecular

weights that give superior coating performance (1-2). Furthermore, we have discovered that

supercritical fluids not only reduce viscosity, they also produce vigorous atomization by a new

mechanism that remedies the defects of conventional airless spraying. Therefore, the supercritical

fluid spray resembles an air spray but without the high air volume, so high quality coatings can be

deposited efficiently.

CONVENTIONAL AIRLESS SPRAY MECHANISM

Airless spray uses a high pressure drop across the spray orifice to propel the coating

formulation through the orifice at high velocity. The conventional atomization mechanism is well

known (3). The coating exits the orifice as a liquid film that becomes unstable from shear induced

by its high velocity relative to the surrounding air. Waves grow in the liquid film, become unstable,

and break up into liquid filaments that likewise become unstable and break up into droplets.

Atomization occurs because cohesion and surface tension forces, which hold the liquid together,

are overcome by shear and fluid inertia forces, which break it apart. However, the cohesion and

surface tension forces are not entirely overcome and they profoundly affect the spray, particularly

for viscous coating formulations.

Conventional airless spray techniques are known to produce coarse droplets and defective

spray fans that limit their usefulness to applying low-quality coating films. Higher viscosity

increases viscous losses within the spray orifice, which lessens the energy available for

atomization, and it decreases shear intensity, which hinders the development of natural instabilities

in the expanding liquid film. This delays atomization so that large droplets are formed (4).

368
Fig. 1 Conventional airless "fishtail" spray Fig. 3 "Feathered" spray with carbon dioxide

Distance Along Coating Pattern Distance Along Coating Pattern

Deposited by Spray Deposited by Spray

Fig. 2 Two nneasured "fishtail" spray deposition Fig. 4 Measured "feathered" spray

patterns deposition pattern with carbon dioxide

The spray characteristically forms a "tailing" or "fishtail" spray pattern. An example is shown in the photograph in Figure 1 . More liquid gathers at the edges of the fan than in the

center, which produces coarsely atomized jets of coating. Sometimes the jets separate from the

spray and deposit separate bands of coating. At other times they thicken the edges so that more

coating is deposited at the top and bottom than in the center of the spray. Examples of measured

spray depositions of coating are illustrated in Figure 2. These deficiencies make it difficult to apply

a uniform coating.

369
SUPERCRITICAL FLUID SPRAY PROCESS

The supercritical fluid spray process uses gaseous fluids such as carbon dioxide and nitrous

oxide in their supercritical state to dilute coating formulations to spray viscosity instead of using

volatile organic solvents. Byproduct carbon dioxide from natural gas wells, fermentation plants,

and ammonia plants is preferred due to low cost and wide availability, but nitrous oxide might be

more environmentally acceptable, because it dissociates to nitrogen and oxygen, so that in effect

air is used as a solvent. Carbon dioxide becomes supercritical at temperatures and pressures above

its critical point, where liquid and gas properties become identical. The critical temperature, 31 C

(88 F), is just above room temperature and the critical pressure, 72 bars (1070 psi), is well within

the operating range of airless spray equipment.

For a wide variety of coating systems, we have demonstrated that 50 to 80 percent of the

organic solvent can be removed from coating formulations with equal or better coating

performance, because little or no change is made to the polymers and pigments. Only the solvent

blend is changed, eliminating the volatile solvents and retaining just enough slow-evaporating

solvent to obtain proper film coalescence and leveling (1). The solvent content of conventional

high-solids coatings can also be further reduced and their performance can be improved because

higher molecular weight polymers can be used (2).

By using the supercritical fluid spray process, conventional low-solids and high-solids

coatings can be reformulated at very high solids levels. These reformulated coatings are called

coating concentrates. They typically have viscosities from 800 to 3000 centipoise, although

viscosities up to 12,000 centipoise have been sprayed. These concentrate viscosities are much

higher than conventional formulations. Even conventional high-solids coatings have a low

application viscosity: 80 centipoise for clear coats and 35 centipoise for base coats. In general, air

spray and air-assisted airless spray require a low viscosity below about 100 centipoise and even

rotary atomizers have an upper viscosity limit below 500 centipoise.

Spray Conditions

The supercritical fluid is mixed with the coating concentrate at the time of spraying. The

spray solution usually contains 10 to 50 weight percent of dissolved carbon dioxide. The proper

concentration for a given coating application depends upon the solubility, viscosity reduction, solids

level, pigment content, and the spray temperature and pressure. The spray solution is heated

above the critical temperature to offset cooling that occurs as the carbon dioxide is released as a

free gas from the spray. The spray temperature is usually 40 to 70 C (100 to 160 F). The spray

solution is pressurized above the critical pressure to increase the carbon dioxide solubility. The

spray pressure is usually 80 to 110 bars (1200 to 1600 psi). Airless spray guns and nozzles are used that have been modified for the special properties of the spray solution.

Generally carbon dioxide solubility is not a limiting factor for spray application. Solubilities

usually range from 25 to 50 percent. Spraying is usually done below the solubility limit at the

spray conditions, so that the carbon dioxide is entirely dissolved. Excess carbon dioxide forms a second fluid phase, but usually no harm results if a small excess is present. A large excess can be

detrimental because solvent is extracted from the polymer solution into the excess carbon dioxide

phase. This can increase the viscosity and reduce the amount of solvent available for film

coalescence and leveling.

The large concentration of dissolved supercritical carbon dioxide produces a liquid spray

solution with markedly different properties. Conventionally sprayed coating solutions are incompressible liquids, but the supercritical fluid spray solution is highly compressible, that is, the

density changes markedly with changes in pressure. Furthermore, temperature and pressure

changes about the solubility limit can cause phase changes. Usually more carbon dioxide is present

than organic solvent, so polymer solvation can be different than in normal solutions. Because

370
carbon dioxide is a small molecule, it has much higher diffusivity and can penetrate polymer

configurations better than normal solvents. Supercritical carbon dioxide is an insulating solvent

that can give the spray solution the proper electrical resistivity for good electrostatic deposition.

Viscosity Reduction

An important function of the supercritical carbon dioxide is to reduce the viscosity of the

coating concentrate to a sprayable level. The spray viscosity is usually less than 50 centipoise,

often as low as one centipoise. It is a function of the polymer system, concentrate viscosity,

carbon dioxide concentration, and temperature. The viscosity is insensitive to spray pressure

provided the carbon dioxide remains dissolved. In most systems, less carbon dioxide than an equal

weight replacement of the replaced volatile solvent is needed to obtain the proper spray viscosity.

The viscosity reduction can be illustrated by using a medium molecular weight acrylic

polymer system. A concentrate with 65 percent polymer has a viscosity of 1000 centipoise at

room temperature. Heating it to a spray temperature of 50 C reduces the viscosity to 300

centipoise. Adding carbon dioxide to 28 percent concentration gives a spray viscosity below 30

centipoise. A concentrate with 75 percent polymer is very viscous at room temperature and the

viscosity is still 2000 centipoise at the spray temperature, but adding carbon dioxide gives a spray

viscosity below 40 centipoise.

SUPERCRITICAL FLUID SPRAY CHARACTERISTICS

Supercritical carbon dioxide not only functions as a good viscosity reducer, more

importantly, it can produce a vigorous new atomization mechanism that can remedy the

deficiencies of conventional airless spraying so that high quality coatings can be applied. The

supercritical fluid spray has the good characteristics of an air spray but eliminates the detrimental

large air volume produced by the compressed air. The vigorous decompressive atomization can

produce the fine droplet size needed for high quality coatings and a "feathered" spray that gives

uniform coating thickness. The coarse droplet size and "fishtail" spray of conventional airless

sprays are eliminated. The carbon dioxide escapes from the spray droplets as free gas, but the

volume is relatively small.

New Airless Atomization Mechanism

The vigorous decompressive atomization is believed to be produced by the dissolved

supercritical carbon dioxide suddenly becoming exceedingly super-saturated as the spray solution

leaves the nozzle and experiences a sudden and large drop in pressure. This creates a very large

driving force for gasification of the carbon dioxide, which overwhelms the cohesion, surface

tension, and viscous forces that oppose atomization and normally bind the fluid flow together into a

fishtail type of spray.

A different atomization mechanism is evident because atomization occurs right at the spray

orifice instead of away from it. Atomization is no longer due to break-up of a liquid film from shear

with the surrounding air. Instead it is due to the expansive forces of the compressible spray

solution created by the large concentration of dissolved supercritical fluid. Therefore, no liquid film

is visible coming out of the nozzle. Furthermore, because the spray is no longer bound by cohesion

and surface tension forces, it leaves the nozzle at a much wider angle than normal airless sprays

and produces a feathered spray with tapered edges like an air spray. This produces a rounded

parabolic-shaped spray fan instead of the sharp angular fans typical of conventional airless sprays.

An example of the feathered spray is shown in the photograph in Figure 3.

The spray pattern deposits coating uniformly in a wide central portion and deposits

progressively less coating at the tapered edges, as illustrated in the measured deposition profile in Figure 4. This feathering is particularly desirable for overlapping adjacent layers of sprayed coating

to produce a coating film of uniform thickness. This is one of the principle reasons why air sprays

371
are used instead of airless sprays to apply high-quality coatings.

The feathered spray pattern is characteristically obtained when the supercritical carbon

dioxide concentration in the spray solution exceeds a transition-point concentration for the given

spray conditions. With no carbon dioxide, the binding forces of cohesion, surface tension, and

viscosity in the incompressible spray solution produce a typical fishtail spray pattern with very poor

atomization. At carbon dioxide concentrations below the transition-point, the binding force

exceeds the expansive force of the carbon dioxide, so a fishtail spray pattern persists, but it becomes more uniform, the spray fan becomes wider, the visible liquid film recedes towards the

orifice, and the spray solution becomes compressible as the concentration increases from zero. At

the transition-point concentration, the expansive force equals the binding force, so neither controls

the spray pattern. The visible liquid film has disappeared and atomization is occurring at the spray

orifice. Surprisingly, as the concentration increases through the transition point, the angular flat

fishtail spray fan first contracts into a narrow transitional oval or flared non-planar spray and then

greatly expands into a much wider parabolic flat feathered spray fan. The transition can be seen

not only in the shape of the spray but also in the greatly improved atomization and deposition of

the coating. The droplet size becomes much smaller, which shows that the binding force is completely overcome by the expansive force of the carbon dioxide, and coating quality is greatly

improved. At carbon dioxide concentrations above the transition point, the spray fan is fully

feathered, much wider and thicker, and exits the spray orifice at a much greater angle. Higher

concentration further decreases the particle size, increases the fan width, and makes the spray

solution more highly compressible, which affects the spray rate.

Laser light scattering measurements and comparative spray tests show that the vigorous

atomization can produce fine droplets that are in the same size range as air spray systems (20 to

50 microns) instead of the coarse droplets produced by normal airless sprays (70 to 150 microns).

For a properly formulated coating, the droplet size range and distribution are ideal for minimizing

orange peel and other surface defects commonly associated with spray application.

This fine particle size provides ample surface area for the dissolved carbon dioxide to very

rapidly diffuse from the droplets within a short distance from the spray nozzle. Therefore, the

coating is essentially free of carbon dioxide before it is deposited onto the substrate. Ideally,

coatings sprayed conventionally and with supercritical carbon dioxide have the same composition

when they are deposited so that coating performance is the same but much less solvent is emitted

to the atmosphere.

The released carbon dioxide gas forms a shroud that tends to displace air, moisture, and

dust from the spray. The lower oxygen level and absence of volatile solvents significantly lowers

spray flammability. The displacement of moisture lowers the local dew point so that moisture

absorption is less likely in humid environments. The drier spray also helps droplets retain their

electric charge during electrostatic spraying.

Spray Temperature Profile

The rapid release of carbon dioxide from solution is evident in the very rapid cooling of the

heated spray that occurs within a very short distance from the spray nozzle. Figure 5 compares

temperature profiles for conventional heated airless sprays and a supercritical fluid spray that is heated to a temperature that compensates for expansive cooling. The conventional spray cools to

ambient temperature within 15 centimeters from the nozzle because the solvent evaporates and

ambient air is entrained into the spray. Therefore the coating is deposited at ambient temperature.

The supercritical fluid spray typically cools to ambient temperature within about two centimeters of

the nozzle. In some cases the spray might briefly subcool, but the entrainment of ambient air ensures that the coating is also deposited at ambient temperature. The amount of cooling depends

upon the carbon dioxide concentration and the spray temperature.

372
60

oo

LU

cc 40

IXI Q.

LU

I- 20

CONVENTIONAL SPRAY

SUPERCRITICAL FLUID SPRAY

J I : I I L—i ' '

'

AMBIENT

0 10 20 30

SPRAY DISTANCE. CENTIMETERS

Fig. 5 Temperature profile of spray

The rapid cooling is beneficial because it reduces solvent evaporation in the spray, so little solvent is lost from a properly formulated coating. Therefore solvent is used efficiently and worker

exposure to solvent vapors in the spray booth is reduced to a low level. Measurements for one

coating system showed that the solids content of the deposited coating was only about two

percent higher than the solids content of the coating concentrate. Therefore, the viscosities of the

coating concentrate and the deposited coating should be similar.

Spray Fan Width

One manifestation of the expansive force of the supercritical carbon dioxide is that the

feathered spray has a much greater fan width than normal airless sprays. The fan width is typically

50 to 1 50 percent greater, and in some cases can be 300 percent greater, than the fan width

rating of the airless spray tip, depending upon the nozzle type, carbon dioxide concentration,

coating formulation, and spray conditions. In particular, much larger fan widths can now be

produced for small orifice sizes than conventional spray tip technology can produce. Generally the

fan width can be altered by varying the carbon dioxide concentration.

Although the spray leaves the nozzle at a much wider angle and the fan width is much

wider than normal airless sprays, the fan width can be indexed to give any fan width from narrow

to very wide by changing the fan width rating of the airless spray tip, as is normally done.

Examples of fan widths measured with and without carbon dioxide are illustrated below for two

coatings, using dome-style spray tips with a 0.23-mm (9-mil) orifice size and measuring the fan

width 30 centimeters (12 inches) from the spray tip.

Fan Width Rating

of Spray Tip

Measured

Widths Without

Carbon Dioxide

Measured Widths

With Carbon Dioxide

Coating 1 Coatinq 2

5 cm 5 cm 10 cm 20 cm

10 cm 10 cm 18 cm 28 cm

15 cm 1 5 cm 28 cm 36 cm

20 cm 20 cm 33 cm 41 cm

373
In general, with the supercritical fluid spray process, cross-cut type airless spray tips produce larger

fan widths than dome-style spray tips with the same fan width rating. For example, a cross-cut

type spray tip with a 0.23-mm (9-mil) orifice size and a fan width rating of 15 centimeters

(6 inches) can produce a fan width of 61 centimeters (24 inches).

CONCLUSIONS

In addition to eliminating up to 80 percent of the organic solvent used to apply coatings,

the supercritical fluid spray process greatly improves airless spray technology by producing a new

atomization mechanism that gives the fine droplet size and feathered spray fan required to produce

high quality coatings. Therefore, we believe the process will be used to reduce solvent emissions

not only in current airless spray applications but also where air spray, air-assisted airless spray, and

rotary atomizers are currently used. Some of the anticipated end uses are automotive topcoats and

component finishing, automotive refinish, wood and metal furniture, appliances, pails and drums,

aerospace, machinery and equipment, trucks and buses, and structural steel.

With continued development of new coatings we expect solvent usage to decrease to very

low levels. Ultimately the supercritical fluid spray process will be able to apply coatings that are the liquid analog of powder coatings, but it will do so with easier implementation, greater

versatility, and better coating performance.

REFERENCES

1. Nielsen, K. A., Busby, D. C, Glancy, C. W., Hoy, K. L., Kuo, A. C, and Lee, C,

"Supercritical Fluid Spray Application Technology: A Pollution Prevention Technology for

the Future" in Proceedings of the Seventeenth Water-Borne and Higher-Solids Coatings

Symposium (R. F. Storey and S. F. Thames, Eds.), pp. 218-239, New Orleans, LA (1990).

2. Nielsen, K. A., Busby, D. C, Glancy, C. W., Hoy, K. L., Kuo, A. C, and Lee, C,

"Application of High Solids Coatings Using Supercritical Fluids", Proceedings Division of

Polymeric Materials: Science and Engineering . Vol. 63, pp. 996-999, American Chemical

Society, Washington D.C. (1990).

3. Dombroski, N. and Johns, W. R., "The Aerodynamic Instability and Disintegration of

Viscous Liquid Sheets", Chemical Engineering Science, Vol. 18, pp. 203-214 (1963).

4. Lefebvre, A. H. and Senser, D. W., "Research Unravels Spray Mysteries", Industrial

Finishing, Vol. 66, No. 6, pp. 16-20 (1990).

374
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 38

DROPLET DIAMETER, FLUX AND TOTAL CURRENT MEASUREMENTS

IN AN ELECTROHYDRODYNAMIC SPRAY

P.F. Dunn and S.R. Snarski

Particle Dynamics Laboratory

Department of Aerospace and Mechanical Engineering

University of Notre Dame

Notre Dame, IN, U.S.A.

ABSTRACT

The measured diameter and flux distributions and total current of ethanol droplets generated

electrohydrodynamicaliy from a single hypodermic needle at two different spray charge densities are presented and discussed. At the lower spray charge density, the droplet diameter distribution having the

smallest mode diameter occurs at the spray centerline, with a diminishing number of droplets and an

increasing modal diameter occurring in the lateral direction toward the periphery of the spray. At the

higher charge density, these droplet diameter distributions become more narrow and approach lognormal

distributions. The droplet fluxes are maximum at the spray centerline and decrease appreciably toward

its periphery. The droplet flux distributions are used to construct droplet charge flux distributions by

assuming various models of the droplet charging process. The resultant charge fluxes are logarithmically

related to lateral position, irregardless of the droplet charging model assumed. The integrals of the

various charge flux profiles are compared to the measured total current. This approach reveals the best

agreement between measured and predicted total droplet currents when an equilibrium end-state process

for droplet charging is assumed.

INTRODUCTION

With the recent development of advanced in situ laser diagnostic systems, it is now possible to obtain

non-intrusive measurements of the velocity and/or diameter of droplets within electrohydrodynamic

(EHD) sprays to a level of detail that was not possible in earlier studies. This paper presents the results

of such laser diagnostic measurements of droplet diameter and flux within a single EHD spray. The

present work further examines the process of droplet charging by using the various models that relate

droplet diameter and charge to predict the total droplet current, which also is measured. As a consequence, the model that best describes the charging process for these experiments can be identified.

The major findings from this study, when considered in conjunction with those reported recently by Dunn

and Snarski [1] conducted under similar conditions using two parallel needles, support that droplet sprays

of ethanol obtained at the subject levels of spray charge density closely approach or already have

reached their equilibrium end-state, as described by Kelly [2].

375
EXPERIMENTS

In the present experiments, a fine spray of droplets in the approximate diameter range of 1 - 50 |j.m

were generated electrohydrodynamically by supplying ethanol at a fixed mass flow rate (4.5 mg/s) from a syringe pump and applying a high positive voltage to a stainless steel hypodermic needle (216 |j.m I.D., 406 |j.m O.D., 1 .905 cm length) located 38 cm above an electrically grounded droplet-collection funnel (see

Figure 1). Droplet diameter and flux distributions were obtained using a Phase Doppler Particle Analyzer

(PDPA) (see Bachalo and Houser [3]) at two axial (z) positions below the needle's tip (15 and 30 mm),

at approximately twenty lateral (x) positions for each axial position and at four applied voltages (15, 20,

25 and 30 kV). Total droplet currents were measured as the collected droplet current to ground through an electrometer. The resultant spray charge densities for these conditions were 12, 42, 106 and 210

C/m3, respectively. A description of the experiment's main components and operating procedure is presented in further detail elsewhere [1].

PDPA Receiver

Linear Traverse

X [

Funnel

Electronic Balance I OO'OO I

I I —

Syringe Pump

High Voltage

Supply

PDPA Transmitter

Electrometer —

Figure 1 . Schematic of the experiment.

Only the measurements acquired at the two highest spray charge densities and first axial position are presented and discussed herein. These essentially constitute those measurements in the Fine Spray

Mode (see Snarski [4] for a discussion of the various BHD regimes encountered in such experiments) in which micrometer-size droplets are produced from short, brush-like filaments emanating directly from the

tip of the needle. This situation is achieved at relatively high spray charge densities (in the range of -100-

200 C/m3), at which the electric field strength at the tip of the needle is -lO^ V/m.

Estimated uncertainties in the measured quantities were 10% for the volumetric flow rate and less

than 1% for the measured voltage. The uncertainty for the probe volume location also was less than 1%,

376
while the typical uncertainty in the measured diameter was 4%, as estimated by Bachalo and Houser

[3]. The overall range of uncertainty in the measured diameter for those measured in the present

experiments was from 2 to 16%, based upon the PDPA droplet sizing calibration studies of O'Hern etal.

[5]. The uncertainty in the current measurements was from 1% at the lowest voltage, where all the

spray was collected by the funnel, to approximately 30% at the highest voltage, where the spray width

exceeded that of the funnel. To compensate for the large uncertainties in current measurement at the

higher voltages, the values of the currents reported herein were obtained by extrapolation of those

measured at the lower voltages and at the higher voltages but at different needle-to-funnel distances.

RESULTS AND DISCUSSION

The droplet diameter distributions measured at the 15 mm axial position for the spray charge

densities of 106 and 210 C/m3 are shown in Figure 2 for three lateral positions. These lateral positions,

representing typical axial centerline (x=0 mm), near-centerline (x=8 mm) and peripheral (x= 32 or 36

mm) distributions, adequately display the evolution of distributions that occurred across the span of the

axially symmetric spray. At the lower charge density, the distributions at the centerline and peripheral

locations are reasonably similar, but both are somewhat different than that at the near-centerline

location. The mode of the distribution (the diameter of maximum count) is smallest at the centerline. At

the higher charge density, the distributions become comparatively more disparate and evolve toward a

lognormal distribution at the centerline. A comparison of the distributions between the two spray charge

densities reveals a progressive increase with increasing spray charge density in the number of smaller

droplets occurring at the centerline and at the periphery. This is especially evident at the axial centerline.

This trend is like that obtained previously by Snarski [4] in similar experiments using two needles.

Individual droplet size distributions can be modelled by various empirical fits (see Lefebvre [6]). For

the present study, the Rosin-Rammier distribution provided excellent agreement with the measurements.

In particular, the measured and predicted volume mean diameters agreed to within approximately 5% for

all cases and to within approximately 2% when averaged over the lateral span of the spray. A

comparison between measured and predicted % droplet volume and count distributions is given in Figure 3. The resultant values for the two Rosin-Rammler fit parameters (the size dispersion coefficient, q, and

the diameter of 63.2% cumulative liquid volume, X) at a fixed spray charge density were minimum at the

centerline and increased by approximately 10 to 20% across the lateral span of the spray. When

averaged across the spray, these values ± their standard deviations were qavg = 3.65 ± 0.27 and Xavg =

26.6 ±1.80 |im for 106 C/m3 and qavg = 3.52 ± 0.18 and Xgvg = 22.5 ±1.90 |im for 210 C/m3. These

resultant values are within the bounds of those obtained for most sprays [6] and are like those found in Snarski's experiments [4]. Also, the parameters q and X were minimum at the spray centerline at each

charge density, implying a diameter distribution of smaller droplets with slightly more dispersion than at other locations. This occurred even more so at the higher spray charge density.

At the higher charge density, however, a progressively larger number of smaller diameter droplets

comprised a majority of the droplet population. For this situation, the Rosin-Rammler empirical fit

inherently did not predict well the lower diameter end of the distribution. Here, the measured diameter

distribution was better described in this region by a lognormal fit, as evidenced by the precision of the

lognormal fit of the count distribution shown in Figure 3. Based upon this trend, it is plausible to expect

that at even higher charge densities a normal distribution will provide the best fit, as predicted by Kelly

[2], if the droplets reach an equilibrium end-state. Such an evolution to a normal distribution would signify a reduction in the complexity of or number of parameters affecting the droplet population [7]. Further insight into the dynamics of the droplets within the spray can be gained by examining the

measured droplet flux distributions. These are presented in Figure 4 for the same conditions as those in

Figure 2. At a fixed spray charge density, the greatest droplet flux occurred at the centerline where the

droplet population was comprised of a majority of small droplets. In the lateral direction toward the

periphery of the spray, the droplet flux decreased appreciably, eventually by 2 to 3 orders of magnitude

377
10

8

O centerline (x=Omm) ; 1 06 C/m^ • centerline (x=Omm); 210 C/m3

^

o

2 -O

o

o

o

10

8

0

10

8

A near-cl (x=8mm) ; 1 06 C/m^

A near-cl (x=8nnm); 210 C/m^

periphery (x=32mm) ; 1 06 C/rrfi

periphery (x=36mm) ; 21 0 C/m^

10 20 30

Diameter, \im

Figure 2. Droplet diameter distributions at the 15 mm axial location and spray charge densities

of 106 and210C/m3.

378
Diameter (pm)

Figure 3. Rosin-Rammler and lognormal fit comparison at the spray centerline, 15 mm axial location and

spray charge density of 210 C/m^.

at the periphery. At the higher charge density, for all lateral positions, the maximum value of the droplet

flux occurred at smaller diameters. This is consistent with the photographic observations reported by

Snarski [4], in which the length and diameter of the brush-like filaments extending from the tip of the

needle decrease with increasing charge density and produce a larger number of smaller droplets.

The measured droplet flux distributions also can be used in conjunction with various models that

relate droplet charge to diameter to construct charge flux distributions. This is accomplished by

incorporating the appropriate droplet charge-to-diameter relation into the droplet flux distribution, directly

yielding the charge flux distribution. In the present study, two models were examined. In one, droplets

were assumed to be charged to their maximum (Rayleigh) limit where charge ~ diameter3/2, and in the

379
periphery {x=32mm); 106 C/m3

periphery {x=36mm); 210 C/m3

0 10 20 30 40

Diameter, urn

Figure 4. Droplet flux distributions at the 15 mm axial location and spray charge densities

of 106 and210C/m3.

380
other, to an equilibrium end-state where charge - diameter [2]. Other droplet charging models, such as

those yielding minimum energy and most probable distributions, were not considered. This was because

the Rayleigh model describes an upper limit and Kelly's model recently has been shown in two independent

experiments [2,7] of EHD sprays to best describe the measured velocity-diameter correlations.

The resultant droplet charge flux profiles at both charge densities are displayed in Figure 5 for each

of the two models. Implicit in determining these profiles is the choice of a diameter that is representative

of the diameter distribution at each lateral position. Here, for the Rayleigh model, the PDPA measured

diameters were used directly and, for Kelly's model, several representative diameters were examined

(dmode. dio. CI20. ^30 and d32 - see Lefebvre [6]). The linear average diameter, dio, gave the best

agreement between experiment and theory in terms of total current measurement. At both spray charge

densities, each model predicted that the maximum charge flux was at the centerline. The Rayleigh model, as anticipated, predicted the highest charge flux. Comparison between the two charge density cases

further showed that the droplet charge flux distribution with increasing spray charge density became

logarithmically related to lateral position. The least-squares fits of the data yielded high correlation

coefficient values of 0.98 and 0.99 at the higher charge density and 0.94 and 0.94 at the lower charge

density for the Rayleigh and Kelly models, respectively. All of these values corresponded to confidence

levels greater than 99.9% when considering the number of lateral positions examined. These findings

imply, irregardless of the droplet charging model assumed, that the droplet charge flux distribution is logarithmically related to lateral position.

Ql 1 1 > 1 ^ 1 . 1 0 10 20 30 40

Position, mm

Figure 5. Droplet charge flux profiles for Rayleigh's and Kelly's models at the 15 mm axial location and

spray charge densities of 106 and 210 C/m^,

381
The model most descriptive of the present experiments can be identified by numerically integrating

these profiles across the lateral span of the spray to yield values that can be compared to the measured

total droplet current. This comparison revealed very good agreement (within the experimental

uncertainty) with Kelly's theory (to within approximately 5% for the lower charge density [0.63 |a.a

predicted versus 0.60 |j.a measured] and to within approximately 20% for the higher charge density [0.94

|ia predicted versus 1.20 |j,a measured]). The total droplet currents predicted from Rayleigh's model were

approximately twice the magnitude of the measured values (1 .27 ^a and 2.08 )j.a for the lower and higher

charge densities, respectively). These results further support that Kelly's theory best describes the

present experimental results.

CONCLUSIONS

The measurements presented herein support that droplet sprays of ethanol obtained at the subject

levels of spray charge density closely approach or may have already reached their equilibrium end-state.

This conclusion was drawn from the observed evolution of the droplet diameter and flux distributions with

increasing spray charge density. Further evidence was acquired by determining that the best agreement

between measured and predicted total droplet currents was achieved by assuming an equilibrium endstate process for droplet charging.

ACKNOWLEDGEMENTS

We gratefully acknowledge the technical advice of Drs. V.J. Novick and E.D. Doss within the

Engineering Physics Division of Argonne National Laboratory. This work was performed under MIPR No.

FY761 68800343 between the University of Chicago and the Department of Defense and under Contract

No. 62622401 between Argonne National Laboratory and the University of Notre Dame.

REFERENCES

1. Dunn, P.F. and Snarski, S.R., "Velocity Component and Diameter Distribution Characteristics of

Droplets Within Two Interacting Electrohydrodynamic Sprays," Physics of Fluids A. Vol. 3, No.3,

pp. 492-494,1991.

2. Kelly, A.J., "Electrostatic Metallic Spray Theory," Journal of Applied Physics . Vol. 47, No. 12, pp.

5264-5271,1976.

3. Bachalo, W.D. and Houser, M.J., Optical Engineering. Vol. 23, No. 5, pp.583-590, 1984.

4. Snarski, S.R., "The Interaction of Electrohydrodynamically Generated Liquid Droplets," Masters

Thesis, University of Notre Dame, Notre Dame, IN, 1988.

5. O'Hern, T.J., Rader, D. J., and Ceman, D.L., "Droplet Sizing Calibration of the Phase Doppler Particle

Analyzer," Annual Meeting of the American Association for Aerosol Research, Reno, NV, 1989.

6. Lefebvre, A.H., Atomization and Sprays, Hemisphere Publishing Corporation, New York, 1989.

7. West, B.J. and Shiesinger, M., "The Noise in Natural Phenomena," American Scientist. Vol. 78, pp.

40-45,1990.

8. Gomez, A. and Tang, K., "Characterization of Liquid Fuel Electrostatic Sprays," Fall 1990 Meeting

of the Western States Section of the Combustion Institute.

382
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 39

INDUCED ATOMIZATION

S.P. Lin and D.R. Woods

Mechanical and Aeronautical Engineering Department

Clarkson University

Potsdam, NY, U.S.A.

ABSTRACT

An intact liquid jet emanating from a nozzle is subjected to a periodic external forcing

in the axial direction. It is discovered that the jet may be induced to atomize and produce

droplets of diameters much smaller than the jet diameter. The induced atomization occurs only

in widely separated discrete frequency bands. The diameter of the droplets produced in these

bands decrease with increasing frequency. The locations of the discrete bands for the induced

atomization depends on the relevant flow parameters.

INTRODUCTION

It is well known since the time of Rayleigh (1) and Plateau (2) that a liquid jet emanated

from a nozzle may break up into droplets of diameter comparable to the diameter of the liquid

jet. The mechanism of the breakup is due to the capillary pinching (3). Taylor (4) demonstrated

that at a relatively high jet velocity there is another mode of breakup which results in a formation

of a spray of fine droplets of diameters much smaller than the jet diameter. Lin & Kang (5), and

Lin & Lian (6) showed that formation of droplets of this scale by spontaneous atomization of a

liquid jet without external excitation can be achieved only in the parameter range Q » We,

where Q = P2/P1 is the ratio of the density of the ambient gas to that of the liquid, We is the

Weber number We = S/pj U^ a, S being the surface tension, U the jet speed and a is the nozzle

inner radius. By use of the energy equation and neglecting the gas viscosity Lin and Creighton

(7) showed that while the mechanism of the jet breakup by the Rayleigh mode is capillary

pinching, the mechanism of atomization is the high frequency pressure fluctuation at the liquidgas interface. Taking into account the effect of gas viscosity, Lin & Ibrahim (8) showed that the

interfacial shear stress fluctuation may be of the same order of magnitude as the pressure

fluctuation in some parameter range. A similar energy analysis for a coaxial core annular flow

was reported by Chen, Bai & Joseph (9). These results suggest that atomization may be induced

even in the range We ~ Q by a high frequency oscillation of the nozzle tip in the axial direction.

This is found to be the case in the present experiments. However, unexpectedly we discover that

the atomization can be induced only in widely separated narrow bands of frequency.

EXPERIMENT

Figure 1 displays the general layout of the experimental equipment. A flask filled with

a test liquid is connected from it upper neck to a flexible tubing. The tubing is connected

through a pressure gauge and a valve to a compressed air line. A nozzle is connected to the flask

383
Sine Wave Genenior OsciUoscope

Motor ComroUer

Processor

r

\

IBM PC/AT

[

Fig. 1 Experimental Layout

through its lower open protrusion. The liquid discharge rate through the nozzle is regulated by

use of the pressure gauge which controls the pressure exerted at the air-liquid interface in the

flask. The flask is mounted on a frame which can be traversed vertically as well as horizontally.

The liquid discharged through the nozzle is collected in a beaker under the nozzle. The

cylindrical nozzle tip is made of Vemitron piezoelectric ceramic. Its inner diameter and length

are respectively 0.05 cm and 1.905 cm. The cylindrical wall thickness is 0.05 cm. The axis of

the nozzle tip is aligned in a vertical direction. The electrodes at the upper end of the

piezoelectric cylinder are connected to a function generator, through an oscilloscope and a

amplifier (Burr-Brown 3854 JM). The amplifier is required to raise the voltage to approximately

100 Volts which is necessary to vibrate the 1.905 cm long piezoelectric piece by as small an

amplitude as 1 |im in the axial direction. The function generator supplies sinusoidal voltages on

the order of 10 Volts peak to peak at frequencies between 1 Hz and 100 KHz. The signal

supplied to the nozzle is monitored on the oscilloscope. The size and the axial velocity

384
distributions of the atomized droplets at a given spatial position are measured non-intrusively by

use of a Phase Doppler Particle Analyzer (PDPA, Aerometrics). The PDPA system consists of

the laser beam transmitter (XMT-1100), a signal receiver (RCV-2100), a motor controller (MCB700-1) Phase/Doppler processor and a data processor. Two transmitted laser beams of equal

intensity intersect at a specified spatial position where the measurement is to be taken. A droplet

passing through this position scatters the transmitted light and produces an interference fringe

pattern. The fringe spacing is related to the drop size by the Mie equation (10). The image of

the fringe pattern appears to move past the receiver at the Doppler difference frequency. Each

of the three detectors in the receiver sees a similar but out of phase Doppler burst. The phase

shift is related to the drop size and the Doppler shift is related to the droplet velocity. The

conversion of the phase and frequency shifts to the particle size and velocity is achieved by the

Phase/Doppler processor (PDP-3100). The statistics of the size and velocity distributions at the

spatial position over a sufficiently long period of time is then obtained by use of an IBM PC/AT.

The detail of the principle of a PDPA system can be found int he paper by Bachalo and Houser

(10). Pure alcohol and 70% -30% by volume of alcohol-water mixture were used as the test

liquids.

The experimental procedure is relatively straightforward. First the receiver is aligned to

receive the forward scattering light making a fixed angle of 30° with the transmitted laser beam.

The flask is traversed to a desired position. The valve of the compressed air line is then opened

and the pressure gauge is set to a preassigned pressure depending on the discharge rate desired.

The function generator is then tumed on, and the frequency is gradually raised until atomization

takes place. The size and the velocity distributions at the intersection of the transmitted beams

are measured with the Aerometric PDPA system.

RESULTS

The results will be presented in terms of the following relevant dimensionaless groups:

Weber number. We = S/pj U a,

Reynolds number. Re = U a/Vj,

Froude number, Fr = U /ga,

Strouhal number, St = U/coa,

amplitude ratio, A = 5/a,

density ratio, Q = P2/Pi»

viscosity ratio, M = V2^V

where the subscript 1 and 2 stand for the liquid and gaseous phases respectively, v is the

kinematic viscosity, g is the gravitational acceleration, and 6 and co are respectively the amplitude

and the frequency of the nozzle oscillation. The amplitude 5 is related to the applied voltage V

385
and the piezoelectric constant d^j (-274 x 10" nVVolt) by

5 = V-L.d3i/t ,

where L and t are respectively the length and thickness of the piezoelectric cylinder. The

amplitude 5 in the present experiment is approximately 1 |im according to the above equation.

Hence the parameter 5/a is fixed to be 4 x 10'^. The values of Q and M for the test fluids are given in Table 1.

Table 1 - Physical Properties of Fluids

Fluid Viscosity Surface tension Density lO^Q M

(% Alcohol - % H^O) (lO'^m^/s) (10"^ N/m) (Kg/m3)

100 - 0 1.45 2.30 789.0 1.53 10.42

70 - 30 2.72 2.71 851.7 1.42 5.52

When the frequency of the function generator is increased gradually from zero, the jet

remains intact until the fu"st critical frequency is reached. When the first critical frequency is

exceeded, the jet suddenly atomizes to form the fine spray. As the frequency is raised further,

the jet continues to atomize in a narrow frequency band. When this band is exceeded, the jet

returns to its intact configuration. Before the maximum function generator frequency of 100 KHz

is reached, three such narrow bands of atomization are encountered. The location of the

frequency bands depend on the relevant flow parameters. The results obtained with pure ethyl

alcohol are presented in Figure 2. Measurements were taken at a point 3.0 mm radially and 5.0

mm axially from the nozzle exit. In each figure the values of A, Q, and M are fixed. The

measured dimensionaless average droplet diameter D = d/(2a), d being the average diameter, are plotted against St in the figure for the set of We, Re, Fr specified in the figure.

D

0.060

1 f 1 1

•

1

1 —f—

•

0.0*0 • • •

•

0 0:30

•

•

0030

•

o nin 1 1 ..J. L. 1, _ 1 X. . —t 1 0.10 OJO 0.30 0.40 0.50 0.60 s

t

Fig. 2 Selective Atomization

Re = 500, We = 0.01383, Fr = 3433, Q = 0.001525, M = 10.42, A = 0.004

386
ACKNOWLEDGEMENT

This work was supported in part by grant no. DAAL03-89-K-0179 of ARO, grant no. MSM-88 17372 of NSF and a New York State Science Foundation Grant. The computation was

carried out with the computer facility at Clarkson University and with the Cornell National

Computer facility, which is funded by the NSF, the State of New York, and IBM Corporation.

REFERENCES

1. Rayleigh, Lord, "On the Instability of Jets", London Math Soc, 10, pp. 361-371 (1879).

2. Plateau, J., Statique Experimentale Et Theorique Pes Liquids Soumix Aus Seules Forces

Moleicularies, Canthier Vallaris (1873).

3. Chandrasekhar, S., Hydrodynamic and Hydromagnetic Stability, Oxford U. Press, Oxford

(1961).

4. Taylor, G.I., The Scientific Papers of G.I. Taylor, Cambridge U. Press, London, Vol. 3

(1963).

5. Lin, S.P. and Kang, D.J., "Atomization of a Liquid Jet", Phys. Fluids, 30, pp. 2000-2006

(1987).

6. Lin, S.P. and Lian, Z.W., "Mechanism of Atomization", AIAA, 28, pp. 120-126 (1990).

7. Lin, S.P. and Creighton, B., "Energy Budget in Atomization", Aerosol Science and Tech.. 12, pp. 630-636 (1990).

8. Lin S.P. and Ibrahim, E.A., "Instability of a Viscous Jet Surrounded by a Viscous Gas",

J. Fluid Mech., 218, pp. 641-658 (1990).

9. Chen, K., Bai, R, and Joseph, D.D., "Lubricated Pipelining HI", J. Fluid Mech.. 214, pp.

251-266 (1990).

10. Bachalo, W.D. and Houser, M.J., "Phase/Doppler spray Analyzer for Simultaneous

Measurements of Drop Size and Velocity Distributions", Optical Eng., 23, pp. 583-590

(1990).

387
1\IIII1iIIiiijjII1ii I '\

'' tI11I11I
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 40

PRODUCTION OF UNIFORMLY-SIZED DUAL CONCENTRIC DROPLETS

FROM COAXIAL SMOOTH JET UNDER APPLIED AC ELECTRIC FIELD

T. Sakai, M. Sadakata, M. Sato, K. Kimura and M. Saito

Department of Biological and Chemical Engineering

Gunma University

Kiryu, Gunma, Japan

ABSTRACT

Hie production of uniformly-sized droplets by means of a mechanical vibration method and an electrical

method has been studied. The relationship between these two methods at the limits of the uniform region are aimed

to establish. The upper and the lower limits of the frequency of a uniform droplet region (synchronous region) for

the mechanical vibration method and the AC electric field method were the same. From the analogy of the surface

wave propagation caused by mechanical and electrical vibrations, it is considered that the AC electric field may cause

an initial disturbance or displacement on the electrically conductive liquid surface. Thus the surface wave growth

along the liquid column surface will result in synchronous disintegration of the smooth Uquid jet

The generation of concentric dual droplets by the AC electric field method and surface wave propagation

rates on co-axial dual liquid columns are discussed. It is proposed that the present method may provide a new

technique for the production of fine uniform droplets or particles, even for non electrically conductive liquids.

INTRODUCTION

Research pe^rs on the production of uniform droplets have essentially describe two methods. One is the

vibration method vMch includes acoustic, mechanical and liquid pressure vibration and the other method is the

application of electrostatic field. Reports on these methods have been published [1 - 14], however, the relationship

between the vibreition method and the electrostatic method has not clarified.

In this pe^r, at first, the disintegration mechanisms during uniform droplet production of both methods are

compared and analyzed. The results are then appUed to produce uniform sized droplets by Explication of high AC

voltages. The second purpose of this work is to produce smaller uniform droplets of non or poorly electrically

conductive liquids using electrostatic method, by applying AC voltages on a smooth dual coaxial liquid column

covered with conductive liquid. This method would provide more convenient process to produce finer uniformlysized droplets with lower injection pressure particularly for high viscous non conductive Uquids when dual

convergent nozzle are adopted.

The vibration method used on dual or multiple coaxial liquid columns to manufacture uniform droplets is

popular [15] such that artificial salmon roe is well known. Though the technique is known, fundamentals of this

389
method have not been discussed.

This work is concentrated in two areas. The

first is to clarify the relationship between the upper and

lower limits of vibration frequency of the uniform

droplet region, and upper and lower limits of applied

AC frequency. The second purpose is to clarify the

fundamentals of generating conditions for dual uniform

sized droplets by means of the electrostatic method from

dual or multiple coaxial smooth jets vMch are

surrounded by electrically conductive liquid, where inner

liquids are not conductive.

- i

k

9)

Q) Dual capillary nozzle

(D High voltage power

source

(D Electrode

(D Flowmeter

(D Tank

(§) Reducing valve

(2) Compressor

(§) Camera

(D TV Camera

@ Stroboscope

Fig.l Schematic diagram of experimental apparatus.

EXPERIMENTAL APPARATUS AND METHOD

Experiments were conducted using the

apparatus shown in Fig. 1. As shown in the Hgure,

inner and outer Uquids are supplied from the storage

vessels (5) using static air pressure to prevent Uquid

flow disturbance. Flow rates of each liquid are

controlled with regulating the air pressiire supplied from

the air compressor (7) through the reducing valve (6)

and measured using flowmeters (4). The inner and outer

liquids are forced into the dual capillary nozzle (1) and

discharged into ambient air. Hie discharged liquids are

consisted of a co-axial dual liquid jet column. The dual

capillary nozzle (1) and the circular plate electrode (3)

are connected to a high voltage AC power source (2). The disintegration patterns of the Uquid column are

recorded by means of shadow and scattered Ught

photographs using a stroboscope (10) as a lighting

source by a still camera (8) or a video camera (9).

The detailed geometry of the dual c^illary nozzle is shown in Fig. 2. In this case, kerosene is discharged

from the inner nozzle as an example of non or poor electrically conductive hquids and from the outer nozzle distilled

water is discharged as the conductive Uquid.

Fig.2 Geometry of dual capillary nozzle.

RESULTS AND DISCUSSIONS

Observations of the disintegration patterns

In Fig. 3, disintegration patterns of the coaxial dual liquid columns of mean velocity of 300 cm/s, A\4ien 2

kV AC potential with frequency of 1.0 - 3.0 kHz is appUed, are conpared with pattern ^^en no AC potential is

appUed. The Uquid velocity at the exit port of the inner and outer nozzle were both kept at 300 cm/s. As seen in

the Fig. 3, spontaneous irregular disintegration of smooth jet column was observed wiien no AC potential is

390
10

20

30

I mm f = 0[kH2]

V = 0[kV]

Fig.3

^plied.

When the Eqjplied

AC frequency was changed

in steps of 0.5 kHz interval,

as seen in the Fig. 3(b) -

(f), the patterns of (b) and

(c) show irregular

disintegration at 1.0 and 1.5

kHz. However, at the

frequencies of 2.0 and 2.5

kHz one can observe that

the disintegrated droplets

were generated at nearly

equal interval synchronizing

with the appUed AC

frequency (Fig. 3(d), (e)).

Furthermore, when the

appUed AC frequency was

increased up to 3.0 kHz, the disintegration interval and

droplet size became irregular. Thus, it was found that

there was a synchronizing region in vAach coaxial dual

Uquid column disintegrated with a uniform interval

distance and droplet size, when the ^pUed AC potential was kept constant.

Upper and lower limits of synchronizing frequency

In the Fig. 3, it is shown that there are upper

and lower limits of frequency in wiiich synchronized

disintegration occurs changing applied AC frequency

with the voltage maintained constant. This region of

the frequency in which synchronized disintegration of

smooth jet column occurs is referred to the "uniform

(synchronous) region". The effects of the Uquid mean

velocity on the synchronizing upper and lower limits of

the frequency of the uniform region was investigated,

when outer and inner Uquid velocity are kept the same. The upper and lower limits of the frequency of

the uniform region was shifted to a larger frequency

range when the Uquid velocity was increased.

Furthermore, it was observed that there is a minimum

potential over which uniform droplets can be produced.

It is also shown that inspite of changing AC voltage

*

t

:

f•

I

1.0

2

(b)

1.5

2

(c)

2.0

2

(d)

2.5

2

(e)

3.0

2

(a) (f)

Disintegration patterns of the dual liquid jet column with and without

AC potential

.

—I

I I I Vac =2[kV]

Key • 'H.obs. o 'L.obs.

100

J I I I 500 1000

Velocity [cm/s]

Fig.4 Upper and lower limits of AC frequency of the

uniform droplets region and comparisons with

calculated values by empirical equations of

mechanical vibration method.

391
the upper and lower limits of AC frequency were kept constant for each liquid velocity. Those phenomena agree

with previous reports on uniform droplets generated by mechanical vibration [1-5, 8]. The authors showed a

parallelogram in wliich uniform droplets were generated. The existence of the minimum applied potential woxild

suggest the essential minimum initial displacement or disturbance on ajet surface by an electrostatic force.

In Rg. 4, an example of upper and lower limits of frequency of uniform region is shown. The upper and

lower limits of frequency line are parallel as in the case of mechanical vibration. Furthermore, the straight lines

calculated by the empirical formulae from the mechanical vibration method [4, 5] agree with the experimental AC

potential method, as expected.

Size distribution of uniformly-sized dual droplets

Uniformly-sized dual droplets were generated within a region of upper and lower limits of applied AC

frequency. The size distributions of droplets with and without AC potential for the 300 cm/s liquid column are shown in Rg. 5. The upper figure (a) shows size distribution of falling spontaneous non synchronized disintegrated

drops \vb.en AC potential was not applied. As seen m the figure, the size spread is within a wide rage of 400 - 850

fim. On the contrary, as is evident in the lower figure (b), the droplet size showed a narrower rage of 500 - 600 fxm

and a sharp peak \vhen an adequate AC potential was appUed.

From the experimental results, it was found that uniform droplets were generated wiien adequate AC

potential was apphed on a smooth jet of coaxial dual Uquid column, the same as for the mono liquid column of

electrically conductive liquid.

Observation of free falling dual droplets

Instantaneous shadow photographs of the free

falling droplets were taken by means of a telescopic

microscope cameras as shown in Rg. 6. The outer and

inner liquid velocities were 200 cnVs, respectively. The

photogrq)h on the left shows a droplet produced by

spontaneous disintegration of a smooth jet when AC

potential was not applied. As seen in the Rg. 3, at the

frequency of 0 and 1.0 kHz, the sizes were unequal,

though each drop consists of both droplets as shown in

the Fig. 6 left. The middle photogr^h of Rg. 6 shows

synchronized uniform droplets as an example when the

AC frequency reaches 1.5 and 2.5 kHz within the

estabUshed uniform region at the same AC potential of

2 kV and the Uquid velocity of 300 cm/s. The

spontaneously and synchronously disintegrated droplets

showed double sphere drops involving one kerosene drop

within each water sphere, though double spheres could

not be detected in Fig. 3, vMch were taken by means of

a usual camera.

However, in the photograph on the right side.

E

a. 50

o

to

Z

<

o

o

a

Q

Vac=0[V] .

Non synchronized region

500 600 700 800

Droplet diameter [fivn]

a) Non synchronized spontaneous disintegration

100

E

a. o

in^

50

Z

<Jo

o

a

a

<

VAc=2[kV]

f = 2.1[kHz]

Synchronized region

400 500 600 700 800

Droplet diameter [^m]

b) Synchronized disintegration

Fig.5 Size distribution of droplets between free falling.

392
o

Vac = 0 [kV]

f = Non

Spontaneous disintegrationiNon electric field

Vac = 2 [kV]

f = 1.5 [kHz]

Within synchronized

region

Vac =2lkVl

f =2.0lkHzl

Over synchronized

region

Fig.6 Instantaneous shadow photographs of droplets between free falling discharged from

a dual coaxial nozzle.

the droplets were generated over the upper lunit of the frequency of uniform region, the outer drop size became

irregular and the number of inner drops were varied. In this case, two, three or more drops were involved, though the

outer size of each droplet seemed nearly equal in size from this photograph. It is considered that in this region of

disintegration, sateUites were produced and sometimes they were absorbed into the main droplets therefore the drop

sizes were not uniform and the sateUite formation could be one of the reasons which determine the number of inner

particles. However, further investigation is needed to clarify the reason why the inner particle number is varied.

Sizes of inner droplets

The sizes of the inner droplets were determined by calculation since to measure directly from the photograph was difficult because of the strong lens effect by the outer droplets. When the outer and the inner liquid mean

velocities are kept uniform, the interface of both liquid should be formed between the thickness of inner tube wall.

Then outer and the iimer flow rate ratio Fo/Fi= y will be expressed using the modified nozzle diameter ratio 6, when

t is the wall thickness of inner tube, Do and Di are defined in Fig. 2.

6 = Do/(Di+t) (1)

Y =62-1 (2)

Then, the outer and the inner drop size ratio p should be

P = (1+y)^^^ (3)

From equations (2) and (3)

p=5^^ (4)

2/3 In this experiment, the center photograph of Fig. 6 is 6 = 2.0, p = 2.0 = 1.59. Therefore, the outer

droplet diameter do determined from the photograph is do = 525 ^m, and inner diameter is di = do/p = 331 (im.

Surface wave growth rate and effect ofappUed ACfield

Grovyth rate and uniform region. The authors investigated theoretically and experimentally on the

relation between surface wave growth rate and mechanical vibration in order to generate uniform droplets [8], using

393
instability theory of Rayleigh [16] and Weber [17]. Rayleigh had indicated that the surface wave displacement along

a smooth jet is

^ = exp(at+ikz) (5)

Here, ^ is surface displacement, a is growth rate of surface displacement in mm/s, t is time, k is wave nimiber, z is

distance in axial direction from nozzle port and i is imaginary. He neglected the viscosity term of Navier-Stoke's

equation and reduced an equation of maximum surface wave growth rate as equation (6), hereafter referred as growth

rate a.

=(ct/p a^)(l-k^a^)kaIj(ka)/lQ(ka) (6)

Here, Iq andl^ are Modified Bessel Functions of the first kind. A, is wave length, a is surface tension, p is density

and a is radius ofjet at any distance, k = 2k/X, and ka is referred as dimensionless wave number.

On the other hand, Weber neglected the inertia term of Navier-Stoke's equation and obtained an equation

regardmg growthrate.

+ a(3}i/p s?)k^^ = (a/2p ahl-\^aVa (7)

Here, (i is viscosity of liquid. The surface wave displacement ^ measured by means of the serial instantaneous

shadow photographs [4, 5], should give a straight line from the equation (5) \vhen log ^ vs. time t is plotted. The

gradient gives growth rate a. The intercept on the vertical axis is corresponded to initial displacement ^q. The

actual smooth jet radius a can not be read from the photographs directly because of the existence of surface waves at

any distance z from the nozzle. The radius a at distance z is calculated by [17]

a = aQUQ^/2(i+2gzAiQ2) (8)

Here, g is the gravity acceleration and Uq is the initial mean velocity of liquid jet. Conversions from measured z to

t are calculated considering gravity acceleration and neglecting air friction. The critical viscosity for using either Rayleigh or Weber's equation for the calculation of surface wave

growth rate is 50 c.p. as pointed out previously [8]. As seen in Fig. 4, the experimental frequency limits of f and

fj^ by the AC method agreed with the calculated lines of f^j and f by the empirical equation of the mechanical

vibration method for low viscosity liquids under 50 c.p. [4, 5].

Growth rate a. along dual jet surface. As described in the former section, growth rate a is strongly related

to the upper and lower limits of vibration frequency regarding to the uniform droplet region, while Weber [17] had

proposed that the breakup length L of smooth jet column should be expressed as L = (u/a)hi(a/eQ) (9)

The value ln(a/^Q) is known as a function of Ohnezorge number. He pointed out that when initial displacement

grows up to ^, as expressed in equation (5) equal to jet radius a, disintegration of Uquid column occurs following

equation (9). However, the jets discussed in many former investigations were merely on mono phase and single

smooth jets.

The growth rate a of the spontaneous wave along coaxial dual smooth jets were measured without applying

mechanical vibration or AC field by plotting log ^ vs. t, when the outer and inner Uquid velocity were 200 cm/s.

The measured points were lined up straightly and both of the gradients of two groups, i.e., water in water jet and

oil in water jet, agreed with the theoretical gradient a which was calculated by the Rayleigh's equation for water

394
\vben the non dimensional wave number ka = Jid/^Qjjg = 0.776. The average of oil in water jet was about 0.1

fim, on the other hand the average of water in water jet was about 0.01 iim.

From this evidence the effect of inner nozzle geometry and physical property of liquid on the variation of

scattering and initial disturbances may suggest the importance of the geometry of inner nozzle port and sharp knife

edge is desirable.

From the experiments it was found that the growth rate a of co-axial dual jet surface wave is ruled merely

with physical property of outer Uquid when the viscosity of inner Uquld is in a same order with outer liquid.

Moreover, the growth rate a agreed to the Rayleigh's equation in spite of the coexistence of inner Uquid.

Key kV Hz

O 0 0

3 1300

Oil in Water Jet ,= 2 m/s

-1

Effect of applied AC field. Fig. 7 shows an exan^le of growth rate a along the surface of the coaxial dual

jet column v/ben AC potential applied within a uniform region is compared with a spontaneous natural jet. The

gradient of fitted straight line of AC appUed jet was calculated from Rayleigh's equation using physical properties of

the outer Uquid but the fitted line of the spontaneous jet was drawn without using regression analysis. When 3 kV

and 1.3 kHz AC was qjpUed to the jet the uniform droplets were generated synchronously with the applied AC

frequency, i.e. one droplet per each cycle, and the jet without the AC generated irregular sized droplets with

spontaneous natural disintegration. The cross point with aQ line corresponds to the breakup length of the jet and

thatofthe AC appUedjet is shorter than the non appUed jet, although the gradients of the two lines were nearly

equal. The breakup length of the spontaneous jet was L =

"1

25 mm, t = 12 ms and \^en 3 kV, 1.3 kHz AC is appUed

it became L = 20 mm and t = 9.5 ms. The intercept of vertical ^ axis increased as AC

potential was ^pUed. Thus this experimental results may

be summarized that as the AC potential is appUed to the

conductive Uquid column the initial displacement on the

jet surface increases, however, surface wave growth rate was

not changed visibly. Furthermore, it was clarified that

theoretical growth rate a along coaxial dual jet surface

foUowed to the Rayleigh's equation despite that AC

potential was appUed considering only its outer liquid

physical properties.

According to the electromagnetic dynamics the

electric field should reduce the surface tension and resulted

surface wave growth rate due to Rayleigh's equation, but the

experimental results did not show this. It is considered that

the effect of the high potential on the surface tension

reduction should be less or negligible, and the electroattractive or repulsive forces caused by the high potential on

the Uquid column should have the affect of increasing the

initial displacement on its surface similarly as the

mechanical vibration, though further physical considerations

stiU remaining.

E

E

10"

10

-3

Gradient by Eq.(6)

u = 1288 mm/s

I I I 8

t

10

[ms]

12

Fig.7 Comparison of surface wave growth along

coaxial dual jet by spontaneous disintegration

and AC applied disintegration within uniform

region.

395
CONCLUSION

The following conclusions were obtained.

1) Uniform droplets were generated wiien adequate AC potential is applied on a coaxial dual smooth jet. A

minimum voltage existed for generating uniform droplets. Also, within the upper and lower limits of AC frequency

uniform droplets were generated and the frequency of both limits were agreed with those of the mechanical vibration,

relating with jet velocity. The uniform dxjal droplets were generated synchronously with AC frequency, i.e. one

droplet per each cycle. Each droplet contained one same sized inner droplet.

2) The surface wave growth rate along coaxial dual smoothjet follows the theoretical growth rate calculated

by Rayleigh's equation using physical property of outer Uquid when the viscosity of both liquid is almost the same.

3) Surface wave growth rate along coaxial dual jet wiien AC potential is appUed agreed with that of

calculated by Rayleigh's equation using induced wave number and physical property of outer liquid only. The

minimum AC potential to produce uniform droplets is considered corresponding to the essential initial displacement

of mechanical vibration method to generate uniform droplets to overcome the back ground noise.

4) The AC potential may cause initial displacement wiiich will propagate along the surface of coaxial dual

smooth jet most quickly than spontaneous disturbances and causes disintegration regularly following appUed AC

frequency and at the end of the jet the displacement grows up to jet radius.

ACKNOWLEDGEMENT

The authors wish to express their appreciate to Kanegafuchi Chemical Co. Ltd. A part of this experimental

work is assisted by their financial support.

LITERATURE CITED

1) Sakai, T. and N. Hoshino: The 7th JCLASS (Japan Conference on liquid Atomization and Spray Systems), pp.

105 (1979).

2) Sakai, T. and S. Senuma: The 9th JCLASS, pp. 1 1 (1981).

3) Sakai, T. and N. Hoshino: J. Chem. Eng. Japan, 13(4), 263 (1980).

4) Sakai, T., M. Sadakata, M. Saito, N. Hoshmo and S. Senuma: ICLASS '82 (Wisconsm), pp. 37 (1982)

5) Matsushita, K., M. Sadakata and T. Sakai: ICLASS '88 (Sendai), A2 (1988).

6) Sakai, T., M. Sadakata and K. Matsushita: The 11th JCLASS, pp. 37 (1983).

7) Sakai, T., M. Sadakata, M. Saito and K Matsushita: The 12th JCLASS, pp. 11 (1984).

8) Sakai, T., M. Sadakata, M. Saito and K. Matsushita: ICIASS '85 (London), VIIB(b)/2/l (1985).

9) Sato, M. and T. Sakai: The 4th JCLASS, pp. 49 (1975).

10) Sato, M., M. Kito and T. Sakai: The 6th JCLASS, pp. 9 (1977)

11) Sato, M., H. Negishi, M. Sadakata and T. Sakai: The 15th JCLASS, pp. 73 (1987).

12) Sato, M.: J. Electrostatics, 15, 237 (1984).

13) Sato, M.: IEEE/IAS Annu. Meet. (Atianta), pp. 1406 (1987).

14) Sato, M., R Miyazaki, M. Sadakata and T. Sakai: ICLASS '88 (Sendai), B3-1 (1988).

15) Fullwiller, M. J. and C. W. Hatcher: Japan Patent, 52-129686 (1977).

16) Rayleigh, L: Phyl. Mag., 177 (1892).

17) Weber, C: Z.A.M.M., 11, 136 (1931).

18) Sato, H. and M. Saito: ICLASS '82 (Wisconsin), pp. 19 (1982).

396
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 41

ATOMIZATION OF LIQUID CYLINDERS, CONES, AND SHEETS BY

ACOUSTICALLY-DRIVEN, AMPLITUDE-DEPENDENT INSTABILITIES

J.L. Dressier

ABSTRACT

Fluid Jet Associates

Spring Valley, OH, U.S.A.

The atomization of several liquid jet configurations by an acoustic

disturbance is studied. The atomizer forms a cylindrical, conical, or flat(a

sheet) liquid jet by means of an electroformed nozzle. A piezoelectric driver,

which has been constructed to efficiently couple energy to the liquid, modulates

the fluid velocity. Operating the piezoelectric drivers at high power levels

produces perturbations with sufficient energy to break the liquid jets into

drops, with a net increase in surface energy. The resulting drop sizes are

influenced by the frequency and amplitude of the driving signal, as well as the

nozzle size. The spatial distribution of the spray is controlled by the spacing

and geometry of the holes in the nozzle plate, the amplitude of the acoustic

signal, and the swirl in the fluid manifold. This device is more robust than the

typical acoustic drop generator because small drops can be made from large holes,

reducing the plugging problem. No air flow is used.

INTRODUCTION

The atomization of a jet or sheet of liquid is a process that in most cases

requires the addition of energy to the fluid. The energy input is required

because the surface area of the atomized fluid is larger than the original fluid

surface area. The energy stored in the increased surface must be supplied from

a decrease in the kinetic energy of the fluid or from an external source.

Most fuel injectors for oil burners or gas turbines use an air blast to

atomize the liquid jet or sheet. A fast moving air stream impinges on the slower

moving fuel. The kinetic energy of the air tears the liquid into filaments and

then into drops. [1,2]

Another technique that has been used for specialized applications is the

acoustic excitation of a liquid cylinder at an unstable wavelength. This

technique does not add energy to the liquid because the fluid surface area is

lower after the disintegration has occurred. Rayleigh[3] derived the dispersion

equation for infinitesimal disturbances on a liquid cyMnder. His theory

correctly shows that the cylinder is stabl e for all but on e of the modes of

perturbation. The jet is unstable only for azimuthally symmetrical perturbations

(the sausage mode) whose axial wavelength is longer than the jet's circumference.

For this one mode, small perturbations will grow and the surface area of the

fluid decreases as the jet disintegrates into a stream of drops. The use of

small amplitude acoustic perturbations to produce a stream of monodisperse drops

from a circular liquid jet has been studied extensively .[ 4 , 5 , 6] The drawback of

this technique for atomizing fuel is that the size range of drops for a given

liquid jet diameter is restricted. The resulting drops have a diameter

approximately twice the diameter of the original jet. This restriction means

397
that small nozzles must be employed to obtain small drops and there is a great

chance of nozzle plugging due to particles in the fuel.

Dressler[7] has derived a dispersion equation for small perturbations on the

surface of a sheet of liquid. This analysis is similar to Rayleigh's analysis

of a circular jet. Dressier found that for perturbations driven solely by

capillary forces, the sheet of fluid is stable for all perturbations. Thus, the

technique of inducing atomization with a small acoustic signal works only for

circular jets; it does not work for sheets.

In addition to the air blast techniques, other methods have been devised to

overcome the capillary force and drive a fluid jet or sheet unstable.

Melcher[8,9, 10] , Crowley[ll] , Ketterer[ 12] , Dressier [ 13] , and Centani [ 14 ] studied

driving and stabilizing instabilities on fluid jets and sheets with electrical

forces. These previous electrohydrodynaraic studies were concerned with systems

that were unstable; an infinitesimal perturbation of the liquid surface would

grow without bound.

Our atomizer described here builds on our previous work[15,16] to

disintegrate the fluid and obtain a spray with a higher surface area than the

initial fluid. We use high amplitude acoustic energy to provide the energy

needed to increase the surface energy of the fluid. This process produces sprays

with the shapes of fans, cones, cylinders, and sheets. The process does not use

an air blast. By setting the frequency and amplitude of the acoustic signal, we

achieve some control of the drop size in the resulting atomized fluid

independently of any air flow.

ATOMIZER DESCRIPTION

Figure 1 shows a photograph of the injector. A piezoelectric driver is

placed in the circular fluid manifold. A nozzle plate, fabricated by a

photof abrication process, is placed across the end of the manifold. With this

design, the acoustic energy produced by the piezoelectric transducer is

efficiently coupled into the fluid, creating additional surface area on the

atomized fluid.

FIGURE 1.

High Amplitude Ultrasonic Atomizer

398
Several photographs of nozzle plates are shown in Figure 2. Figure 2a is

a rectangular slot, Figure 2b is a rectangular slot with width perturbations, and

Figure 2c shows a common round nozzle.

FIGURE 2a.

Rectangular Slot. 500 x 50 Microns

FIGURE 2b.

Rectangular Slot with Perturbations.

FIGURE 2c.

Common Round Nozzle. 50 Microns.

EXPERIMENTS ON CIRCULAR JETS:

Zero Order Azi muthally Symmetrical Mode (Sausage or Varicose)

Dressier and Jackson have previously shown that azimuthally symmetrical

disturbances, which are predicted to be stable by Rayleigh's criteria for small

perturbations, can be made unstable by the application of a large perturbation

in the jet's velocity. The energy in the perturbation, if large enough, can

cause the jet to break into a stream of drops. Because the breakup is dependent

on the acoustic amplitude, it is called an amplitude dependent instability.

399
Figure 3 shows a jet that is unstable for the applied perturbation. The

varicose disturbances grow in amplitude as the jet moves from the top to bottom

of the photograph. This process for jet disintegration is commonly called

Rayleigh breakup.

Figure 4 shows a jet that is stable for the applied perturbation. The

wavelength of the perturbation is smaller than the jet's circumference and the

perturbation decays as the jet moves from top to bottom of the photograph.

Figure 5 shows the jet of Figure 4 driven with a larger amplitude signal.

With the higher acoustic signal, the jet is unstable and breaks into a stream of

drops at the driving frequency. For this jet, the Rayleigh wavelength is 314

microns. The wavelength of this disturbance is about 280 microns. The drops in

Figure 5 have a volume that is 10% less than the smallest drop predicted by

Raleigh's theory for small perturbations.

FIGURE 3. FIGURE 4. FIGURE 5.

Circular Jets Exhibiting Circular Jet Exhibiting Circular Jet Exhibiting

Raleigh Break-Up Decaying Perturbations. Amplitude Dependent

Instability.

H igher Order Modes

Rayleigh' s Theory predicts that all of the higher order modes of a liquid

cylinder are stable for small perturbations. One method of countering the

stabilizing effect of surface tension is to apply an electric field to the liquid

surface. This technique has been shown to be useful for producing very

reproducible instabilities of the first mode. [10]

We have succeeded in driving a jet of liquid unstable in the higher modes

by the application of a high amplitude acoustic drive. Again, this instability

is dependent on the amplitude of the excitation. Figure 6 shows a liquid jet

with radial perturbations that are caused by our acoustic excitation. These

perturbations do not exceed the stability threshold and they decay as the jet

400
moves. Figure 7 shows a jet with a larger perturbation. These perturbations

exceed the threshold and the jet is unstable. The instability causes drops to

be ejected in the radial direction. There are 14 to 16 drops ejected from each

perturbation on the jet and therefore we must be driving the m=14 or 16 azimuthal

mode.

LIQUID CONE INSTABILITY

The photographs of the atomized jet show that there is a region of large

drops in the center of the spray pattern. To modify the spatial distribution of

the spray, we added swirl to the injector manifold. With this added swirl, a

hollow cone of liquid is produced from the circular nozzle.

Figure 8 shows the cone after an acoustic drive was applied. The excitation

forces the cone to break into rings of liquid which then disintegrate into drops.

The spray pattern formed is thus hollow.

FIGURE 6.

Stable Perturbations on

a 750 Micron Circular Jet.

FIGURE 7.

Unstable High-Order

Perturbations on a 750

Micron Circular Jet.

FIGURE 8.

A Cone Shaped Jet

Breaking into Rings Due

to an Acoustic Drive.

401
SHEET INSTABILITY

Dressler[7] developed a dispersion equation for small disturbances on a thin sheet of liquid. The derivation was made into a homework problem[9] and in

the final publication, the capillary force was changed to a magnetic force.

Since it is not available elsewhere, the original equation is given here without

derivation. This equation predicts that the sheet is stable for all small

disturbances.

- P

0)' coth2Kd(2K2T) + (k2T)2=0

(1)

where k is radians/length, w is radians/second, p is the fluid density

T is the surface tension, and d is the half-width of the fluid layer.

The dispersion equation can then be further factored into two equations.

The first represents disturbances where the sheet of liquid has a thickness

perturbation and the deflections of the sides of the sheet move in opposite

directions. The second equation is for kinking perturbations where both sides

of the sheet move in the same direction. These two dispersion equations are then

simplified with the long wavelength assumption that K is small. Because they are

not readily available elsewhere, the dispersion equations for the thickness mode

and the kink mode, for a liquid sheet driven only by capillary forces are

presented here.

po)2 = K^Td (2)

d

(3)

The kink mode, (3), of the sheet has a typical wave type of dispersion

equation while the thickness mode equation, (2), is

more interesting. For the thickness mode of the

fluid sheet and the sausage or azimuthally

symmetrical mode on the circular jet, the

dispersion equation is fourth-order in space and

second-order in time. Thus two of the

characteristics are horizontal, representing

infinite velocity of propagation. This strange

behavior results from the incompressibility

assumption, which gives more spatial derivatives

than time derivatives to the dynamic equations for

both the circular jet and flat sheet. If one has

an interest in studying flat or circular jet

deformations, he should be aware that this infinite

propagation velocity means that the initial

conditions at the nozzle exit are not sufficient to

determine the liquid behavior. One downstream

boundary condition must be considered.

Figure 9 is a photograph of a liquid sheet

originating from a slot nozzle. Small FIGURE 9

perturbations are applied to the nozzle by s^eet of Fluid Exhibiting acoustically perturbing the fluid velocity and gtable Perturbations

402
spatial perturbations are applied to the sheet thickness by putting perturbations

in the nozzle width as shown in Figure 2b. Figure 9 shows the sheet is stable

for these perturbations. The sheet was also stable for all other small

perturbations.

Figure 10 shows the sheet after it was made unstable in the thickness mode.

Here a large acoustic signal is applied to the sheet and it breaks into tubes as

a result of this large signal. Since a large excitation is required to produce

this sheet breakup, this instability is called amplitude dependent. The nozzle

plate used for Figure 10 contains thickness perturbations and they cause the

tubes formed from the sheet breakup to break into drops.

The higher order azimuthal modes on the circular jet were shown to have an

amplitude dependent instability. For this instability, drops were ejected

radially from the jet. Similarly, the second order or kink modes of the sheet

can be forced to become unstable if driven with a large enough amplitude.

Photograph 11 shows fingers of liquid extending from the sides of a sheet due to

the application of a moderate driving force. These perturbations are not large

enough to be unstable and Figure 11 shows them collapse back into the sheet.

When a larger perturbation is applied to the sheet, the perturbations become

unstable and drops are ejected from the edge of the sheet. This is shown in

Figure 12.

^ >>

FIGURE 10.

Sheet of Fluid Unstable

in the Thickness Mode.

Tubes Break into Drops.

Instability. Side View.

FIGURE 11. FIGURE 12.

Stable, High-Order Kink Unstable, High-Order Kink

Mode on a Fluid Sheet. Modes on a Fluid Sheet.

Side View. Amplitude Dependent

403
SUMMARY

The effect of high amplitude acoustic signal to the stability of fluid

sheets and cylinders has been examined. Without the large acoustic signal, a

fluid cylinder is unstable only for sausage mode disturbances with wavelengths

longer than the circumference of the cylinder. The sheet is stable for all small

disturbances. When a large acoustic signal is applied to a cylinder, it becomes

unstable in the sausage mode with wavelengths shorter than the jet's

circumference. For cylinders with large diameters, we found that we could drive

the surface unstable for higher order azirauthal modes. The instability caused

drops to be ejected radially from the cylinder.

Swirl was added to our atomizer and this changed the circular jet into a cone. A high amplitude acoustic signal forced the cone to break into a series

of expanding rings which then broke into drops.

Amplitude dependent instabilities were exhibited on a fluid sheet. The

thickness mode was driven and this stimulated the sheet to break into ligaments

which then broke into drops. The kink mode of the sheet was driven and small

drops were ejected from the sides of the sheet.

The acoustically driven atomizer can provide a reasonable flow rate with

no air mass. The device will be useful in situations where it is desired to

control the atomization without depending on the air mass flow. The drop size

and mass flow can be controlled by the acoustic amplitude and frequency and the

fluid supply pressure. One potential use for the acoustic atomizer is as a fuel

injector for reciprocating engines that operate at low speeds. By speeding the

vaporization of the fuel, it may be possible to have a fuel/air mixture which can

be ignited by a spark plug and which is not as rich in liquid fuel droplets as

present mixtures.

ACKNOWLEDGEMENT

This work was supported by contract F33615-89-C-2973 from the Air Force

Wright Laboratory ( WL-POSF) . The photographs were made by Marcia Dressier. The

author also wishes to acknowledge the assistance of Dr. Tom Jackson, Cindy

Obringer, Gary Switzer, and Lt. Robert Hancock.

REFERENCES

1. LeFebvre, A.H. , "Airblast Atomization", Progress in Energy and Combustion

Science, Vol 6, pp. 233-261 (1980).

2. Mansour, A. and Chigier, N. , "Disintegration of Liquid Sheets", Phys. Fluids,

A2(5) (May 1990).

3. Rayleigh, J.W.S., "On the Stability of Jets", Proceedings of the London

Mathematical Society, Vol 10, p 4 (1878). Available in U.S. in Chandrasekhar

,

S. , Hydrodynamic and Hydromagnetic Stability , Chapter 12, Oxford University

Press, London (1961).

4. Pierce, J.R. , "An Interesting Wave Amplifier" , IRE Trans. Elec.Dev., pp 73-74

(April 1960).

5. Mason, B.J. and Brownscombe J.L., "Production of Uniform Size Drops at a Controllable Frequency and Spacing from a Vibrating Capillary", Journal of

Scientific Instruments, Vol 41 (1964).

6. Dressier, J.L., and Kraemer, G.O. , "A Multiple Drop-Size Drop Generator for

Calibration of a Phase-Doppler Particle Analyzer", in Liquid Particle Size

Measurement Techniques (Hirleman, Bachelo, and Felton, Eds.), ASTM, Philadelphia

PA, STP 1083 (1990).

404
7. Dressier, J.L., "Active Electromechanical Control of Fourth-Order Continua",

Master of Science Thesis, MIT, Cambridge, MA (May 1966).

8. Melcher, J.R., Field-Coupled Surface Waves , Chapter 6, MIT Press, Cambridge,

MA (1963).

9. Melcher, J.R., Continuum Electromechanics , Prob. 8.12.1, MIT

Press, Cambridge, MA (1981).

10. Melcher, J.R., "Complex Waves", IEEE Spectrum, pp 86-101 (Oct. 1968).

11. Crowley, J.M. , "Growth and Excitation of Electrohydrodynamic Surface Waves",

Phys. Fluids, Vol 8, No. 9, pp 1668-1676 (Sept. 1965).

12. Ketterer, F.D. and Melcher, J.R., "Electromechanical Costreaming and Counter

Streaming Instabilities", Physics of Fluids, Vol 11, No. 10, pp 2179-2191 (Oct.

1968).

13. Dressier, J. L. ,"Videotype Sampling in the Feedback Stabilization of

Electromechanical Equilibria", in Feedback and Dynamic Control o f Plasmas , (T.K.

Chu and H.W. Hendel, Eds.), American Institute of Physics, NY (1970).

14. Centani, F.A. , "The Dynamics of an Electrified Sheet of Fluid", B.S. Thesis,

MIT Dept of EE, Cambridge, MA (1966).

15. Dressier, J.L., "Shock Model of Liquid Jet Breakup", Lawrence Livermore

National Laboratory, UCRL-78480, Rev. 1 (Oct 20,1976).

16. Dressier, J.L. and Jackson, T.A., "Acoustically Driven Liquid Sheet

Breakup", ILASS-Americas 4th Annual Conference, Hartford CT (May 1990).

405

ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 42

ATOMISATION AND COMBUSTION PERFORMANCE OF A

STEM-JET TYPE TWIN-FLUID ACOUSTIC ATOMISER

N.R. Ramesh' and R. Natarajan^

Propulsion Division, National Aeronautical Laboratory, Bangalore, India

tDepartment of Mechanical Engineering, Indian Institute of Technology, Madras, India

ABSTRACT

A modified version of the Hartmann-type stem and cavity air

jet acoustic generator coupled with a twin-fluid air-assist

atomiser has been used to form a unique practical device for

atomisation, emulsif ication and combustion applications. The

effects of acoustic and fluid dynamic fields on the atomisation

parameters have been investigated theoretically. These results

have been employed to evolve a step-by-step procedure for

designing the important parameters of the twin-fluid acoustic

device. Detailed experimental work has been performed to study

the effect of various flow and fluid properties on the mean drop

size of the spray. The results have been correlated to arrive at

a semi-empirical equation for estimating the SMD of sprays from

the acoustic device, applicable in the range of fluid properties

and test conditions employed. Incorporation of the acoustic

device in a practical combustion chamber has proved to be

successful

.

1 . INTRODUCTION

It has been postulated and observed that the acoustic field

in the spray zone enhances liquid sheet disintegration into

ligaments and drops resulting in better atomisation and efficient

combustion [1]. The Hartmann whistle and its modified form,

namely, the stem jet generator [2] which generates acoustic field

by directing moderately high pressure air through the centre of

the device into a suspended resonant cavity has been found to be

best suited for atomisation applications. The present work deals

with a detailed investigation of the design aspects of the

acoustic device and its performance evaluation by experiments.

2. DESIGN CONSIDERATIONS

The design of the acoustic device using air is generally

based on the empirical relations developed by Hartmann himself,

and with a few modifications as suggested by the later

investigators [2,4]. However, for the purpose of its application

to atomisation, emulsif ication and combustion, with the addition

407
of a liquid it is essential to consider some of the two-phase

flow parameters.

2.1 Equivalent Diameter Concept

The empirical equations of Hartmann [2] for designing the

acoustic generator have been suitably modified to account for the

presence of the stem rod by using an equivalent diameter

concept by replacing by D^^ and by D^^.

2.2 Acoustic Frequency

Acoustic frequency is calculated by considering the resonant

cavity as a quarter wave tube with one end closed, for which the

natural frequency of operation at the fundamental mode is given

by

(2n - 1)

a

4 (h + ^ ) fn = —

In practice it has been observed that for the given cavity

dimensions, f^ varies with the pressure of the driving fluid, and

the sonic velocity of the air-liquid mixture. Hence it is

required to calculate the actual f^ corresponding to the

operating conditions.

2.3 Velocity of Sound in Air-liquid Mixture

In an air-liquid mixture the resultant sound velocity is a

function of the void fraction ( o'- ) and the individual sonic

velocities which can be estimated using the equation based on

mean density and bulk modulus of the mixture [4] and can be used

for calculating the actual operating frequency

= 1/ [(<^/ c^; ) + (1 -o()/

5^ c^] (2)

2.4 Relation Between Drop Size and Frequency

A liquid drop, suspended in a gaseous medium subjected to an

acoustic field, undergoes periodic compression and rarefaction.

By considering the equation of motion and neglecting the gravity

effects it has been shown [3] that

= 2K (T / SF^ (3)

Using the general relation between driving freauency f and

the wave frequency F, as suggested by Lang, F = f/2, and the

equation for can be written as:

X= (8Tcr/ S f2)^/3 (4)

It is also well known that the median diameter of the liquid

droplet is proportional to X , and hence we can write:

408
d = kI, so that d = K (8T^/ 3| f^)^/^ (5)

where K is a constant to be determined experimentally.

2.5 Design Procedure

Based on the foregoing discussion, guidelines for a stepby-step procedure for the design of the twin-fluid acoustic

atomiser is suggested.

3. EXPERIMENTAL PROGRAMME

3.1 Configuration of the Acoustic Device

The sectional view of the acoustic device chosen for the

present work is shown in Fig.l.

Using the design procedure, a nozzle diameter of 6 mm was

selected, and the other dimensions were derived. Atomising air

was injected into the main body through a 10 mm dia tube. The

test liquids, and water (when used for emulsif ication) , fed

through 6 mm dia tubes, filled the annular groove in the

convergent-divergent nozzle and entered the main air stream

tangentially through three orifices of 0.8 mm diameter.

3.2 Test Facility and Instrumentation

The overall layout of the facility is shown schematically in

fig. 2. In all the atomisation experiments an optical bench

incorporating the forward diffractive light reattering technique,

built indigenously, was used for the measurement of spray drop

size.

3 . 3 Experimental Work

Extensive tests were conducted to determine the parametric

effects on the atomisation performance of the acoustic atomiser.

As a practical application, the acoustic device was then

incorporated into a gas turbine combustion chamber, and lean

stability limits, were determined. The results were compared

with those obtained by repeating the same experiments using a

conventional pressure jet atomiser.

4. RESULTS AND DISCUSSION

4.1 Results of Atom isation Studies

The influence of flow and fluid properties on the Sauter

mean diameter is shown plotted in figures 3,4,5. The increase in

liquid flow rate increases the drop size almost linearly, whereas

increase in atomising air flow rate and in turn the air velocity

decreases the drop size. The behaviour is similar for all the

four liquids. The optimum air liquid ratio lies between 4 and 8

409
105

Section X-X

FIG.l SECTIONAL VIEW OF THE ATOMIZER

Air

main

V

PR

PR

G Pressure gauge

PR Pressure regulator

V Control valve

Atomizer

Orifice, meter \^)~~\

Liquid

tank

-XManometer

FIG. 2 SCHEMATIC OF THE EXPERIMENTAL SET UP

410
for low viscous liquids, and is around 10 for furnace oil.

Generally the SMD of furnace oil is 4 to 5 times the SMD of low

viscous liquids.

Results of the present work have been compared with those

obtained by Nagai et al [8] (Fig.6). It can be seen that there

is good agreement between the two. Similarly the results of SMD

obtained for furnace oil have been compared with those calculated

from Jasuja's equation.

4.2 Drop Size Correlation

In the present investigation, the large amount of

experimental data obtained during atomisation studies have been

used to arrive at a semi-empirical equation for the prediction of

SMD, taking into account as many parameters as possible :

SM. = o.o.s.(f^y'X^JT^%, <e,

4.3 Results of Combustion Studies

The results of combustion studies have been analysed in

terms of the lean stability limits. Stable combustion was

possible at low fuel to atomising air ratios below a fuel rate of

about 7 g/s. At higher fuel flow rates, combustion was stable

for fuel/air ratios above 0.4. Visual and photographic

observations indicated that the flame cone angle was as high as

180 , and tended to spread backwards if the gap setting was

reduced. As the fuel flow rate was increased, the flame became

more luminous, with an increase in length and diameter.

The changes in the shape of the flame, viz., length and

diameter, and cone angle were observed as the fuel flow rate was

increased. It was seen that for kerossene and diesel fuels a

stable blue flame could be obtained with flow rates upto 5 g/s.

The flame was short and broad, and had a wide angle of coverage

with symmetric distribution.

Sound level and frequency were strong functions of the

nozzle diamter and other geometrical and flow parameters. In the

present case, the predominant sound frequency for the atomizer

was around 10 kHz. The sound level was 130 dB, which decreased

by 15 dB with liquid injection, and a further decrease of 10 dB

was observed during combustion. Using the measured acoustic

power output, the acoustic generator efficiency was calculated to

be around 47o.

While burning kerosene with water addition, an almost

transparent flame was observed. The reasons presumably are: (i)

homogeneous dispersion of water in kerosene spray (ii) the

isolated liquid droplets being re-atomised by the boiling of

water, thus achieving better mixing and in turn high intensity

combustion, (iii) reduction in soot formation by the lower

combustion temeperature with water. Flame length considerably

decreased during emulsion burning compared to neat fuel burning.

411
30

E

^ 20

a

s

10

0

Kerosene

J L

1 2 3

Liquid flow rate (g/s)-

Kerosene

J L

0 4 8 12

Air/ liquid ratio (by weight) -

16

Symbol o • A A

Atomizing air pressure kPa 160 239 298 367 436

(psig) 10 20 30 40 50

Atomizing air flow rate g/s 5.« 80 10.7 13.3 16.0

FIG. 3 EFFECT OF FLOW RATE ON FIG A EFFECT OF AIR/ LIQUID

SMD RATIO ON SMD

35

30

n

"eT 20

CO 10

Symbol o

Air/ liquid ratio 4

-r- Water

Kerosene

J L

Furnace oil

Atomizing air pressure = 298 kPa (30 psig)

J I

1.0 1.3 6.0

Viscosity ( \i) Hs/rr? X lO"''

300

100

-80

a. a^ 60

FIG. 5 EFFECT OF LIQUID VISCOSITY ON SMD

Jasuja's eqn. 40

Furnace oil

-L J.

t

i. 20

a

Nagai's result.

Kerosene

0 2 A 6

0 2 4 6

J Air/liquid ratio (by weight)—»> Air/liquid ratio (by volume ) X10.

FIG. 6 COMPARISON WITH AIR BLAST ATOMIZER

412
The effects of water addition on lean stability limits are

shown in Fig. 8. It is seen that water addition in general has a

deteriorating effect on the lean stability of the acoustic

atomiser

.

^•^ Results of Combus t ion Chamber Tests

Stability limits were obtained by carrying out a series of

extinction tests at constant, pre-determined levels of air

temperature and pressure. With the fuel flowing and the mixture

ignited, the fuel flow was gradually reduced until flame

extinction occurred. This process was repeated at increasing

levels of air mass flow rate until the complete stability loop

was drawn. Figure 9 illustrates the main features of the

stability loop obtained by this technique.

NOMENCLATURE

C Velocity of sound

D Diameter

f Frequency

h Depth of resonating cavity

1 Gap between the nozzle and the resonating cavity

m Mass flow rate

P Pressure

Q Volume flow rate

SMD Sauter mean diameter

V Velocity

W Power

oC Void fraction

t\ Efficiency

JA- Absolute viscosity

$ Density

cr Surface tension

REFERENCES

1. Ramesh, N.R. , Sridhara, K., and Natarajan, R., Fuel (London),

Vol.2 pp. 1677-1680 (1985).

2. Borisov, Yu.Ya in "Source of High Intensity Ultrasound" Vol.1

(Part 1), Ed., Rozenburg, Plenum Press, New York, (1969).

3. Lang,R. J., J. Ac. Soc. Am., Vol.34, No.l, pp 6-8 (1962).

4. Ramesh, N.R. , Rao, K.V.L. , and Natarajan, R., Proceedings of

the 10th National Conference in I.C. Engines and Combustion,

Rajkot, India, (1987).

5. Nagai, I.N. , Kumazawa, T. , Hayakawa, Y. , and Okazaki, H., Prod ICLASS, Tokyo, Japan ,pp . 209-219 (1978)

6. Jasuja, A.K. , ASME Paper, 78/GT/83.

413
Q 0.6

SIcn

oa

.t 0.2

. o

a.

-2^ A

^ Unstable region

I Subson/'c regioni *-]-» Scjnic regiop « 6 8 10

Atomizing air flow rate ( g/s) •

I L

12 16

100 200 300

Atomizing air pressure (k Pa) ^

FIG. 7 . LEAN BLOW OFF LIMITS FOR KEROSENE

400

y 0.6

<

O

au.

.i: 0.2 a

Symbol o A V o

•/. of water 5 10 15 20

0 _L

7

From Fig. 7-

2 4 6 8 10

Atomizing air flow rate Ma(g/s)

12 14 16

FIG. 8 EFFECT OF EMULSIFICATION ON STABILITY LIMIT

Combustor Air Flow (m;t • hiq 1 kgls

FIG. 9 LEAN STABILITY LIMITS FOR ACOUSTICALLY ATOMISED KEROSENE

414
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 43

CALIBRATION FACTORS FOR LASER DIFFRACTION RING DETECTORS:

THEORETICAL MODELING

S.B. Kenney and E.D. Hirleman

Mechanical and Aerospace Engineering Department

Arizona State University

Tempe, AZ, U.S.A.

ABSTRACT

Photodiode arrays used in laser di^ction particle sizing instruments must be calibrated to account for detectorto-detector variations in sensitivity. We have calibrated an Insitec EPCS-P detector by scanning a small laser beam

across the detector surface. A deconvolution of the known intensity distribution of the laser beam from the

measured signal resulted in detector response as a function of position. Detector response was approximately

constant over the region of the ideal active detector and it decreased exponentially in the region beyond the ideal

detector boundary. A diffusion length constant of SO^^I gave the best fit to the measured data. Theoretical

predictions of calibration factors based on measured detector response agreed reasonably well with Malvern and

Insitec calibration factors obtained from the traditional uniform light illumination method. This indicates that edge

effects in ring detectors made by different sources are similar.

INTRODUCTION

Particle sizing instruments based on near forward scattering signatures are commonly used to measure droplet

size distributions in sprays [1]. Scattering signatures are collected by the instrument and are processed by a mathematical inversion scheme to determine the size distribution. The equation which models the scattering

signature is a classical Fredholm integral equation. This integral equation is often approximated by a linear system

of discretized equations and the calculated size distribution is only an approximation of the actual distribution. The

corresponding coefficient matrix resulting from discretization contains elements that must be estimated. Two of

these components are the distribution of particle size within each size class and detector response characteristics.

The ultimate accuracy of the calculated size distribution is limited by the accuracy to which these components are properly estimated.

Since it is the particle size distribution that is sought, it is doubtful that any information on within class size

distributions would be known a priori. However, it is possible to experimentally probe the detector response.

Literature references [2-4] discuss calculation of the scattering matrix assuming an ideal step response at the

boundary of detector elements. Any deviation in the actual detector edge response characteristics from this assumed

ideal behavior would result in a bias error in the matrix, and hence, also in the calculated size distributions. Further,

all scattering matrix calculations reported to date have assumed no detector-to-detector variation in responsivity, the

ideal situation. However, since interlaboratory studies [5] highlighted the significant effect of between-detector

responsivity variations on the overall accuracy of laser diffraction instruments, calibration factors have been

introduced to correct for these between-detector effects. The process of calibrating a photodetector array using

uniform illumination [6] can e:q)erimentally correct for two effects, 1) spatial variations in the local responsivity

(amps/watts) either within or between detectors, and 2) edge effects. Qearly, details of these local effects are masked by the averaging effect of uniform flood illumination. In this paper we are concerned with the details of

detector array characteristics on a spatial scale much smaller than the detector dimensions. In particular, we report

415
e;q)eiimental measurements of local respoasivity both within and at the boundaries of the detector elements. We

have used the results to develop a theoretical model for the edge effects and we therefore are able to theoretically

predict calibration factors.

MATHEMATICAL FORMULATION

Scattering measurements in laser diffi^ction instruments are typically made with annular ring shaped detectors

which cover a finite range of scattering angles as determined by the detector ^rtures. It is convenient to assign a

particular scattering angle 0^ to represent the range of scattering angles in the ith detector aperture. The detection

process can be represented as:

iw(®i) = / we(e,ei)i(e)de (D

0

where the weighting function vfQ(Q,Q^) describes the relative responsivity variations across the detector and l^Q^ is representative of the signal obtained from the ith discrete detector (the subscript w indicates a dependence on the

weighting function). The azimuthal, or q> dependence on the responsivity variations are significantly less important

than the 9 effects, and have been neglected. i(9) is the intensity (W/sr) diffracted at near-forward scattering angles

0. The weighting function W0 is what is to be determined experimentally. It is also noted that the particle

distribution should also be discretized in a manner similar to that used for the scattering intensity [7]. In that case we obtain a system of mg equations in m^ (a is the size parameter) unknowns where mg is the number of discrete

detectors and m^ is the number of discrete size classes. The linear system is written as:

I = K • N (2)

In Eq. (2) the mg elements of the vector I are 1^(90 as given by Eq. (1); the m„elements of N contain the bth

partial moments of the number of particles in the size class; and K is the instrument or system matrix whereby

element represents the diffitiction contribution of a unit measure of particles in the jth size class onto the ith

detector. Tne elements of K are given by:

eooo

Kij = / / kb(a,9i) w„(a,aj) wg(9,9i) da d9 (3)

00

where WQ((a,aj) is a weighting function for the jth size class and k|j(a,9p is a general scattering function which

gives the scattering contribution of a unit quantity of particles of size a into angle 9. The solution or measured

particle size distribution indicated by N in Eq. (2) can in theory be obtained by inverting the matrix K.

EXPERIMENTAL PROCEDURE

Fig. 1 shows a schematic for a laser diffiaction particle sizing instrument. A laser beam is spatially filtered,

expanded and collimated to a few millimeter 1/e^ diameter. Particles within the coUimated beam scatter light which

is collected by a receiving lens and reflected onto the detector plane. A commercially available Insitec particle

sizing detector head and electronics were used for the experiment . The layout of the Insitec detector is well-suited

for examining detector response, given its odd numbered rings on one half of the detector and even numbered rings on the other. This design has large non-responsive sections between adjacent rings, which provides good isolation

between adjacent detectors.

The Insitec detector was mounted in a cantilevered position on a programmable x-y translation stage which

moved in a plane normal to the beam axis. A HeNe laser beam was focussed to a 1/e^ spot 25)xm in diameter and

the spot was positioned in the small hole in the center of the detector. (The hole allows the strong unscattered light

at the center of the detector to pass through, thereby reducing stray reflections and inner ring cross-talk). Coarse

centering was performed by moving the detector until the light reflecting off the region surrounding the hole

disappeared (indicating the light was passing through the hole). The beam power was adjusted so that when the

laser was positioned in the middle of an active detector the resultant signal was just below the saturation level.

416
Beam Expander/

Spallal Filter

Receiving Lens

( Transform Lens

)

Detection Plane (Transform Plane ) vj^-

Fig. 1. Schematic of conventional laser diffraction particle sizing

system.

n-type

Fig. 2. Cross section of an Insitec photodiode detector showing the

ideal responsive region (R-active), the transition region (gap), and

the non-active region (masked).

Several runs were made by scanning across the entire detector on a line passing through the center in one direction

and then rotating the detector 90 degrees and repeating the scan in the other direction. Signals were recorded at l^m increments and the entire e^tperimental apparatus was shielded fix)m stray room light.

Before scanning the detector, tfie rings were examined under a microscope. Three distinct regions were noticed

which are shown in Fig. 2 and are labeled as follows: the ideal responsive or active region (Ractive)' transition

region (gap), and the non-active region (masked). Bounding each detector is a thin aluminum film, which has been

deposited to cover most of the area between detectors. It is assumed that the metal helps block radiation so that

masked regions are not responsive. Though Fig. 2 shows just one detector, all detectors were examined and found

to be similar.

To measure the gap width, the detector was placed on a two-axis translation stage that was coupled to LED

displays showing relative stage position. The resolution of the display was l|im. A camera connected to a 19-inch

TV screen was mounted on the microscope which allowed for easy visual inspection of the object. A small piece of

tape was placed on the screen to mark the starting and ending points for various features. An inner gap edge was

visually lined up with an imaginary tangent line on the edge that would be parallel to the tape. The stages were moved until the outer gap edge was reached and the difference between stage positions was recorded. In this

manner the width of each gap (both inner and outer per detector), was determined.

RESULTS AND DISCUSSION

The average gap width was approximately 8^m (results from all detector rings can be found in Table 1). The

largest measured gap width was 10mm while the smallest gap width was 6mm, though most widths were 8mm. The

table begins with detector 5 because the first four inner rings were damaged when a hole was laser drilled through

the detector center. Also included in Table 1 is the visual measured width of the Insitec detectors (from

microscope).

Fig. 3 displays the individual ring signal vs. radial distance for the first five inner odd-numbered rings. The

rectangular boxes represent the specified width of the detectors and therefore the ideal top hat response distribution,

while the curves represent the measured signal. The measured signal is a convolution of the detector spatial

response and the Gaussian intensity profile of the laser beam. Since the measured curves are centered around the

specified points, it was concluded that the laser beam was approximately centered about the detector. The

maximum signal is close to 8 volts, however this value is not quite reached on rings 4 and 5 because the width of the

rings is less than twice the beam width, (where twice the beam width represents more than 99% of the total beam

energy). Also, overlj^) between rings occurs because of the beam width and edge effects resulting from electron

hole pair and photon diffusion. Note how overiap decreases with increasing ring size, which is due to the larger

regions of non-responsive area between rings. The curves appear smooth except for the saturation region where

small deviations in signal are noticed. It should be mentioned that 1000 signals are averaged at each position.

In order to be able to calculate wq for use in Eq. (1) it is necessary to model the gap response. This response can be estimated by deconvolving the measured signal from the known Gaussian intensity profile of the laser beam.

After several runs were taken of the inner odd rings (results of rings 11, 13, IS) an average response in the gap region

417
Table I

Iiuitec ideal detector width and gap width. Visual dimensions were acquired by viewing the detector under a microscope.

Detector Visual detector Visual gap width

width (pm) inner (ftm) outer (|im)

12

3

4

5 27 7 7

6 29 8 8

7 38 7 8

8 47 7 6

9 55 8 8

10 65 7 8

11 77 7 8

12 92 8 8

13 108 7 7

14 127 7 7

15 149 8 9

16 177 8 9

17 209 8 6

18 244 7 7

19 285 7 8 20 333 10 6

21 392 7 8 22 455 9 10 23 532 8 10 24 625 8 8

25 730 '7 8 26 855 8 7

27 997 8 9

28 1165 8 7

29 1361 10 8

30 1586 7 9

31 1855 7 8

aI

I

200 400 600

radial ditanc* (pm) 800 1000

Fig. 3. Plot of detector signal vs. radial distance for five inner odd

rings. The rectangles represent the specified cfetector dimensions.

The curves are the signals resulting from a 1/e^ 25^m diameter

laser beam of wavelength 0.6328|Xm. Measurements were made at Iftm increments.

1.0

0.8

0.6

0.4

0.2

0.0

\

\

\

\

MmI adlv* ragton 9«P \^ masked region \

\V..,.,^^^^^^

20 40 60

radial distance (|iin) 80 100

Fig. 4. Plot of the average edge signal of rings 11, 13, and 15 vs. relative distance. The signals of the three rings are superimposed

to form one curve.

average

was calculated by determining the average signal for each ring at each l|ini position (over a 100|im distance) and

then superimposing the data to calculate an overall average. This average gap response curve is shown in Fig. 4.

Initial insight into a possible model for the gap response is found by looking for symmetry within Fig. 4. The

symmetry present in the curve points to a potential model based on a simple function, most likely exponential. A

logical model to predict the gap response based on physical conditions would include uniform response over the

idealized detector region, decaying exponential response over the 8^m gap width and no response in the masked or non-responsive regioa

Photons incident on the gap region are absorbed by the silicon where they create electron-hole pairs. The

electron-hole pairs begin diffusing through the material until they either reach die p-n junction where there is a

possibility that they will contribute to the overall current, or they lose their energy through collisions and recombine.

For an overview of the physics of photodiode operation see Sze [8]. As far as a model is concerned only one

parameter will need to be determined; the "diffusion length constant" governing the decaying exponential.

However, it should be mentioned that when convolving the Gaussian beam with the gap response function the

value of the diffusion length constant has only a slight impact on the slope of the curve, thereby making it difficult to estimate the proper value of the constant to match the measured data. Fig. 5 confirms this observation, where

three curves representing the same beam diameter, but different diffusion length constants are displayed. As can be

imagined the gap response resulting from different diffusion constants is highly variable, yet because the beam

width is much greater than the gap width the slopes of the curves are nearly equal. If the beam width was less than

the gap width then the slopes of the three curves would be much different. Furthermore, if the beam was reduced to a point source then the response would simply be the exponential itself. However, even with large beam widths a

418
l-o-xd-I.O -i>-xd-10.0^xd-100.0

maskad g ld*al activ* rsglon I

1

radal distane* (pm)

Pig. 5. Plot of theoretical signal curves calculated by convolving a 2^m beam diameter with the theoretical detector response. The

specified detector width is 55^m which corresponds to ring 9 of

the Insitec detector. Calculations are made with different diffusion len^h constants as indicated. The ideal active, gap, and masked

regions are also indicated on the figure.

radial diatano* (Mm)

Fig. 6. Comparison of theoretical and measured signal vs. relative distance for rin^ 9. The theoretical signal was calculated based on

a 25)lm beam diameter, 8^m edge width, 50)im diffusion length,

and 5S)im detector width.

difference exists between the curves. The results are sensitive to diffusion length in that the curves expand with

increasing diffusion length constant. Since the specified width of each detector is known, a theoretical calculation could use diffusion constants as a means to expand or contract the curve to match the measured data. First, the total laser power in the Gaussian beam

is normalized to one and the beam is moved in one micron increments over the idealized, gap, and masked regions.

At each position the expected signal is calculated by convolving the response function (assuming a diffusion length)

with the Gaussian beam intensity. (As can be seen from Fig. 2 there are instances when the width of the beam

covers all three regions of the detector.) These values are then compared to the measured data to find the di^sion

length constant that provides the best fit. E^qperimental data from ring 9 were compared to theoretical data calculated based on different diffusion length

constants. Assumed values in the theoreticjd calculations include a 25\un 1/e^ Gaussian intensity profile diameter

laser beam, gap width of 8^m, and total detector width obtained from visual inspection. Results showed that a

diffusion length constant of SOumtlO^im (which falls within observed measurements [8]) gave the best fit between

e^rimental and theoretical values. A plot of ring 9 normalized signal vs. distance is foutxl in Fig. 6, and as can be

seen fi-om the graph the fit is quite good. As a test for the model, comparisons were made between experimental

and measured results for many of the other rings and similar results were noted. Thus, it was concluded that a

simple exponential function s^pears to adequately predict the gap response of the Iiisitec ring detector. The

weighting function over the gap width can be written as:

weedge = P/P = e"^"^ (4)

where p(6) is the local responsivity, p is the maximum responsivity in the ideal active region of the detector, xd is the diffusion length constant, and x is the position measured fiom the ideal edge into the gap.

Now that the weighting function wg has been determined a new instrument matrix K can be calculated by

substituting the measured gap weighting fiinction of Eq. (4) into Eq. (3) in place of the usually assumed step

function. To quantify the effect of the detector weighting function on the instrument matrix, two matrices were

calculated; one assuming a step detector response, the other the actual measured gap response (and weighting values

of 0 and 1 in the masked and idealized regions, respectively). A plot of the difference matrix (element by element)

between the two matrices is presented in Fig. 7. The matrices were calculated on an area basis with rectangular

weighting functions, uniform by volume within-class distributions, and a Fraunhofer diffraction approximation.

The surface plot shows almost no change on the outer rings, while the difference increases as ring geometry

decreases. This is expected, as the ratio of gap area to visual detector area is much greater for smaller rings than

larger ones.

419
Fig. 7. Plot of the difference matrix between K based on a step response detector weighting function and K obtained with

theoretical detector response weighting function (which accounts

for detector response in the gap region between the mask and

idealized detector boundary).

0.«

10 100 1000 10000

detector whtti (Mm)

Fig. 8. Plot of calibration factors vs. detector width for various instruments and calibration techniques. The numbers in the legend

correspond to the following: 1 ) Malvern 2600, instrument #1,

uniform incoherent illumination [6,1 1]. 2) Malvern 2600,

instrument #2, factory supplied calibration factors [9]. 3) Malvern

2600, instrument #3, uniform incoherent illumination [10]. 4)

Insitec, uniform incoherent illumination. 5) Insitec, theoretical predictions based on gap response method.

It has abeady been mentioned that it is possible to predict calibration factors based on gap or transition region

response as opposed to systematic between-detector variations in the responsivity. The accuracy of the "gap"

calculated factors is based on the premise that the major contribution to calibration factors is from gap response. A

predicted calibration factor is defined here as the ratio of the ideal detector ou^ut to the actual detector output when

illuminated with light of uiiiform intensity. The ideal detector output is calculated by integrating the incident light

intensity over the ideal, visual detector area while assuming that the weighting function wg is constant and equal to one. The actual detector output is found in the same manner except the area over which the integration is performed

includes the gap region and the associated weighting function wg^^g^. In equation form we have:

/ 1 <1A J dA(ideal)

Cpred=-^

=— ; (5)

Jwe I dA J Weedge «*^gap) + J <^idcaD

where Cpj^ is the predicted calibration factor and I is the intensity (W/m^).

Traditionally, calibration factors have been determined based on illuminating the detectors with uniform light

intensity. We determined two sets of calibration factors for the Insitec detector. One set of factors was calculated

using Eq. (5), while the other set was determined using the uniform light method. Results of the two methods are displayed in Fig. 10 alcMig with calibration factors from 3 different [9-11] Malvern 2600 instruments determined

empirically with the uniform light method. Since the Insitec detector geometry differs slightly from the Malvern

detector geometry, the calibration factors are plotted vs. detector width. As is seen from the plot the curves are in

reasonable agreement. Noticed trends are that all factors decrease in value as detector area decreases.

Consequently, the inner detectors over-respond relative to the outer detectors. Calibration factors determined from

Eq. (5) are limited by the fact that e:q>erimental data were collected along two lines running through the center of

the detector instead of scanning the entire detector surface. Additionally, in Eq. (S) it was assumed that the

weighting function equals a constant over the ideal detector region, however, Fig. 4 shows slight variation in this

regioa One other area of uncertainty which we did not investigate is the ends of each detector. Though we

accounted for the end region in the calculations the region is different from the rest of the perimeter regions in that a masked area is not found 8^m from the ideal detector boundary. A drawback to the uniform light method is that

linear detector response is assumed with no cross-talk. If either of the above assumptions is not correct then errors

in calibration factors wiU propagate through the remaining factors that need to be determined. This results from the

requirement that values acquired from different light levels must be spliced together [3]. Overall, it appears that

calibration factors calculated using Eq. (5) compare reasonably well with factors determined with the uniform light

method, especially considering the fact that factors are for different detectors from different laboratories.

420
CONCLUSIONS

In conclusion, the transition region response has been measured and modeled for an Insitec detector. A simple

e;q)onential with a diffusion length constant of 50^m accurately predicted the experimental data. The detector

weighting function has been calculated and incorporated into the instrument matrix K. Calibration factors for the

Insitec detector have been predicted based on edge effects and were compared with calibration factors derived from

the uniform light source technique. Both methods indicated that the inner rings over-respond relative to the outer,

larger rings. It is difficult to conclude which method produces the more accurate results. The uniform light method

has the disadvantage that detector signals resulting from several light levels must be measured and the calibration curves spliced together. The gap response model has the disadvantage that it is based on 2-dimensional results and

it must be assumed that gap response is axisymmetric. Another independent technique for calculating calibration

factors could help verify whether the gap or uniform light source is more accurate. The effect of using the

theoretical calibration factors instead of the uniform light source factors would, in general, translate into slightly

larger particle sizes because of the higher signals on the iimer rings.

ACKNOWLEDGMENTS

This research was supported by the Air Force Office of Science Research, Air Force Systems Command, USAF,

under grant AFOSR-90-03S8, Dr. Julian Tishkoff, program manager. The U.S. government is authorized to reproduce and distribute reprints for governmental purposes notwithstanding any copyright notation thereon.

REFERENCES

1. E. D. Hirleman, "Particle Sizing by Optical, Nonimaging Techniques," in Liquid Particle Size Measurement

Techniques, STP 848, J. M. Tishkoff, R. D. Ingebo, and J. B. Kennedy, Eds. (American Society of Testing

Materials, Philadelphia, 1984), pp. 35-60.

2. J. Swithenbank, J. Beer, D. S. Taylor, D. Abbot, and C. G. McCreath, in Experimental Diagnostics in GasPfiase Combustion Systems: AIAA Progress in Astronautics and Aeronautics, B. T. Ziim, ed., AIAA V. 23, pp.

421-447. (1977).

3. E. D. Hirleman, V. Oechsle, and N. A. Chigier, "Response Characteristics of Laser Diffraction Particle Size

Analyzers: Optical Sample Volume Extent and Lens Effects," Optical Engineering, V. 23, No. 5, pp. 610-619

(1984).

4. M. Heuer, K. Leschonski, "Results Obtained with a New Instrument for the Measurement of Particle Size

Distributions from Diffraction Patterns," Particle Characterization, V. 2 pp. 7-13 (1985).

5. E. D. Hirleman and L. G. Dodge, "Performance Comparison of Malvern Instruments Laser Diffraction Drop

Size Analyzers," ICLASS Conference Proceedings, The Institute of Energy, London, V. 2, pp. rVA.3.1-

IVA.3.6 (1985).

6. L. G. Dodge, "Calibration of tiie Malvern Particle Sizer," Applied Optics, V. 23, pp. 2415-2419 (1984).

7. E. D. Hirleman, "Uncertainties in Matrix Formulations of the Fraunhofer Diffraction Particle Size Problem,"

Proceedings of ILASS America Conference, Institute of Liquid Atomization and Spray Systems, pp. 1-5

(1989), Irvine, CA.

8. S. M. Sze, Physics ofSemiconductor Devices, John Wiley and Sons, New York (1981).

9. O. S. Samuelsen, University of California, Irvine, CA (private communication).

10. E. A. Hovenac, NASA Lewis Research Center, Cleveland, Ohio (private communication).

11. L. G. Dodge, Southwest Research Institute, San Antonio, Texas (private communication).

421
i I
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 44

PARTICLE DIAGNOSTICS AND TURBULENCE MEASUREMENTS IN AN

ISOTHERMAL SPRAY UNDER THE INFLUENCE OF STABILIZING BODIES

A. Brena de la Rosa and W.D. Bachalo

Aerometrics, Inc. Sunnyvale, CA, U.S.A.

Abstract

The present work reports a study of the effect of three stabilizing bodies on the turbulence

properties and on the dynamics of an isothermal spray. A vane type swirler having a high swirl

number S = 1.23, a bluff body, and a bluff body with orifices were used to create a turbulence

field with recirculation. A pressure swirl nozzle was used to atomize water downstream of the

stabilizing bodies and produce a spray to study its turbulence properties and the interaction with

the surrounding air. Properties of the spray such as particle size distributions, particle velocity

distributions, and particle number density were abo obtained for the sprays in the flow fields of

these recirculating flows. The results of the paper show that the recirculation regions produced

by the three stabilizers are very different with very high turbulence intensities generated by

the bluff body with orifices. Spatial distribution of the particle size is also very different, most

notably the liquid drops in the swirling field seem to foUow well ordered trajectories which reflect

the effect of the circumferential velocity imparted to the drops. In contrast, the drops in the

wake of the bluff bodies exhibit trajectories which differ greatly for the various particle size

classes.

Introduction

The importance of stabilizing bodies such as disks, cylinders, V-gutters, and swirlers in reacting

and non-reacting flows hcis been recognized for some time. Beer and Chigier (1983). Combustion

in the wake of stabilisers causes changes in the shape and strength of the recirculation zones. In

addition, reaction in the recirculation region results in an increase in temperature and thus in

volumetric flow rate. Moreover, the rate of fuel evaporation, mixing and the rate of formation

of chemical species are greatly influenced by the structure of this recirculation region in the flow.

Mecisurements in the reactive case for gaseous fuels have been made by Winterfeld (1965) and

Bespalov (1967) in the wake of a V-shaped flame holder. In combustion systems swirl generation

is, by far, the most extensively used mechanism for stabilizing the flame and for controlling the

efficiency and generation of pollutants, Gupta et al. (1984). While swirl burners have been used

widely in combustion, a major effort has been put recently to elucidate the basic mechanisms that

govern swirl flows, Sislian and Cusworth (1986) - Cameron et al. (1988). In the application of swirl

combustion devices to utility and industrial gas turbine systems, special emphaisis has been placed

423
in understanding the mechanism of formation and emission of pollutants such as CO, unburnt

hydrocarbons {UHC), NOx, and soot; especially when it is recognized that measures customarily

taken to reduce, for example, CO, UHC, and soot tended to increase the emission of NOxAn important aspect in the design of combustion systems is the ability to predict and control

their performance and efficiency. One of the routes to achieving this is through numerical modelling

of the flow under study. However, this task is usually very complex if the combustion designer

does not possess some basic knowledge of the flow, which in turn must be obtained only from

experimental evidence. The purpose of this paper is, therefore, to provide some insight of the eff"ect of three difi'erent stabilizing bodies on the structure of the liquid spray.

Experimental Setup

The tests were carried out in a 46 X 46 cm^ wind tunnel with a free stream velocity of Vfg = 20

m/s for all the experiments reported here. Each of the stabilizing devices was placed on the liquid

supply tube of a pressure swirl atomizer and the assembly surrounded by an acrylic chamber 300

nmi ID which had a screen fitted in the upstream end to dampen turbulence fluctuations of the air

flow. This whole assembly was placed, in turn, inside the wind tunnel test section, see Figure 1. The swirler was manufactured of stainless steel and consisted of a hub 19 mm ID and eight

straight vanes welded equidistantly to it and forming an angle of 60° with the longitudinal axis of

the swirler. The design of the swirler followed the general guidelines depicted by Beer and Chigier

(1983). A more detailed description of the swirler configuration is given by Brena de la Rosa et al.

(1990). The 60° vane swirler chosen for this study provided a strong recirculation region and gave

a nominal swirl number of 5 = 1.23, Beer and Chigier (1983).

The solid disk and the disk with orifices were 76 nmi in diameter and 3.18 nmi thick. Two sets

of eight synrmietrically located orifices were drilled in one of the disks. One set of holes was located

at 32 nun from the center of the disk and its orifices were 6.35 mm in diameter, the other set was

located at 21 nmi from the center of the disk and the orifices were 4.76 mm in diameter, see Figure

1. The location and the size of these holes was arbitrary.

The measurements were obtained with an Aerometrics two-component Phzise Doppler Particle

Analyzer (PDPA) whose principle of operation is given by Bachalo and Houser (1984). The system

consists of a 3 Watt Argon-ion laser, however, the power used in the present experiments was 100

milliwatts measured at the probe volume. The PDPA was configured in the 30 degree off'-axis

forward scatter, where the scattering of light by refraction is the dominant mode and yields a linear

relationship of particle size versus phase between the photodectors over the detectable size range of

the instrument. Particle sizing was done with the 0.5145 /im green beams which also gave the first

component of velocity, while the 0.488 fim blue beams provided the second orthogonal component

of velocity. Frequency shifting was provided by a rotating grating. The optical components selected

to obtain the turbulence data were an / = 300 mm collimating lens and an / = 200 mm transmitter

lens which gave an eflfective particle detection range of 0.5 to 40 fim in diameter. For the spray data

an / = 160 mm collimating lens and an / = 500 mm transmitter lens yielded a particle detection

range of 4.0 to 300 /im in diameter.

Two sets of experiments were carried out for the three stabilizing bodies. Firstly, the flow was

seeded with steam particles having a mean diameter jDio = 1.8 /im and a Sauter mean diameter

Ds2 = 3.9 /im to obtain the turbulence properties of the flow in the absence of the spray. Secondly, a

pressure swirl atomizer operating at P = 415 kPa with a nominal flow rate of 3 gph, 45° solid cone,

was used to atomize water downstream of the stabilizing bodies. For each of these experiments

the axial, radial, and tangential velocity components of the flow were evaluated. The experiments

consisted of radial scans of the flow field which were obtained at five different axial locations,

424
namely, X = 10, 20, 50, 100, and 150 mm downstream from the face of the atomizer to evaluate the

velocity and turbulence fields. At these same positions (except at X = 10 mm where only velocity

information was taken), meaisurements of drop size, drop velocity, and particle number density

were obtained from the water spray in the recirculating field. At eeich axial position the radial

measurements were taken at 3-5 mm intervals from the centerline of the nozzle to the periphery of

the flow field.

Results

Figure 2 shows the radial distribution of the mean axial velocity of the air U obtained at 10

mm from the face of the atomizer for the three stabilizing bodies. Several interesting features can

be observed in this plot. Most notably is the diflference in the velocity gradients observed amongst

the three bodies. In the case of the bluff body, the gradients of mean axial velocity are lowest,

in fact, the velocity distribution changes between U = —2.5 m/s to about U = —7.5 m/s over most of the flow field at this location. The vane swirler S = 1.23 yields an almost constant axial

mean velocity from R = —25 mm to i? = 25 mm and subsequently increases to the edge of the

swirling field. The bluff body with orifices yields the sharpest changes in mean axial velocity with

two peaks in the radial distribution. The first occurs at jR = 20 mm just before the place where

the row of small holes is located. The trough observed at /2 = 30 mm in the radial distribution of

U is located where the large holes are placed. From this point the velocity gradients are steepest

and the velocity increases towards the edge of the solid disk and eventually decreases towards the

boundary of the acrylic chamber.

Although not shown, the radial distribution of the radial and tangential velocities of the air

were obtained. As expected, the bluff body and the bluff body with orifices provide essentially no

rotational velocity to the flow. The air field produced by the swirler showed a steep gradient in

azimuthal velocity at the edge of the swirler close to the nozzle.

Figure 3 (left) gives a plot of the turbulence Reynolds shear stresses uV for the three stabilizing

bodies at three axial locations X = 10 mm, X = 20 mm, and X = 50 mm from the atomizer. The

interesting features about this figure are that the bluff body shows regions of relatively high shear

at /2 = 20 mm and R = —20 mm probably caused at the boundaries of the recirculation eddy.

The vane swirler produces a practically uniform region of shear except at the boundaries with the

external air flow field. The bluff body with orifices exhibits a radial distribution of shear stress u'v'

whose peaks are consistent with the locations where the gradients of the mean velocity field U are found, i.e., at i? = ±20 mm and R = ±32 mm, see top of Figure 3. The third peak is result of the

flow being sheared by the external air flow field. At X = 20 mm the bluff body depicts fluctuations

of the shear stress profile throughout the flow field. The bluff body with orifice shows peaks in

the shear stress profile caused by the jets issuing from the holes. The swirling field does not show

regions of high shear except at the boundaries with the external air flow. At A" = 50 mm the shear

stress profiles do not exhibit regions of high shear except at the boundaries of the flow. Notice that

for the swirler, half-radial profiles of the shear stress at X = 20 mm and X = 50 mm were only

available.

Also shown on Figure 3 (right) is a comparison of the mean axial velocity of discrete drop sizes

having 10 /xm, 20 /im, 30 fim, 40 fim, and 50 /xm in diameter with the mean velocity of the air.

At X = 20 mm, in the case of the bluff body and the bluff body with orifices, some of the particle

sizes reflect sharp changes in the mean velocity as a result of the vigorous mixing and of the jets

issuing from the orifices. In the case of the swirler at X = 20 mm, the mean axial velocity profiles

of the drops are similar to each other and the particles are decelerated up to the boundaries of the

recirculation region, i.e., R = 26 mm. Outside of the recirculation region the external air flow field

425
— F10« STWIGHieCF

Figure 1: Schematic of the swirl vane assembly fitted to the liquid supply tube of the atomizer

inside the wind tunnel test section. The same overall configuration was used (with the swirler

removed) for the bluflF body and the bluff/orifice stabilizers. Right: Schematic of the bluff body

with orifices.

E

20

15

10

5

0

-5

-10

X=10 mm o —o bluff body

^ ^ bluff/orifice

o S=1.23

A A A^

A-A-'^-A-A-A"

o-o-o-o-o-o-o^o^o- o-oH 1 H -I 1 H H 1 < 1 1 1 H

-50 -40 -30 -20 -10 0 10 20 30 40 50

Figure 2: Radial distribution of the mean axial velocity of the air at X = 10 mm from the atomizer

for the three stabilizing bodies.

426
(Nm\CM

IS

(/) LU

OH

I—

a:<UJX

CO

15

10

5

0

-5

-10

-15

o ——o bluff body

^ bluff/orifice

o —o S=1.23

X=10 mm

H—I—I

>—

I I —

-50 -40 -30 -20 -10 0 10 20 30 40 50

CM

CO \CME

>

c/i

on

a:<

bJ

X

(/)

CMm

CM

>

COCO

UJ

q:

ICO

a:<UJX

CO

10

5--

O-

-5-

-10

o —o bluff body

a — bluff/orifice

o —o S=1.23

/ \. o-o -a—

\,/

4- +- -+ -I- -+- X=20 mm

-I—'—I—<— -50 -40 -30 -20 -10 0 10 20 30 40 50

20

15

10

5

0

-5

-10

-15

-20

o —o bluff body

^ —^ bluff/orifice

o —o S=1.23

X=50 mm

H ' 1 •—

E

o

o

_1

>

_J<

X

<

-50 -40 -30 -20 -10 0 10 20 30 40 50

RADIAL POSITION, mm

15 J

CO \

E

10--

5--

o

o 0--

_J

Ld

> -5--

1VI — 1 n -

.

X

<

-15--

-1

15-rCO

E

1

1 nu • -

5--

.OCI 0--

UJ

> -5--

< -10--

X

<

-15--

A —4 10 /im

o —o 20 fjun

o —o 30 nm

V —V 50 fun

• —• Air

bluff body

X=20 mm

10 20 30 40 50 60

bluff/orifice X=20 mm

I

a —a 10 /im

o —o 20 /im

0 —o 30 /im

' —' 50 /wn

• —• Air

1 ' ' I

I -10

20

16

12

8

4

0

10 20 30 40 50 60

-4

S=1.23

X=20 mm

a-—a 10 /im

o —o 20 /im

« —o 30 /im — ' 50 /im

• —• Air

I

' I

I -10 0 10 20 30 40 50

RADIAL POSITION . mm

60

Figure 3: Radial profiles of the turbulence Reynolds shear stresses uV (left) and comparison of

the mean axial air velocity U with the mean axial velocity of discrete drop diameters for the three

stabilizing bodies.

427
is dominant and the drops exhibit large gradients of mean axial velocity, with the highest velocities

associated with the smallest drop sizes. Notice that negative mean velocities of the air are present

for the bluff body and the bluff body with orifices for most of the flow field at this location, which

is an evidence of the strong recirculation flow.

For reasons of space limitations, the profiles of the turbulence intensity and of the kinetic

energy of turbulence, Et, are not presented. Turbulence intensity shows a fast decay for the three

stabilizers and only the bluff body with orifices shows a region of turbulence at the shear layer with

the external air flow field. In the case of Et, the results show that the regions of high energy content

are located where the fiow is sheared by opposing high velocity air streams, i.e., at the boundaries of

the recirculation region in the core and at the boundaries with the external air stream. In addition,

the bluff body with orifices produces the highest Et close to the nozzle.

Figure 4 depicts the radial distribution of the Sauter mean diameter D32 for the three stabilizing

bodies at three axial locations. Larger particles were measured in the spray produced by the bluff

body at all axial positions when compared to the other two stabilizing bodies. It is evident that

the smaller particles are found in the core of the recirculation region for all the stabilizing bodies.

The radial distributions of D32 for the bluff body and the bluff body with orifices do not follow a

particular trend but rather reflect the intense mixing of the spray and the effect of the jets issuing

from the orifices, see for example, the D32 profile at X = 20 mm.

Also shown on Figure 4 are the radial distributions of particle number density Nd (number of

particles per cubic centimeter), for the flows produced by the three stabilizing bodies. Peak values

of Nd for the bluff body with orifices and for the swirler are about four times greater than those

produced by the bluff body. For the latter, this may be the result of the strong recirculation region

which prevents a relatively large concentration of drops at that location, (see Nd for the bluff body

at X = 20 mm, Figure 4). In general, the radial distributions of Nd are strongly affected by the

dynamics imposed by the different stabilizers, especially at locations close to the atomizer, X < 50

mm.

Conclusions

Meeisurements of the turbulence properties and of the particle dynamics of an isothermal spray in

the wake produced by a bluff body, a bluff body with orifices, and a 60° vane swirler have been

made. Due to space limitations, only some of the results have been presented here, however, a

complete discussion of the present investigation will be published in a journal.

The following conclusions can be drawn from the present work:

• The air flow fields produced by the three stabilizing bodies tested here produced drastic

differences in both the turbulence field and the dynamics of the liquid particles. Flow reversal

was observed for the three flow fields, with the bluff body yielding the highest negative axial

mean velocities at all axial locations, see Figure 4.

• Regions of highly sheared flow are present in the fleld produced by the bluff body with orifices

which persist to about X/D = 0,66 {X — 50 mm, where = 76 mm) from the nozzle.

• The turbulence properties of the flows show that the bluff body with orifices yields the highest

kinetic energy of turbulence and turbulence intensity, caused by the strong jets issuing from

the holes.

• The trajectories of discrete liquid particles are also very different for the three flow fields, the

largest differences amongst them observed close to the nozzle. In the case of the vane swirler,

428
140

B 110-

<

UJ

80-

< 50

in

bluff body

o —o X= 20 mm

A » X= 50 mm

0 —» X=100 mm

\

,o—

-10 10 20 30 40 50 60 70

(/) Z

UJQ

q:: m

ISZ

UJ

-Jo

H-<

Q.

1000

750-

500

250-

E

UI

<

<

UI

on

UI

I-<

C/)

100

90-

80--

70--

60--

50-

-10

bluff/orifice o —o x= 20 mm — X= 50 mm

0 —o X=100 mm

10 20 30 40 50 60

a:

UIm

z

UI

_l

o

I—

a:<

a.

4000

UIa 3000 -^

2000-

1000-

70

0-

-10

bluff/orifice o —o X= 20 mm

* —" X= 50 mm

0 —« X=100 mm

Vc

—o

10 20 30 40 50 60

E

a.

q:

UI

<

Q

<

UI

UII—<

120

80--

40--

S=1.23

-10

o —o x= 20 mm — X= 50 mm

<> —o X=100 mm

/

/

0 10 20 30 40 50

RADIAL POSITION, mm

60 70

z

UIQ

q:

UIm

UI

_i

o

I-<

Q.

4000

3000

2000 -

1000

0 10 20 30 40 50

RADIAL POSITION . mm

Figure 4: Radial distribution of the Sauter mean diameter Ds2 (left) and of the particle number

density Nd (right) for the three stabilizing bodies at three axial locations.

429
a clear evidence of the effect of the recirculation region (in the core) and of the external air

flow field on the droplet trajectories was present in all the results, see for example Figure 3.

Particle number densities resulted in peak values which are about four times greater for the

bluff body with orifices and for the swirler than those produced by the bluff body.

Acknowledgements

The authors gratefully acknowledge the support of this work by the National Science Foundation,

Contract ISI-8860783 and by NASA Lewis Research Center, Contract NAS3-25204. The authors

would also like to express their gratitude to Mr. Mark Yoneyama and Mr. Robert Kennedy for

their contributions during the experimental stage of the program.

References

• Bachalo, W.D., and Houser, M.J., 1984, "Phase/Doppler Spray Analyzer for Simultaneous

Measurements of Drop Size and Velocity Distributions'*, Optical Engineering, Vol. 23, No.

5, p. 583.

• Bachalo, W. D., Brena de la Rosa, A., and Rudoff, R. C, 1988, "Advances in Diagnostics

for Complete Spray Characterization" , Proceedings of the 4th International Conference on

Liquid Atomization and Spray Systems, Sendai, Japan, August 22-24.

• Beer, J. M., and Chigier, N. A., 1983, Combustion Aerodynamics, Robert E. Krieger Publishing Company, p. 111.

• Brena de la Rosa, A., Wang, G., and Bachalo, W. D., 1990, "The Effect of Swirl on the

Velocity and Turbulence Fields of a Liquid Spray", Presented at the 35th ASME Gas Turbine

and Aeroengine Congress and Exposition, Brussels, Belgium, June 11-14.

• Bespalov, L V., 1967, "Physical Principles of the Working Process in Combustion Chambers

of Jet Engines" , Translated from Russian, Clearinghouse for Federal Scientific and Technical

Information, USA, AD 658 372, p. 366, May.

• Cameron, C. D., Brouwer, J., Wood, C. P., and Samuelsen, G. S., 1988, "A Detailed Characterization of the Velocity and Thermal Fields in a Model Can Combustor with Wall Jet

Injection" , ASME Paper No. 88-GT-26, Presented at the Gas Turbine Congress and Exposition, Amsterdam, The Netherlands, June 5-9.

• Gupta, A. K., Lilley, D. G., and Syred, N., 1984, Swirl Flows, Abacus Press.

• Lilley, D. G., 1988, "Lateral Jet Injection into Swirling Combustor Flowfields", AIAA Paper

No. AIAA-88-3183, ALAA/ASME/SAE/ASEE 24th Joint Propulsion Conference, Boston,

July 11-13.

• Sislian, J. P., and R. A., Cusworth, February 1986, "Measurements of Mean Velocity and

Turbulence Intensities in a Free Isothermal Swirling Jet" , AIAA Journal, Vol. 24, No. 2, pp.

303-309.

• Winterfeld, G., 1965, "On Processes of Turbulent Exchange Behind Flameholders" , Tenth

Symposium International on Combustion, The Combustion Institute, pp. 1265-1275.

430
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 45

ATOMIZATION WITH SPILL-CONTROLLED SWIRL PRESSURE-JET NOZZLES

M. Loffler-Mang and W. Leuckel

University of Karlsruhe, Engler-Bunte-Institute

Combustion Technology Section

Karlsruhe, Germany

ABSTRACT

The entire atomization process using spill-controlled swirl pressure-jet nozzles was investigated. Part

I deals with the flow fields inside scaled-up plexiglass nozzles. Tangential and axial mean velocities and

velocity fluctuations were measured within the nozzle chamber by using Laser-Doppler-Velocimetry (LDV).

Outlet conditions were evaluated from the measurements and, in addition, photographs of the flow patterns

were taken. Part II deals with the drop behaviour in the spray cone of real nozzles. Axial and radial drop

velocities (mean values and RMS-values) and, simultaneously, drop sizes were measured at various nozzle

distances with a Phase-Doppler Particle Analyser (PDPA). Finally, some comparative considerations of

angular exit momenta (perspex models) and drop sizes (real nozzles) were drawn from the results,

exhibiting agreement between the two parts of the investigation. Also an energy balance for the nozzle as a whole was formulated and evaluated.

INTRODUCTION

Atomization with pressure-jet nozzles is widely used in mechanical engineering, e.g., for paint and

laquer spray nozzles, within spray towers, in humidifiers, but also in combustion engineering for liquid fuel

nozzles (domestic heating burners, industrial furnaces, gas turbine combustors). The great advantage of

spill-controlled nozzles compared to those without spill return is their much wider control range and the

relative constancy of spray quality over a wide throughput range. Their main drawbacl< is the lack of

equations for predicting drop sizes, velocities, etc. They still have to be developed, designed and

dimensioned empirically.

The swirl pressure-jet atomization process can be characterized as follows. The fluid enters the swirl

chamber of the pressure-jet nozzle through tangential channels. The energy necessary to create a free film

surface is taken from the pressure-induced kinetic energy of the fluid to be atomized. As a result of the

radial pressure gradient caused by the swirling motion (centrifugal forces) within the nozzle, an air-filled hollow core forms about the nozzle chamber axis. Hence, the liquid issues from the nozzle exit as a

concentric film which forms an expanding hollow cone, which finally disintegrates into an atomized spray as a result of inertia and viscous forces acting upon the film.

Obviously, the generation of the final drops is strongly influenced by the outlet conditions and the

liquid film properties at the orifice. Those in return depend on the flow fields inside the nozzle chamber.

Keeping those dependencies in mind, the motivation for stepwise investigations is evident. The first step

(PART I) consisted of flow field measurements inside the nozzle chambers of enlarged perspex models

(scale 20 : 1). In the second step (PART II), investigations in the spray cone of real nozzles were carried

out. In order to characterize the flow within the nozzle, and to compare the results obtained inside and

431
outside the nozzle, three dimensionless numbers were used:

Reynolds number Re =

swirl number

spllt.feed ratio

So =

V

RV=^

PART I : FLOW-FIELD INSIDE NOZZLE CHAMBER

Experimental Set-Up

forward feed

/

pressure—vessel

nozzle

flowmeters

spray chamber

pump storage tank

Fig. 1: Experimental set-up I

For the experimental investigation of

the flow-field inside the nozzle chamber

scaled-up perspex nozzles were mounted

into a pressure vessel, which was placed

above a large spray chamber as can be

seen in Fig. 1 . Vessel and spray chamber

had two windows each, so tiiat on the

one hand velocity measurements inside

the nozzle using Laser-Doppler

Velocimetry could be conducted, and on

the other hand photographs of the spray

cone could be taken.

A ternary test liquid was fed into the

nozzle chamber by a pump. One part of

the liquid stream left the nozzle through

the orifice, the other part escaped through

the central spill-bore in the rear wall of the

nozzle, flowing directly back to the

storage tank. The part of the liquid being

atomized was collected at the bottom of

the spray chamber from where it also

passed into the tank. The volume flow

rates of the feed and the spill were

measured and controlled by flowmeters.

Typical feed volume flow rates were in the

range between 10 and 90 l/min.

Air bubbles in the liquid stream

served as tracer particles for the LDVmeasurements. The bubbles were fed into

the test fluid by means of a valve at the

low-pressure side of the circuit.

Results

Fig. 2 exemplifies flow fields for different spilLfeed ratios (RV) at a constant feed rate of 35 l/min. In the lower left corner of each diagramm the values of the spilhfeed ratio (RV), the swirl number (Sq) and the

Reynolds-number (Re) are indicated. On the left hand side of each diagramm, the contour of the air-filled hollow core has been included, the corresponding dimensions being extracted from photographs of the

flow. The behavior of the core under the influence of various spiil:feed ratios is obvious : at a ratio RV = 0.0

the core diameter increases progressively towards the nozzle outlet, whilst for RV = 0.86 the core assumes

an almost cylindrical shape of very large diameter. On the right hand side of each diagramm tangential and

432
= 0.00

s. - 3.73

Re - 66250 85

Z [mm]

35 1/min

0 1/mln

1.46 m/s

V 2.2 est

RV = 0.43

s. = 3.73

Re - 66250 85

Z [mm

35 1/mln

15 1/mln

1.46 m/s

2.2 est

33\r40

RV = 0.86

s. - 3.73

Re - 66250 85

Z [mm

V,- 35 1/mln

30 1/mln

»*. 1.46 m/s

V 2.2 est

Fig. 2: Mean velocity profiles u,w; variation of spill:feed ratio

433
axial velocity profiles are illustrated. The

velocity components were measured at 6

different nozzle cross-sections each. The

plots represent the velocities as ratios of

the mean tangential inlet velocity through

the tangential ports. The axial components

are drawn in the downward direction, the

tangential components to the rear of the

drawing plane.

In all cases, the tangential velocity

profile resembles a Rankine vortex, with a

solid body vortex characteristic in the

vicinity of the nozzle axis, and tangential

velocity decrease towards the chamber

periphery. The tangential velocity profile

varies only slightly in the downstream

axial direction, the peripheral zone being

simply cut-off due to the convergent

nozzle geometry.

From the axial component profiles a

clear influence of the spill-return may be

observed. Whilst at low spill :feed ratios

two forward flow zones appear (one at the

nozzle periphery, the other in the vicinity

of the hollow core), which merge with

each other in the region of the orifice, at high values of RV (RV > 0.75) a clear

backflow along the hollow core is apparent for nearly the whole nozzle

chamber length. Near to the nozzle outlet

the values of tangential and axial velocity

components are rather similar.

In addition to the mean values of the

velocity components also the fluctuations

of the tangential and axial components

Fig. 3: Velocity fluctuations u',w' were measured. Again the spill :feed ratio

showed the strongest influence on the

profiles, as to be expected from the observations of the mean velocities. The variation of the swirl number

indicates a smoothening effect of swirl on turbulence. Measurements with a higher viscosity test fluid

resulted in lower fluctuations as can be seen in Fig. 3. Here again the velocities are plotted as ratios of the

mean tangential inlet velocity. From Fig.3 it is also evident that the fluctuations are always high in regions

with high gradients of the mean velocities (see Fig.2), i.e., the largest fluctuations appear at the periphery

and near to the hollow core, whilst they are damped in the region between those maxima.

PART II : SPRAY CONE INVESTIGATIONS

Experimental Set-Up

Measurements in the spray cone of real nozzles were conducted according to the arrangement as

schematically illustrated in Fig.4. Water was fed from a storage vessel by a piston pump to a spill-controlled

swirl presssure-jet nozzle which generated the droplets. The volume flow rate could be controlled and, in

addition, the nozzle pressure was metered for both the nozzle feed and spill rates. For the experiments, a

nozzle with an orifice of 1mm exit diameter was used which, at an applied pressure of approx. 10 bar,

delivered a volume flow rate of 40 l/h. Different standardized units (swirl generators and nozzle chambers)

RV = 0.43

s. - 3.73

Re - 66250 85

Z (mm

velocity fluctuation

35 1/mln

v.- 15 1/mln

w,- 1.46 m/s

2.2 est

RV - 0.43

s. - 3.73

Re - 10052

velocity fluctuation

35 1/mln

15 1/mln

1.46 m/s

V 14.5 est

434
tl Nozzle on coordinate slides

Fig. 4: Experimental set-up II

could be combined to a complete nozzle. Hence the spili:feed ratio, the swirl number and the nozzle

geometry were experimental variables also in this second part of the investigation.

The atomization occurred in free atmosphere. The spray configuration was of the hollow cone type

with a spray angle of approximately 80°. The spray was collected in the supply reservoir. In order to

reduce the influence of secondary backflowing mist to a minimum, the surrounding air was uniformly drawn

from the vessel by a suction fan. The nozzle was mounted on a 2-dimensional slide, thus allowing different

spray flow sections to be probed, following a 2-dimensionaI coordinate system.

The optical axis of a Phase-Doppier system (PDPA) was located right above the rim of the vessel.

PDPA delivers the probability density distributions of drop size (d) and drop velocity components (u,v).

Results

On the right hand side in Fig.5 drop velocity information is being presented. For three different nozzle

distances, the mean values of the axial velocity and the related RMS-values are shown. The velocity is rather low in the center, then increasing to relatively high values in the main spray cone shell. This velocity

maximum decreases with increasing nozzle distance, because the drops are decelerated by the

surrounding air. For detailed knowledge of the structure of the spray not only mean velocities but also RMS-values are needed which are responsible for the droplet dispersion process. There is a clear maximum of fluctuation

at the inner side of the main spray cone shell, which is the region of the largest velocity gradients. In the

center region the drops move slowly and exhibit low tubulence. On the left hand side of Fig.5 arithmetic

mean diameter {d^^) and Sauter Mean Diameter (SMD) are plotted. As from the velocities, also from the

drop sizes the hollow cone type of the spray becomes evident. In the center region a rather narrow size

distribution exists with a SMD of about 25 fjm. For comparison, the main spray cone shell exhibits a wide

distribution and a SMD around 200 jjm.

435
E

3.

E

3.

E

3.

Diameters and velocities

RV=0.5, Re=68307. 5^=3.68, Nozzle 1

SMD: Sauter meon diameter, d,„: arithmetic mean diameter

u : axial mean velocity, u': axial RMS

250^

200-^

150-

100

50-1

o-B SMD

nozzle

distance

25mm

250^0

200-

150-j

100-j

50-

' I' ' '

I ' 60 40 20

' '

I

'

'

'

I

' ' '

I

' ' 0 20 40 60 80

nozzle

distance

50mm

I'll

25q80 60 40 20

200^

150-

100

nozzle

SO'j distance

75mm

r15

-10

-5

\

E

15

-10

-5

(A

E

' '

I

'

I

'

I

' ' ' I

'

' Q 20 40 60 80^^

-10

-5

E

~i—I—

I—I—I—I—|~*i—r-i—I—1—I—I—

j—I—r—I—

I—I—r—i—

|—i—i—i—|—r i i 80 60 40 20 0 20 40 60 80

Radius [mm]

Fig. 5: Drop diameters and velocities

for different nozzle distances

Another aspect can be pointed out in Fig.5. Tlie maximum of the arithmetic

mean diameter (in the spray cone shell)

increases slightly with nozzle distance.

The spray cone aspirates surrounding air

into its center. The small drops follow the

air flow without slip, leading to decreased

concentration of smaller drops and,

hence, increased mean drop size in the

shell.

EXIT MOMENTUM AND SAUTER MEAN

DIAMETER

Figure 6 compares LDVmeasurements obtained for the liquid flow

inside the nozzle chamber of the perspex

model with drop size measurements in the

spray cone. In Fig.6a reduced angular

momenta (as evaluated from velocity

measurements) near to the nozzle exit are

plotted as functions of the spill :feed ratio

RV. For the low viscosity test fluid, i.e. high Reynolds number of the swirl

chamber flow, a minimum angular

momentum at the nozzle exit occurs in the range of a spill:feed ratio of about 0.4,

which could be expected to lead to coarser atomization in this range. Indeed,

this fact is being confirmed by the spray

measurements (Fig.6b). Here, the Sauter

Mean Diameter (SMD) is plotted as a

function of the radial position, varying the

spill :feed ratio. As expected from the

evaluated exit momenta, SMD values in the main spray cone shell (radial positions

between 30 and 50mm) at first increase

up to a maximum value at RV = 0.6, then

for higher spill :feed ratios decrease again.

Former measurements of drop sizes in the

spray cone with a Malvern Particle Sizer

showed the same trend. In Fig.6c

cumulative volume distributions are

plotted for different spillifeed ratios. Again

the drop size at first increases with

increasing spill:feed ratio up to a value of

RV = 0.5 and then decreases, indicating

that atomization improves for higher

spill :feed ratios. This observation is important when operating spill controlled

nozzles over a wide control range and will be a topic of major interest in future

investigations.

436
0.8-

E

D-t^C

0)E

o

E

>

on

D

o

a; 0.2-

0.0

high viscosity (14,5 cSt)

swirl number SO = 1 ,77

nozzle 2

nozzle 1

nozzle 2

—'

1

1 1' 1 '

I 0.0 0.2 0.4 0.6 0.6 1.0

spill/feed rotlo RV

Fig. 6a: Reduced angular exit momentum over spill:feed ratio

250 r—

1

E

3

1 1 Q 200-

to

i_

0) -f 150-

0)E

o

b 100-

c

o

<u :§

50-

0)

-4-' 3

O

(/)

nozzle G14

Re = 68307

So = 3,68

10 20

I 30

I 40

0—0 RV = 0.8 - RV = 0.6

Q-a RV = 0.4 e-e RV = 0

50 60

spray cone radius r [mm]

Fig. 6b: Profiles of Sauter Mean Diameter

437
1000

Particle size (jjni)

Fig. 6c: Cumulative volume distributions

ENERGY BALANCES

In this final section a sinnple energy balance is carried out, based on measured data from the nozzle

investigations.

The energy necessary for atomization is taken from the high-pressure volume flow to the nozzle, the

feed energy Ey. When operating the nozzle with spill-control, a portion of the energy flows into the spill

return, the spill energy Er. The kinetic energy in the feed and spill lines is negligible compared to the static

pressure energy. Part of the remaining energy Ey - Er is converted into kinetic energy at the nozzle oriffice

Enina, a higher portion being lost as a result of internal fluid flow friction inside the nozzle chamber. At a

nozzle distance h, a portion of E^ina is expended to create the surface of the new drops Eqi,, other one

delivers the kinetic energy of the drops E.^^, and a third is lost due to friction between the liquid and the

surrounding air. All this energy flow rates can be estimated from measurements. For the drop energies in

the spray it is necessary to integrate over the droplets of the entire spray cross section. All the equations

are shown in Fig. 7 . For a nozzle distance of 50 mm, the results of the estimated energy balance can be

summarized as follows; all the energy terms are being expressed as ratios of the initial energy Ey - Ef,:

- 10-30 % of the initial static pressure energy appear as kinetic energy flux at the nozzle orifice;

- 0.1-0.4 % remain as surface energy in the spray;

- 5-25 % can be identified as kinetic energy of the droplets in the spray.

438
g. 7: Nozzle scheme and energy terms

CONCLUSIONS

The results of the first part of the investigation, as

obtained from measurements inside enlarged perspex

models, may be interpreted as follows:

In all cases the radial profiles of the tangential

velocity resemble a Rankine vortex type distribution.

Most of the measured axial velocity profiles show two

maxima, one at periphery of the chamber and the other in the vicinity of the central hollow core. For very high

spill :feed ratios, a back-flow along the hollow core was

evident instead.

The RMS-profiles exhibit that the velocity fluctuations are always high in regions with high gradients of the mean

velocities.The variation of swirl number delivered a damping effect of swirl on the turbulence.

The second part of the study, which dealed with

spray cone investigations, yielded the following results:

A hollow cone structure of the spray was clearly

evident.

There were slow drops with low turbulence in the

center region of the spray, whilst in the spray cone shell

fast drops with high turbulence were found. Also, in the

center the drops were rather small and had a narrow size

distribution, but in the spray shell there were large drops

with a wide size distribution.

The maximum of the arithmetic mean diameter in the

spray shell increased slightly with the nozzle distance, to

be explained by the fact that the spray cone aspirates

surrounding air together with small droplets into its center.

From comparing calculated nozzle exit conditions

(perspex model) with drop size measurements in the spray

cone (real nozzle), agreement was found between a minimum of angular exit momentum and a maximum of

SMD occurring in the spill.feed ratio range of about 0.5.

An estimated energy balance for the droplet mass

was carried out. The calculated examples in the spray

delivered a remaining surface energy between 0.1 and

0.4% only, and a remaining kinetic energy between 5 and

25% related to the initial static pressure energy.

439
NOMENCLATURE

A tf^tol LUlal inlot liilcL sroQ dit^a

own 1 ui idi 1 luvi uidi 1 icic^i

arithmptip ai lil II 1 IwLlw mp^n 1 1 IwCll 1 Hiamptpr UlCII 1 ICLCI

•n Qnnnljir pvit mnrnpntiim

«n distnmilar i^uidi pntrv/ x^i III mr^mpntiim y II i\ji 1 id iLui 1 i

Unotir^ IMIIULIU onorn\/ c^ilciyy atdl Hictanr^o UloLdliUc^ KII F'^kin.a kinptip r\il ICLIV^ ^nf^THM Clld^y atCll nrifipp Wl IIIL^C F ciirfapo ouiidut; pnornw c;iimyy

spill GnGrgy

p iccU ciiciyy

IIh liU^^lc UloldiiUU

Pr spill pressure

Pv lesu pr6ssur6

oriTic6 rauius

licyiiUIUo llUlilUci

'exz iduius 01 cXuciurioiiy oi iric iiiici C/nanrici

RV ^nill'fppH oLJIll.lCCLi ratin 1 uLlvJ s OVVIII Qwirl niinnhpr 1 ILJI 1 SMD Saiitpr Mpan niflmptpr

• Ol^ili fl^\ill fo^A spill now raie

«Vv feed flow rate Wo tangential inlet velocity

w,u tangential and axial velocity components

w',u' tangential and axial velocity fluctuation components

V viscosity

p density

a surface tension

ACKNOWLEDGEMENT

The authors wish to express their gratitude to the Deutsche Forschungsgemeinschaft (DFG) for the

financial support of the investigations reported here, which have been performed in the frame of the joint

research project "Sonderforschungsbereich 167/ Hochbelastete Brennraume" of the Karlsruhe University.

440
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 46

MULTIPLE SCATTERING AND SIZE DISTRIBUTION EFFECTS ON THE

PERFORMANCE OF A LASER DIFFRACTION PARTICLE SIZER

T. Paloposki and A. Kankkunen

Helsinki University of Technology

Espoo, Finland

Abstract

The effect of multiple scattering on the performance of the Malvern Particle Sizer

was studied experimentally. Two particle samples with the same volume median diameter

but with different size distributions were used in the experiments. One sample could be

described by a log-normal distribution function, the other by a Rosin-Rammler distribution

function. The particle samples were specifically designed to simulate drop size distributions

in liquid sprays.

The model independent, log-normal, and Rosin-Rammler models were used in the

data analysis, and the results were compared. It was found that even when the sample

concentration was low and mtdtiple scattering did not occur, the resiilts obtained with the

log-normal and Rosin-Rammler models could be misleading, particularly in the high end

of the distribution (large particles). When the log-normal test sample was used, the model

independent and log-normal models gave good residts, but the Rosin-Rammler model

seriously underestimated the fraction of large particles. When the Rosin-Rammler test

sample was used, the model independent and Rosin-Ranm[iler models gave good results,

but the log-normal model seriously overestimated the fraction of large particles.

When the sample concentration is high and multiple scattering occurs, data analysis

with the model independent model is not feasible. Correction equations axe available

in hterature for data analysis with the log-normal and Rosin-Rammler models. It was

found that when the correction equations were used, both the log-normal and RosinRammler models provided the mean particle size reasonably well. However, there were

considerable difficulties in the high end of the distribution. This was not surprising, since

the experiments with a low particle concentration had already shown that the data analysis

with the log-normal and Rosin-Rammler models can produce misleading results.

441
Introduction

The Malvern Particle Sizer was developed by Swithenbank et al. [1] for the

measurement of drop size distributions in liquid sprays and is conmierciaUy manufactured

by Malvern Instruments Ltd. The instrument is based on the diffraction of laser light by

the drops. The light energy distribution in the near-forward direction is measured with

a multi-element detector and analyzed with a microcomputer. A non-linear least-squares

method is used to find the best-fit drop size distribution i.e., a size distribution which

gives the most closely fitting Hght energy distribution.

Swithenbank et al. assumed a Rosin-Rammler distribution of drop sizes and single

scattering of light by the drops. Negus and Azzopardi [2] discussed the validity of

these assumptions and tested the instrument against particle samples with known size

distributions and concentrations. The particle samples were made of glass spheres and

suspended in water in a stirred test cell with windows for optical access. They concluded

that some limitations arise from the assumption of a Rosin-Rammler size distribution

and estimated the critical particle concentration above which multiple scattering problems

become significant.

The assumption of a Rosin-Ranomler size distribution is no longer necessary, since

other models have been added to the data analysis software of the commercial instrument.

The user can now choose between three different distribution functions — Rosin-Raxomler,

normal, and log-normal — and a model independent distribution. AU three distribution

functions are two-paxameter functions i.e., the values of two independent parameters

specify the distribution. The model independent distribution is a presentation of volume

fractions in 16 discrete size classes; thus, there are 15 independent parameters. Obviously,

the model independent distribution is much more flexible than the two-parameter

distributions. Any distribution that can be described by a two-parameter distribution

function can also be described by the model independent distribution. Furthermore,

bimodal distributions and some other difficult distributions can be described by the model

independent distribution but not by any two-parameter distribution functions.

The two-parameter functions can stiU be useful in spray studies, since the drop size

distributions in liquid sprays are usually unimodal. However, the normal distribution has

seldom been applied in spray studies and wiU not be discussed here. The Rosin-Rammler

and log-normal distributions have been applied more frequently and wiU be studied in our

paper.

The multiple scattering problems have been widely investigated. A most practical

solution has been presented by Felton et al. [3]. They presented correction equations

which can be applied both to results obtained using the Rosin-Rammler model and to

results obtained using the log-normal model. The correction equations were derived

using a theoretical model of multiple scattering, and model predictions were compared to

experimental results obtained using dense suspensions of glass beads in water. Excellent

agreement was obtained. Felton et al. also found that multiple scattering effects were not

significant for obscurations less than 50 %; obscuration is the fraction of light diffracted

by the drops and is recorded in each measurement.

Felton et al. did not present correction equations for restdts obtained using the

model independent model. The correction equations for residts obtained using the Rosin442
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 48

TIME-BASED POLARIZATION RATIO MEASUREMENTS OF DROPLET SIZE

J.R. Zurlo, C. Presser and N. Chigier*

Chem. Sci. and Tech. Lab., National Institute of Standards and Technology, Gaithersburg, MD, U.S.A.

*Camegie Mellon University, Pittsburgh, PA, U.S.A.

ABSTRACT

Nearly instantaneous droplet size measurements were carried out using an ensemble light scattering

technique based on measurement of the polarization ratio. Measurements of droplet size were performed in a pulsed impinging diesel spray and in a steady hollow-cone spray. The measurements in the pulsed diesel

spray show a change in size with time, with the largest droplets at the beginning of the spray. Measurements

of droplet size and number density in the hollow-cone spray show changes in both droplet size and number

density with time. A method of quantifying measurement error of average droplet size under conditions of

low number density was demonstrated.

INTRODUCTION

Knowledge of the droplet size in sprays is important for spray combustion, pesticide and herbicide

spraying, spray cooling, fire sprinklers and many other industrial applications. Time-based measurements of

droplet size and number density are important for pulsed sprays. Not as obvious is the need for time-based

measurements in "steady" sprays. High speed movies of the atomization process have shown that "steady"

sprays are not uniform in time but can have clusters of droplets. This droplet clustering may have significant

ramifications in combustion applications for soot production. A rapid change in spray angle, called fluttering,

has been observed in airblast atomizers [1]. Quantitative measurements of droplet size and number density

with time are difficult to obtain with commercially available instruments.

Techniques commonly used to determine droplet size in sprays include laser diffraction [2], phase/Doppler

interferometry [3] and light-intensity deconvolution [4]. The laser diffraction particle sizer is a line-of-sight

measurement with a maximum samphng rate of 45 Hz. The phase/Doppler interferometer and light intensity deconvolution instrument are single particle counters in which individual droplets are detected as they

pass through the measurement volume. The phase/Doppler interferometer has a maximum samphng rate of

125,000 particles/s. The intensity deconvolution technique does not provide temporal measurements due to

mechanical switching of the measurement beam that is necessary to measure the full range of droplet sizes.

One ensemble technique, which is based on measurement of the polarization ratio, has been used successfully to measure droplet size in sprays [5]. Its advantages include insensitivity to beam steering, excellent

spatial resolution, rapid response time, and insensitivity to multiple scattering [6]. Its disadvantages include a dependence on the index of refraction and the need to assume a form for the size distribution. During the

past several years, various approaches have been used to measure the polarization ratio. Details of these

optical configurations are described elsewhere [5, 7, 8, 9].

EXPERIMENTAL SETUP

Two different configurations of the droplet sizing instrument were used in these experiments. The first arrangement consists of a randomly polarized helium-neon laser light source and two photomultipher tubes

mounted on an optical bench. Light from the laser passes through the spray and the ratio of horizontally to

457
vertically polarized light scattered at 105° from the forward direction is collected to determine the Sauter

mean diameter of the spray. The scattered light is collimated with a lens and then passes through a

focusing lens, a circular aperture (which reduces the measurement volume to approximately 1 mm^), a narrow band pass filter and a 50/50 beam splitter cube, which produces two beams. Each beam passes

through a linear polarizing filter and onto a photomultiplier tube. The polarizing filters are oriented normal

to each other to provide a vertical and a horizontal component of the scattered light. The output current of

each photomultiplier tube is amplified, and then measured with a microcomputer-based digital oscilloscope

peripheral. The data acquisition system has four 8-bit analog-to-digital converters that allow simultaneous

sampling of all channels at a rate of 500 kHz.

The pulsed spray was generated by an experimental diesel atomizer [10] impinging on a flat plate. The

measurements were performed at ambient pressure and temperature. More details of the experimental setup

are available in reference [11].

For measurements in the hollow-cone spray, an argon-ion laser was used as the light source. A schematic

of the setup is shown in Fig. 1. Since the output of the laser is polarized, a different optical arrangement was

used. The blue (488 nm) beam was used for the measurement. A polarization rotator was used to rotate

the polarization of the laser beam through 45° . The beam was directed through the spray and the scattered

light was collected with a lens and directed to a 50/50 beam splitter. Half of the light went through a narrow band pass filter, a horizontal polarization filter and then onto a photomultiplier tube, while the other half

of the light went through a narrow band pass filter, a vertical polarization filter and then onto the other

photomultiplier tube. The use of a 45° linearly polarized, or a circularly polarized laser beam is better than

the use of a randomly polarized laser because of the possiblity of "polarization slipping" with a randomly

polarized laser. The data acquisition system was upgraded to a 1 MHz, 12-bit system because of the wider

dynamic range encountered in the steady spray.

The steady spray measurements were performed on a kerosene spray generated by a hollow-cone pressureswirl atomizer of nominally 60° spray angle. The measurements reported in this paper were at a location 3

cm downstream and 1.5 cm off" the spray centerline. This position is on the outside edge of the spray cone. LIGHT SCATTERING CALCULATIONS

Since the form of the droplet size distribution is assumed, Mie scattering calculations can be performed

to determine the polarization ratio as a function of the Sauter mean diameter (SMD). The Sauter mean

diameter is commonly chosen to characterize atomization quality for many applications, and is defined as: irp(D)D3^

(1)

where D is the droplet diameter and P{D) is the probability distribution function. A log-normal size

distribution was chosen because it is often used to represent droplet distributions in sprays. To characterize a collection of particles with a log-normal size distribution requires two parameters, the geometric mean

diameter Dg, and the geometric mean standard deviation, CTg. For all calculations CTg was set equal to 0.3. The probability distribution function for a log-normal size distribution is expressed as:

{HD/Dg)r P{D) =

y/2TTcrgD

exp (2)

The polarization ratio is defined as:

T =

Qhh

Qvv

(3)

where Qw and Qhh are two elements in the Mueller matrix. The relationship between the Mueller matrix

and the Stokes parameters , [/„, Ih, U, V], is given by:

Qvv Qvh

Ih 1 Qhv Qhh

u 0 0

V 0 0

0

0

i(5i5^-f-525?)

0

0

-|(525? - -2(5i52 -

ho

ho

Uo

(4)

458
where k = 27r/A and r is the distance from the scattering volume. is the intensity of the vertically

polarized light, similarly Ih is the intensity of the horizontally polarized light. The 0 subscript on the Stokes

parameters indicates incident light. Si and 52, the scattering amplitudes, can be calculated using Mie

scattering theory. and are the complex conjugates of the scattering amplitudes. The elements of the

Mueller scattering matrix that are used for droplet sizing are:

Qvv — Si (5)

and

Qhh = Si (6)

Figure 2 shows the relationship between polarization ratio and Sauter mean diameter for a scattering angle

of 105°, an index of refraction of 1.45 (dodecane), and a laser wavelength of 632.8 nm. Figure 2 shows the

high degree of sensitivity of the polarization ratio for particle sizes less than 15 //m; for droplet sizes larger

than 40 ^m, there is little sensitivity.

Travereo

BB - Beam block F Inlerierence liKof L - Lens M- Mrnr PA - Polafizalion analyzer PH - Pin holo PMT - PholomuH. lube deledor PR - Polartzalion folalor RO • Receiving optics SF • Spatial tmer

Mcro- A/D oonv

-I r

30 45

SMD (fim)

Figure 1: Schematic of the polarization ratio particle Figure 2: Calculated value of polarization ratio versus

sizer used for measurements of the hollow-cone spray. Sauter mean diameter for 105° scattering angle, an

index of refraction of 1.45, and a wavelength of 632.8

nm.

Qyh and Qhv are zero for single scattering from an ensemble of spherical particles. The combination

of the four Stokes parameters, Iv,Ih, U, V, and the Mueller matrix are useful for describing light scattering

from any object. If the Mueller matrix of an object is known, then calculating the interaction of a known

light source, [IvO, Iho,Uo,Vo], with the object is a simple matter of matrix multiphcation. An unpolarized

light source of unit intensity has Stokes parameters of [1,1,0,0]. Use of an unpolarized incident light source

allows the immediate determination of two elements of the Mueller matrix, Qw and Qhh since /„ and Ih are measured. A vertically linearly polarized light source of unit intensity has Stokes parameters of [1,0,0,0].

Use of a vertically linearly polarized light source allows one to verify that one of the cross components of

the Mueller matrix (Qhv) is zero. Similarly, using a horizontally linearly polarized light source allows one to

verify that Q^,h is zero. Once the droplet size has been determined, the droplet number density can be obtained using Eq. 7.

N = (7)

459
When taking measurements at high data acquisition rates, it is necessary to verify that sufficient numbers

of droplets are in the measurement volume to ensure a representative sample. As the number of particles in

the measurement volume decreases, the accuracy with which the average droplet diameter can be determined

decreases. With the assumption that the droplets in the spray have a log-normal size distribution, varying

size groups of droplets were generated using a random number generator. The uniform random number

generator was transformed to a log-normal deviate, so that given an infinite number of random numbers,

a perfect log-normal distribution would result. The Sauter mean diameter and polarization ratio were

calculated directly by summing the contribution from each individual droplet in the group of droplets. As

the number of droplets in each group increased, the variation in the calculated polarization ratio decreased.

Figure 3 shows calculations for 200 droplets per group. Figure 4 shows calculations for 1000 droplets per

group.

Figure 3: Polarization ratio versus Sauter mean di- Figure 4: Polarization ratio versus Sauter mean diameter calculated using 200 droplets per group ameter calculated using 1000 droplets per group

The standard deviation of polarization ratio for number of droplets and Sauter mean diameter was computed using the randomly generated log-normal distribution. This allows one to generate error bars on

the Sauter mean diameter for a given polarization ratio and number of droplets in the measurement volume.

Figure 5 shows the error bars for one standard deviation for a measurement volume that contains 1000

droplets. As expected, the measurement error for Sauter mean diameter increases as the polarization ratio

decreases. Figure 6 shows the error bars for one standard deviation for a measurement volume that contains

200 droplets. The measurement error increases with decreasing number of droplets in the measurement

volume. The standard deviation of polarization ratio, S, can be determined from:

367J8

((n + 237.206) *(5MD-F 5.3))' ^ ^

where n is the number of droplets in the measurement volume and SMD [//m] the Sauter mean diameter.

Equation 8 is only valid for a laser wavelength of 488 nm, index of refraction of 1.45, and a scattering angle

of 105° . The computations would have to be repeated for any change in the above variables.

460
EXPERIMENTAL RESULTS

Before accurate measurements of number density can be made, the polarization ratio droplet sizing

instrument must first be calibrated. Other researchers have used scattering from propane as an absolute

standard. The time-based polarization ratio measurements were not sensitive enough to detect scattering

from propane so another calibration technique was needed. Latex spheres of known size suspended in

distilled water provided that standard. The previous, more sensitive system [12] was used to measure the

number density of latex spheres and then the latex spheres were used as a secondary calibration standard

for calibrating the time-based polarization ratio measurements.

The polarization ratio particle sizer was used to make measurements of droplet size for the portion of a pulsed diesel spray that impacted on a plate. The impaction plate was wiped clean of any liquid between

runs. Data was taken at a radial distance of 2.5 cm from the spray impaction centerline and a height of 7

mm above the plate. This distance is sufficiently far from the impaction zone so that the spray is moving

radially outward parallel to the plate. The linear depolarization ratios were measured to verify that spherical

droplets were present in the measurement location. For measurements of linear depolarization ratio, the laser

light was made alternately vertically or horizontally polarized by placing a linear polarization filter in front

of the laser. Measurements of both vertical and horizontal linear depolarization ratio were made at the

measurement location. The vertical and horizontal linear depolarization ratios remained below 0.2 for most

of the measurement time indicating that significant depolarization is not present. Figure 7 shows the Sauter

mean diameter as a function of time.

The Sauter mean diameter does not remain constant with time at this position. The initial Sauter mean

diameter is larger than at later times. This transient in droplet size lasts for about 0.1 ms and is attributed

to the increased momentum-to-drag ratio of the larger droplets. A range of droplet sizes is produced by the

impaction of the spray on the plate. Since the larger droplets have a higher momentum to drag ratio, the

effect of aerodynamic drag is less pronounced and these droplets arrive first at the measurement location.

Figure 8 shows the Sauter mean diameter and number density of the hollow-cone spray as a function

of time. Evidence of droplet clustering can be seen by noting the large change of number density with time.

Figure 9 shows a portion of Fig. 8 with the error bars for one standard deviation of polarization ratio added.

As expected, for times with high number density and small droplet size, the error bars are small and for

low number density and large droplet size the error bars are larger. This illustrates a trade off between

large measurement volume and high accuracy, and small measurement volume and low accuracy inherent in

time-based ensemble techniques.

The evenly spaced measurements in time that this technique generates are amenable to statistical

analysis. Figures 10 and 11 show autocorrelations and cross-correlations calculated from measurements in

the hollow-cone spray. The autocorrelation plot can be used to look for specific frequencies in the spray. The

cross-correlation plot shows that for this location in the spray there is an inverse relationship between the

measured Sauter mean diameter and the measured droplet number density. It is also possible to produce

power spectral density plots from the measurements. This type of information would give valuable data

about the basic atomization process and useful practical information for spray nozzle designers. It is however,

difficult at this time to use these preliminary results to provide quantitative information on spray structure.

CONCLUSIONS

Modifications to the polarization ratio particle sizing technique allow nearly instantaneous measurements of

Sauter mean diameter and number density. Measurements were performed in a pulsed spray and in a steady

hollow-cone spray. The measurements in the pulsed spray show the expected result of the larger droplets at the spray tip, followed by smaller droplets. Measurements in the steady hollow-cone spray show that there

are changes in both average droplet size and number density with time. At the one measurement location

there was an inverse relationship between measured Sauter mean diameter and measured droplet number

density. Quantification of measurement error of average droplet size under conditions of low number density was demonstrated.

ACKNOWLEDGEMENTS

One of the authors (JRZ) would like to acknowledge the support of the Postdoctoral program at NIST,

which is administered by the National Research Council.

461
REFERENCES

Mao, C.-P., "Fluttering in Airblast Atomizer Sprays", ILASS-Americas, Proceedings of the 2nd Annual

Conference (1988).

Swithenbank, J., Beer, J.M., Taylor, D., Abbott, D., and McCreath, G., "A Laser Diagnostic Technique

for the Measurement of Droplet and Particle Size Distribution", in Experimental Diagnostics in Gas

Phase Combustion Systems, (B. Zinn, Ed.), Progress in Astronautics and Aeronautics, Vol. 53, pp.

421-447, AIAA, New York, NY (1977).

Bachalo, W.D. and Houser, M.J., "Phase/Doppler Spray Analyzer for Simultaneous Measurements

of Drop Size and Velocity Distributions", Optical Engineering, Vol. 23, No. 5, pp. 583-590, September/October (1984).

Holve, D.J. and Annen, K., "Optical Particle Counting and Sizing Using Intensity Deconvolution"

,

Optical Engineering, Vol. 5, No. 23, pp. 591-603, September/October (1984).

Beretta, F., Cavaliere, A., and D'Alessio, A., "Drop Size and Concentration in a Spray by Sideward

Laser Light Scattering Measurements", Comb. Sci. and Tech., Vol. 36, pp. 19-37 (1984).

Beretta, F., Cavaliere, A., and D'Alessio, A., "Experimental and Theoretical Analysis of the Angular Pattern Distribution and Polarization State of the Light Scattered by Isothermal Sprays and Oil

Flames", Combustion and Flame, Vol. 49, pp. 183-195 (1983).

Beretta, F., CavaHere, A., and D'Alessio, A., "Ensemble Laser Light Scattering Diagnostics for the

Study of Fuel Sprays in Isothermal and Burning Conditions", Twentieth Symposium (International) on Combustion, pp. 1249-1258. The Combustion Institute (1984).

Presser, C, Gupta, A.K., Santoro, R.J., and Semerjian, H.G., "Droplet Size Measurements in a Swirling

Kerosene Spray Flame by Laser Light Scattering", ICLASS-85, pp. VIIC/2/1-13 (1985).

Hodges, J.T., Baritaud, T.A., and Heinze, T.A., "Planar Liquid and Gas Fuel and Droplet Size Visualisation in a DI Diesel Engine", SAE 910726 (1991).

Sinnamon, J.F., Lancaster, D.R., and Steiner, J.C., "An Experimental and Analytical Study of Engine

Fuel Spray Trajectories", SAE 800135 (1980).

Zurlo, J.R. and Chigier, N., "Impinging Diesel Spray Dynamics", Atomization and Sprays, Vol. 1, No.

3, pp. 303-318 (1991).

Presser, C, Gupta, A.K., Semerjian, H.G., and Santoro, R.J., "Application of Laser Diagnostic Techniques for the Examination of Liquid Fuel Spray Structure", Chem. Eng. Comm., Vol. 90, pp. 75-102

(1990).

462
60

50

E

^30

a

w 20

10

0

0.5 1 1.5

Polarization Ratio

' ' ' «,f t y a «,> > 0.5 1 1.5

Polarization Ratio

Figure 5: Error bar plot for 1000 droplets in the mea- Figure 6: Error bar plot for 200 droplets in the measurement volume. surement volume.

1.75 2.25 2.75

Tlm« (mi)

3.25

Figure 7: SMD as a function of time. Data taken at a radial distance 2.5 cm from the diesel spray centerli

and a height of 7 mm above the plate.

me

E

en

1 0'

4 5 6

Time ( m s )

Figure 8: Sauter mean diameter and number density as a function of time for the hollow-cone spray.

463
30

Figure 9: Sauter mean diameter and number density as a function of time for the hollow-cone spray. Error

bars for one standard deviation of polarization ratio are shown.

200 300

Lag

500

Figure 10: Autocorrelation of number density for the hollow-cone spray.

Figure 11: Cross-correlation of SMD and number density for the hollow-cone spray.

464
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 49

GENERATION OF SPHERICAL AND MONODISPERSE POLYMER

PARTICLES BY ATOMIZATION OF MONOMERS OR DISSOLVED

POLYMER PRECURSORS

T. Panagiotou and Y.A. Levendis

Department of Mechanical Engineering

Northeastern University

Boston, MA, U.S.A.

ABSTRACT

A technique to produce spherical and monodisperse particles of selected plastics is presented. Liquid precursors of either mixtures of organic monomers and initiator catalysts, or polymers dissolved

in organic solvents were sprayed inside a vertical thermal reactor. The temperature range in the

reactor was 400-670 K and the experiments were conducted in a nitrogen atmosphere. Atomization

was achieved by an acoustically excited aerosol generator. Batches of equal size particles of two

thermoplastic materials, poly(styrene) and poly(methyl methacrylate), were obtained in the range

of 30-60 )um in diameter. The thermal environment, the atomization conditions, and the residence

times the particles experienced in the reactor were explored using numerical techniques; residence

times in the order of 10 s were estimated.

INTRODUCTION

Production of small, homogeneous polymer particles of controlled size are of technological interest to the plastics, catalyst and others industries[l]. Moreover, if the particles are produced

spherical and monodispersed, they can find further applications in fundamental studies of material

properties, calibration of instruments etc. Commonly, monodispersed particles have been produced

from polymer colloids by emulsion or suspension polymerization. Everett and co-workers[2,3] have

discussed the preparation of near spherical, 100 /im, poly(vinyl chloride) (PVC), poly(vinylidene

chloride) (PVDC) and submicron poly(acrylonitrile) powders from dispersion polymerization. A

similar technique was followed by Pendleton et al.[A] to produce submicron (0.1 ^m) PVDC particles. Howard and Knutton[5] prepared large spherical carbon particles (ftiSOO /im) from the thermal

degradation of cross-linked polymeric beads, made from the suspension polymerization of mixtures

of PVDC and ethylene dimethylacrylate. These materials yielded porous beads upon carbonization.

Bhanti et a/.[6] generated poly(styrene) particles by nebulizing a polymer solution in xylene

and evaporating the solvent. Subsequent size-classification and second stage atomization after

resuspending the colloids in water was necessary to produce a monosized aerosol. Matijevic and

co-workers[7] produced fairly uniform, micron size poly(urea) particles by reacting liquid aerosol

droplets of toluene diisocyanate with ethylenediamine vapor. The same aerosol technique was used

for polymerization of styrene monomer droplets in vapors of the initiator trifluoromethanesulfonic

465
INLET

VIRTUAL IMPACTOR'-' COLLECTION OF

PARTICLES

Figure 1: Schematic of the acoustically excited aerosol generator and the two stage thermal reactor

used for the production of spherical polymer particles.

acid (TFSA)[8]. The above investigators also produced (8//m) poly(divinylbenzene) particles using

the same aerosol technique[9]. These particles shrunk to porous 3-5 spheroid residues upon

carbonization at 500° C [10].

An alternative technique was used by Levendis and Flagan[ll,12] to produce monodisperse

solid carbon particles (5-200 //m) to be used in combustion studies. The method utilized an aerosol

generator to atomize various mixtures of poly(furfuryl alcohol) and pore-forming agents dissolved

in acetone. The mixtures were sprayed inside a thermal reactor where secondary polymerization

(curing) took place. The work described herein pertains to the production of poly(styrene) and

poly(methyl methacrylate) particles using the above technique to either atomize solutions of these

two polymers in solvents, or atomize and polymerize their liquid monomer precursors. In both cases

solid polymer particles were formed in the thermal reactor. In the latter case the polymerization

was conducted and completed inside the reactor. In this manner the rate of polymerization of the

two thermoplastics was expedited and plastic powders were produced in a matter of seconds.

EXPERIMENTAL APPARATUS AND PROCEDURE

(A) Aerosol Generator. In order to produce spherical and monodisperse drops from a polymeric

fluid an aerosol generator was constructed based on the design of Levendis and Flagan[12], shown

in Fig. 1.

466
In this instrument a continuous flow of liquid is forced through a small orifice. Concurrently

the liquid in the generator's cavity is oscillated at a high frequency using a piezoelectric transducer.

The vibration creates an instability which leads to the break-up of the liquid jet into uniformly sized

droplets. The size of the droplets is controlled by the size of the orifice, the oscillator frequency

and the liquid feed rate. Consequently, the size of the dried particle depends on all of the above

parameters in addition to the degree of dilution of the polymer (mass fraction of solvent). In the

present experiments in order to produce particles of sizes in the range of 30 - 60 //m, monomers or polymers under various degrees of dilution were used in conjunction with 75 fiva orifices, frequencies

of 16 -18 kHz, and a flowrate of 1.6 cc/min. The piezoelectric elements used were silver electrode

bimorphs {Vernitron PZT5H a,nd Piezo Kinetics PZT-550) 1.25 cm in diameter, 0.62 mm thick.

They were glued with a conductive silver epoxy ( Tra-Con) on thin stainless steel metallic membranes

25-75 /im (thickness). These transducers were driven by a function generator at 20 V amplitude.

The pinholes were laser drilled for precision {Melles Griot). A high pressure syringe pump

{Harvard Apparatus 909) was used to feed liquid polymers to the generator via {Becton Dickinson)

plastic syringes.

(B) Thermal Reactor. A thermal reactor was constructed for the drying and curing of the aerosol

droplets that were generated at the top of the reactor, where the aerosol generator was mounted,

Fig. 1. This stainless steel flow reactor is 1.5 m high, 0.1 m i.d. and is externally heated in two

stages by cylindrical (Thermocraft) heating elements. They are controlled by separate (Omega)

temperature controllers and they supply 3800 W at the top and 3000 W at the bottom sections

which provides waU temperatures up to 1000 K. The aerosol generator is placed at the top of the

reactor on a special flange that incorporates provisions for introduction of dilution gas (nitrogen).

This gas passes through a flow straightener and flows downwards engulfing the stream of droplets.

A secondary stream of gas fiows through a small orifice concentric with the aerosol jet. Fig. 1, to

disperse the droplets and prevent coagulation.

A quartz observation window is mounted on the top to facilitate monitoring of the operation

and uniformity of the jet. To check for monodispersivity of the droplets provisions for a small gas

jet to impinge perpendicularly to the droplet stream have been made. In this manner droplets

of equal size deflect at the same angle meanwhile polydisperse droplets deflect at a multitude of

angles (satellites). Monodispersion can be accomplished by adjusting the flowrate and finely tuning

the frequency. Theoretically, the frequency range for monodispersivity is < / < ssD where

Dj is the orifice diameter[13]. Directly below the observation window an orifice plate has been

inserted to improve the flow conditions, as discussed in a following section. At the bottom of the

reactor an virtual impactor stage has been constructed to aid in the collection of dry particles and

eliminate small micron-size particles formed by various mechanisms such as condensation of vapors.

A sampling filter stage has also been incorporated.

Production of Polymers.

In order to produce the polymers required for this study two different techniques were employed: (i)

aerosol production of dissolved polymers with subsequent evaporation of the solvent in the reactor,

and (ii) aerosol production of partially polymerized precursors in a solvent with subsequent evaporation of the solvent and sequential completion of polymerization in the reactor. Both techniques

can also be applicable to any other polymers that can be completely dissolved in solvents or to

monomers whose polymerization can be achieved in bulk and can be readily controlled.

467
Figure 2: SEM micrographs of polystyrene particles, (a) xlOOO and (b) x500.

(A) Poly(styrene). Polystyrene particles were produced with both of the above techniques:

1. An aerosol of commercially available polystyrene (Styron 685-D, Dow Chemical) dissolved in

toluene was generated. Solutions of 1 part of polystyrene in 40 parts of toluene were used in

conjunction with a 75/im orifice, a flowrate of 1.6 cc/min and an excitation frequency of 17 kHz.

The gas (nitrogen) flowrates in the reactor were « 1 1pm dispersion and 2.5 1pm dilution. The

maximum wall temperature at the upper section of the thermal reactor was 300 °C and that of the

lower section was 400 °C. Collected particles were smooth spheres having a diameter of « 33//m as deduced by optical and SEM microscopy, Fig. 2 a and b. Any larger particles formed by fusion

of two or three particles while still liquid (doublets, shown in the left side of Fig. 2 b, or tripplets)

were separated out by subsequent sieving. Dissolution of styrene beads in toluene required a few

hours at room temperature and sonication prior to atomization was necessary to achieve uniform

mixing. Insufficient mixing resulted in bubble formation. Furthermore, solutions thicker than

30:1 were too viscous to handle properly and produce monodisperce particles. Extensive testing

for monodispersivity of the droplets was done at the onset of each experiment using the test jet,

described before. Optimum conditions for the reactor temperature and flow characteristics were

determined by a combination of numerical modelling and experimental trials, with the following

criteria in mind:

(i) To achieve vaporization of the solvent slowly enough to avoid formation of bubbles.

(ii) To ensure that the residual polymer phase remains in the molten state to enable formation of

spherical particles.

(iii) To enable solidification of the particles, in the cooler bottom section of the reactor, before

they reach the collection stage.

Overheating should be avoided since it might alter the chemical composition by pyrolyzing the

polymer.

2. Production of polystyrene particles of w60 //m was achieved by generating an aerosol of partially

polymerized styrene monomer. The monomer was mixed with benzoyl peroxide (2% by weight)

and was heated at 83-84°C, in a nitrogen atmosphere, for 30 minutes. Subsequently, the semipolymerized precursor was dissolved in toluene at a ratio of 1:1 and passed through the thermal

reactor in nitrogen atmosphere. Production was conducted at the same conditions as before with

the exception of temperature, which was optimized for the current situation. The temperature

468
Figure 3: SEM micrograph of poly-methyl-methacrylate particles produced in the lab from the

monomer precursor (x800).

profile in the reactor was controlled to prevent flash-vaporization of the monomer. Thus, the top

section of the reactor was set at a low temperature (240°C) to provide slow initial heat-up, and the

bottom section was hot enough (300° C) for the polymerization reactions to proceed to completion.

(B) Poly(methyl methacrylate). Batches of Poly(methyl methacrylate) particles of 30-60/im in

diameter were produced by the above two techniques at similar conditions.

1. Solutions of 2% poly(methyl methacrylate) in acetone were atomized by the aerosol generator in conditions similar to the ones used for the poly(styrene) particles.

2. Methyl-methacrylate was pre- polymerized in bulk using benzoyl peroxide (2% by weight) as

initiator. The polymerization was carried out at 75° C for 15 min resulting in a partially reacted,

medium viscosity liquid. This was then dissolved in acetone at a ratio of 2:1 (acetone to polymer)

and conducted to the aerosol generator. The polymerization was then completed in the thermal

reactor, with wall temperatures of 260 °C and 290 °C at the top and bottom sections, respectively.

Solid particles, 45/xm in diameter, were collected, shown Fig. 3.

NUMERICAL SIMULATION OF THE PARTICLE HEAT TREATMENT

To model the velocity and temperature environment in the thermal reactor the "FLUENT" software package, (Creare, 1989) was utilized. Numerical solutions for the gas phase were obtained by

simultaneously solving the continuity, momentum and energy equations. The program was used to

study the effects of the reactor geometry, gas flow rates and wall temperature on the velocity and

temperature proflles. Furthermore, an additional feature of this program dealt with the behavior

of liquid drops introduced in the gaseous environment. Thence, the trajectory of the drops, the

time required for solvent vaporization, the residence time of the solid particle, and the particle

heat-up characteristics were estimated. Results so obtained were used to find the optimum reactor

temperatures for droplets of the polymer-solvent mixtures investigated herein. Various droplet sizes

were modelled in the range of 30-144 /im in diameter.

A brief description of the calculation with FLUENT is given in the following: A finite difference

grid is set up which divides the domain of the problem into a number of computational cells. Boundary

conditions are specified at the inlet and walls of the domain. The equations of continuity, momentum and

energy for the gas phase are reduced to a set of simultaneous algebraic equations and an iterative scheme is

469
used to find the solution of the system. When the solution of the gaseous phase converges sufficiently, the

introduction of droplets takes place and the mass, heat and momentun transfer between the gaseous phase

and the droplets are calculated. The iterations continue until the solution converges.

Input to the model were the following:

• The geometry and the dimensions of the reactor. The overall length was 1.6 m and the inner

diameter 0.1 m.

• The inlet velocities of the dispersion and the dilution gas streams. Their inlet flowrates were

1.0 1pm and 2.5 1pm respectively; the corresponding velocities were calculated to be 0.0055

m/s and 7.46 m/s.

• The wall temperature of the reactor, which was estimated by measurements taken by four

thermocouples attached to the outer surface of the reactor at different heights. The temperature between the thermocouples was assumed to change linearly. The maximun and minimun

wall temperatures were set according to the experimental conditions used for the generation

of the various particles.

• The flow type, which was considered turbulent because of the vigorous mixing of the high

velocity dispersion jet with the slow moving dilution gas. Turbulence was assumed despite a low Reynolds number, {Re = 25), calculated far away from the mixing zone (at mid-reactor

height).

• The initial size of the droplets, as calculated from the formula[13]: Dd = (^)^^^5 where Dd

is the initial droplet diameter, Q is the volumetric flowrate of liquid forming the aerosol and

/ is the imposed frequency. For the values of Q and / mentioned earlier, the initial droplet

diameter was calculated to be 144 //m. The corresponding initial axial velocity of the droplets

[u = -^^1 where Do is the diameter of the orifice) was calculated to be 6.06 m/s. Furthermore,

to investigate the behavior of drops at various radial positions in the reactor, initial radial

velocity ranging from 0.1 m/s to 0.5 m/s were input. The droplets consisted of a volatile

solvent (toluene or acetone) which evaporated and diffused into the reactor gaseous phase,

and a solid content (2.5% by weight) which heated-up but did not evaporate, simulating the

polymer. It should be noted here that the devolatization of the polymer was not taken into

account in the present study.

• Finally, the physical constants of the gaseous (nitrogen) and the droplet (predominantly solvent) phases were input as a function of temperature. Since the model assumes average

properties for the two constituents of the droplet, the solid particle diameter (after the va- porization stage) was overestimated because the density was biased towards that of the more

abundant solvent. Corrections for the real polymer density had to be applied.

Results of the Numerical Modelling

The theoretically predicted thermal environment inside the reactor is shown in Fig. 4 for the case where the maximum wall temperatures were 400° C (673 K) and 300° C (573 K) for the upper and

lower stages, respectively. The axial velocity profiles along the reactor are depicted in Fig. 4a.

Both the dispersion and the concentric dilution streams are shown with the former diffusing

relatively fast in the slow moving dilution gas stream. The presence of the orifice plate situated

downstream of the jet inlet, was found helpful in preventing boyant gases from rising, along the

walls, to the very top of the reactor. The presence of these hot gaises in close proximity to the

470
Figure 4: Results of numerical simulation for gas flow in the thermal reactor (a) axial velocity

profiles and (b) axial velocity contours (c) flow streamlines and (d) temperature profiles. Maximum

wall temperatures were 400° C (673 K) and 300° C (573 K) in the upper and lower stages. Wall

temperatures and centerline temperatures (in parentheses) are shown on Fig. d. (e) Results

of numerical simulation for the droplet-particle phase showing tracks of equal size droplets (144

fim) having different initial radial velocities (0.1-0.5 m/s). Asterisks denote completion of solvent

evaporation.

471
Figure 5: Results of numerical simulation depicting (a) particle size and (b) particle temperature

as a function of time for initial droplet sizes 60 ( ) 100 ( ) and 144 ( ) fim. The solvent

was toluene. Conditions corresponding to Fig. 4.

inlet of the droplets can cause rapid evaporation of the solvent and result in bubble formation, as mentioned before; moreover, as hot gases condense on the cooler surfaces of the observation window

obstruct the view to the jet.

In Figs. 4b and c the contours of the axial velocities and the streamlines are shown. It can be

observed that the presence of boyancy highly disturbs the flow pattern and induces circulation of

gases near the walls of the reactor. Thus, in addition to the presence of the orifice plate, it was

determined that application of suction at the exhaust of the reactor was necessary to control the

flow field in the reactor. The temperature profiles along the reactor are shown in Fig. 4d, where

it can be seen that the center-line temperature of the inlet streams is an average of 30 K below

the wall. Similar temperature and velocity profiles were obtained for other experimental conditions

used in this study.

Droplet trajectories are shown in Fig. 4 e. All droplets depicted in Fig. 4e had the same

initial size, 144 /xm, but different radial velocities (0.1-0.5 m/s). The droplets experienced an initial

heat-up and evaporation period where the solvent diffused to the gaseous phase; the end of this

period is marked with asterisks in Fig. 4e. Times for solvent vaporization were calculated to be

« 0.17 and « 0.07 s (for toluene and acetone, respectively). Fig. 5a and b. It should be noted

here, that according to the model used, the vaporization of most of the solvent was completed

before the boiling point of toluene or acetone were reached, (383 K and 330 K) and hence, the

process was diffusion controlled. Upon evaporation the particles foUow the trajectories depicted

in Fig. 4 a and they experience further heat-up till they reach the local gas temperature. Total

residence times were calculated in the range of 7.5 to 14 seconds, depending on each particle's path.

Therefore, in order to produce particles that experience the same residence time in the reactor and

result in uniform properties, the particles should flow in a rather narrow stream around the reactor

centerline. Furthermore, narrow streams minimize the deposition of particles on the upper surfaces

of the funnel, as shown at the bottom of Fig. 4 a. Thence, even if the above arguments indicate

that narrow droplet streams are desired, a certain amount of dispersion is required to avoid droplet

coagulation, so a compromise must be determined experimentally. Moreover, the problem becomes

more complicated since, even small instabilities of the droplet stream can initiate a rapid succession

of collisions and result in a polydisperse aerosol[13].

472
Figure 6: Results of numerical simulation showing (a) gas velocity profiles in the virtual impactor

stage of the furnace and (b) particle tracks through the reactor for two droplet (dry particle) sizes:

144 (42) and 30 (8) fim.

To investigate the effect of size on the residence time of the particles Figs. 5a and b, are included

where calculations of the variation of droplet diameter and temperature with time, respectively, are depicted for particles of diameters ranging between 60 - 144 fim, but same radial velocities. The

total residence times for those particles flowing around the centerline range between 9-7 s. Velocity profiles of the gas as it flows through the virtual impactor are shown in Fig. 6a. This

particular impactor was designed to operate with a particle collection efficiency of 60% for particles

having diameters bigger than 10 ^m[14]. In Fig. 6b the trajectories of two particles of different

sizes that flow through the reactor and pass through the impactor stage can be seen. The bigger

particle, 42 ^m in diameter after evaporation, is trapped in the collection stage while the smaller,

8/Lim, exits the reactor following the gaseous phase. Thus, any fine particles that are formed by

nucleation and condensation reactions in the gas phase, escape the impactor following the exhaust

stream.

SUMMARY

Spherical and monodisperse particles of poly(styrene) and poly(methyl methacrylate) in the size

range of 30 to 60 /xm were produced by an acoustically excited aerosol generator. The generator

formed a monodisperse spray from either mixtures of polymers and their solvents, or monomers

with initiator catalysts and solvents. As the spray passed through a thermal reactor in nitrogen

atmosphere the solvent evaporated and solid, spherical polymer particles were formed. The above

technique can be used for the production of a variety of monodisperse particles of (a) organic or inorganic solids that can dissolve and result in solutions of relatively low viscosity, or (b) polymers

from their monomer precursors whose polymerization can be conducted in bulk.

The thermal treatment of the droplets was studied theoretically using numerical techniques.

Results concerning the velocities and temperatures of both gas and particles, as well as the solvent

evaporation and particle residence times were obtained.

473
ACKNOWLEDGEMENT

This research was supported by the NSF Initiation Grant CTS-8908652. The authors gratefully acknowledge assistance from Marina Belopolski, Jonathan Doughty in the experiments and

construction of equipment.

REFERENCES

1. Sicilia, D.B. "A Most Invented Invention" Article in Invention & Technology, p. 45, Spring/Summer

1990.

2. Adams, L.B., Boucher, E.A., Cooper, R.N. and Everett, D.H., "Third International Confer- ence on Industrial Carbon and Graphite," p. 478. Soc. Chem. Ind., London, 1970.

3. Everett, D.H. and Rojas, F., Chem. Soc, Faraday Trans. 1, 84(5), 1455 (1988).

4. Pendleton, P., Vincent, B. and Hair, M.L., J. Colloid Interface Sci., 80, 512 (1981).

5. Howard, G.J. and Knutton, S., J. Appl. Polym. Sci., 19, 683 (1975).

6. Bhanti, D.P., Dua, S.K., Kotrappa, P. and Pimpal, N.S., J. Aerosol Sci., 9, 261 (1978).

7. Partch, R.E., Nakamura, K., Wolfe, K.J. and Matijevic, E., J. Colloid Interface Sci., 105,

560 (1985).

8. Partch, R.E., Matijevic, E., Hodgson, A. W. and Aiken, B.E., J. Polymer Sci. Polym. Chem.

Ed., 21, 961 (1983).

9. Nakamura, K., Partch, R.E., and Matijevic, E., J. Colloid Interface Sci., 99, 118 (1984).

10. Gangolli, S., Partch, R.E. and Matijevic, J. Colloids and Surfaces, 41, 339 (1989).

11. Levendis, Y.A. and Flagan, R.C., Comb. Sci. and Tech. 53, 117 (1987).

12. Levendis, Y.A. and Flagan, R.C., Carbon, 27, 265 (1989)

13. Berglund, R.N. and Liu, B.Y.H., Environ. Sci. and Tech., 7, 147 (1973).

14. Mulholland, J.A., Yue, G.X. and Sarofim, A.F. "The Formation of Inorganic Particles During

Suspension Heating of Simulated Wastes." Manuscript in Preparation (1990).

15. Biswas, P. "Impactors for Aerosol Measurements: Developments and Sampling Biases" Ph.D.

Thesis, Caltech (1985).

474
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 50

CHARACTERISATION OF SPRAYS OF BITUMEN-IN-WATER EMULSIONS

N. Marcano and A. Williams

Department of Fuel and Energy

Leeds University

Leeds, United Kingdom

ABSTRACT

Sprays and atomisation studies of bitumen- in-water emulsion were carried out

using a Delavan air assisted external mixing atomiser. A Malvern 2600

particle sizer analyser and a Dantec 58N10 phase doppler signal processor

were used for measurements of the spray droplet size distribution at

different air/fuel ratios . Photography were also used to study the spray

break-up mechanism.

Results indicate that bitumen- in-water sprays, under the conditions covered

in this study, will produce smaller droplets than predicted using

correlations developed for air assisted atomisers with low viscosity

liquids. Apparently the addition of bitumen particles to water produces a

different spray break-up mechanism in comparison with water or medium fuel

oil. This results in the formation of smaller droplets. Finally, it was

demonstrated that the use of the Malvern 2600 particle sizer analyser

instrument, using a Model Independent mode, was the most adequate to

characterise the spray droplet size distribution for bitumen- in-water

emulsions

.

INTRODUCTION

The growth in energy consumption has increased the demand for

substitute fuels or ways of extending existing supplies of conventional

fuels. The formation of emulsions of oils with water has been considered as

an alternative way of obtaining liquid fuels from bitumens [1], heavy crude

oils and petroleum heavy fractions from refineries [2]. Recently a new

development in the form of bitixmen- in-water emulsions is being

commercialised as Orimulsion [1,3]. The primary benefit of the dispersion

of bitumens droplets in the water continuous medium is that the viscosity is

reduced which makes easier handling and pumping and in consequence assists

atomisation and combustion generally.

Although two-phase fluids, such as coal-water slurry or bitumen- inwater emulsions, have been burned successfully in industrial combustors,

there is still a lack of detailed knowledge of the atomisation behaviour

A75
particularly in the latter case. However, oil-ln-water emulsions behave

differently to a medium fuel oil and in some respects can parallel coalwater atomisation [4]

.

Efficient atomisation of the fuel is probably the most important

factor in the combustion of liquid fuels. This paper reports a study of the

atomisation of bitumen- in-water fuels using different techniques to provide

accurate results as well as shown the strengths and limitations of each

technique

.

EXPERIMENTAL APPARATUS AND PROCEDURE

The liquid fuel used was a commercial bitumen- in-water emulsion

(Orimulsion) supplied by INTEVEP, S.A. The mass median diameter of the

bitumen particles in the emulsion was 14 jlm and the other characteristics

and properties are given in Refs. 1, 3 and 5.

The atomisation facility consisted mainly of a fuel tank, a gear pump

and a Delavan air assisted external-mixing atomiser model SN 30610-1.

Electric heating was used in the fuel tank and the line to the nozzle in

order to keep the atomisation temperature between 65-68''C. Injection fuel

pressures from 1.2 to 2.8 Bar and atomising air flow rates from 700 to 1400

1/hr, at ambient temperature, were used.

Droplet size measurements as a function of the air/fuel ratio were

undertaken using two different techniques with different instruments: a Malvern 2600 particle sizer analyser using a 300 mm lens and Dantec 58N10

phase doppler signal processor used in a forward scatter configuration and a

scattering angle of 30° . In the measurements performed with a Malvern

instrument the Model Independent (MI) and Rosin-Rammler (RR) modes were used

in order to compare the accuracy of fitting the data and seven hundred

sweeps of the diodes were collected for each experimental measurement [6] . Additional experiments were carried out with water in order to compare them

with the results obtained with the emulsion.

Direct photography using a normal camera and a strobe light and an

Imacon high speed camera at a speed of 1.0 x 10* frames per second were used

to study the spray break up mechanism.

RESULTS AND DISCUSSION

Choice of Droplet Size Distribution Function

Results for Orimulsion atomisation were obtained and compared with

water sprays at the same atomising conditions. Each data point represents

the average of 2 to 3 experiments.

In the case of bitumen- in-water emulsions there is currently no

information about the best choice of drop size distribution function.

Consequently both the RR and MI modes were used to find the best fit of

data. The Log difference (Log diff) term is the parameter used to express

the ability of the Malvern's calculated energy distributions to fit the

experimental data [6] . For RR most of the computed light energy

distributions generated fall in the adequate to good range (Log diff between

476
4.0 to 5.5), but there is a clear tendency to be excellent when the air/fuel

ratio increases. The MI distribution has more strict log diff. requirements

and the results are from good to excellent with log diff no higher than 4.5

in the whole range of experimental work.

Because of the good performance of the MI in fitting this data, this

model was chosen to represent the data and calculate the droplet size

distribution of the spray. The obscuration for all these experiments was in

the range from 0.10 to 0.43, thus the measurements were not affected by

multiple scattering [7].

Measurements of SMD as Function of Flow Parameter

Figure 1 shows the Sauter Mean Diameter (SMD) of the spray versus air

flow rate for different atomisation pressures . Air flow rates lower than

900 1/hr, or air/fuel ratios lower than 0.27, result in a very unstable and

inhomogeneous spray with consequential scattering in the results . At higher

air/fuel ratios, e.g. 0.97, the spray is more homogeneous and the results

are more consistent. A significant reduction in droplet size by increasing

the air/fuel ratio is achieved, but there seems to be a limit to the

effectiveness of increased air flow rates beyond which no further

significant reduction in droplet diameter is reached. These results follow

the normal tendency of all liquids sprayed, Newtonian or not. Figure 1 also

shows the results obtained with the Dantec instrument. These are

significantly different from the ones obtained with Malvern, even increments

in the air flow rate have little effect on the SMD. Water sprays under the

same conditions follow the same behaviour.

The differences between the results, in the first instance, are

because two different types of sampling are employed to determine drop size

distribution. The Malvern measures a line-of -sight average through the

complete spray. It also registers a signal proportional to the population

of drops within the sample volume during the sampling interval, i.e. the

spatial average. The Dantec instrument, on the other hand, samples in a

very small volume, and registers a signal proportional to the number of

drops passing through the volume, i.e. temporal sampling. Generally the

results obtained by spatial and temporal sampling are different because in

many sprays systems the velocities are correlated with size due to

variations in drag with size [8].

Secondly the extremely small sampling volume of the Dantec instrument

(ca. 10"^ cm'^) , in comparison with Malvern could also give a different

measure of the average spray droplet size distribution especially if the

spray is not homogeneous

.

Further problems relate to the nature of the atomisation of two phase

fluids. The droplets produced may consist of droplets of water, or of a

single bitvimen particle surrounded by water, or of a water droplet

containing multiple bitumen particles. Sprays of Orimulsion show this as is

presented in Fig 2. This fact introduces errors in the measurements due to

different reflection or refraction patterns and variations in the particle's

refractive index caused by the different types of droplets. The Dantec

instrument is very sensitive to optical alignment, especially to the

scattering light collection angle, and this should be chosen in such a way

that just one scattering mechanism, ie . reflection or refraction, must

477
dominate for there to be a very smooth relationship between phase and drop

size. This is determined by the polarisation of the incident light beam and

the refractive index of the liquid droplet.

The Malvern's measurement technique is relatively insensitive to the

particle refractive index and shape so that the particles may be absorbing

or non- absorbing, spherical or irregular in shape. Furthermore, the

scattered light pattern is insensitive to the particle velocity and to the

position of the particle in the measurement volume. This represents an

advantage for the measurements of droplet size distribution in sprays of

this kind of fuel.

In general the distribution droplet size of the spray for Orimulsion

presents a wide range of drop sizes and an increase in the air/fuel ratio

reduces the fraction of large droplets in the spray.

Prediction of Sauter Mean Diameter in Orimulsion Sprays

A number of functions are available to describe the SMD of sprays

produced by air assisted atomisers. Lefebvre presents a good survey of

these correlations in his book [9] . The ones related to air assisted

atomisers were used to correlate the experimental results obtained with

Orimulsion.

In general, the majority of the correlations used give high SMD in

comparison with the experimental ones, which is expected considering that

they were developed for low viscosity single phase fluids such as kerosine,

diesel, water, etc. That is the case for the Elkotb [9] equation namely:

SMD = A dg Re"°-2^ We'°-^^ (m^/m^)°-2^ (1)

Reynolds and Weber numbers are expressed as Re = p^^d^/^l^ and We = . p^, and O are the density, viscosity and surface tension of the fuel

respectively, d^ the nozzle orifice diameter, U^^ the air relative to liquid

velocity and ni^,ni^ the mass flow rate for fuel and air. The constant A was

found to have a value of 51 for kerosine sprays.

In an attempt to fit Orimulsion experimental data with this

correlation over the whole range of experimental work, the A value was found

to be a function of the air/fuel ratio according to the following:

A = 7.706 (m^m^) -0.438 for m^/m^ < 1.687

and A = 9.824 for m^m^ > 1.687

Simmon's correlation [9] developed for air-assist atomisers is given

in equation (2) as a function of the atomiser's characteristics. C is a

constant, p^ is the air density and Uj^, the fuel and air velocity. Other

terms are as defined above

.

SMD = C

^^0.06 ^0.375

Pa

0.375 1 (

0.55

(2)

478
The C value was found to be 802 and the experimental data fits very

well with the theoretical ones, but for a very narrow range of air/fuel

ratios . Figure 3 illustrates the fit of the two correlations with the

experimental data.

For air-assisted atomisers the major parameter that affects the

atomisation quality is the relative velocity between the liquid and the

surrounding air (Uj^) . At a constant fuel velocity lower than the air

velocity, increments in the air flow rate represent increments in the

relative velocity; the size of the ligaments decrease, their life becomes

shorter, and on their collapse when exposed to the high velocity air

stream, much smaller droplets are former. This could be the explanation for

the good fit of the Elkotb correlation, which takes into account the air

relative to liquid velocity.

Mechanism of Spray Break-Up

Flat sprays were used to identify the break-up mechanism for bitumenin-water emulsions. High speed photographs shown in Fig 4 indicate that

this occurs mainly as a perforated sheet. The photographs show the

behaviour near the jet and further down the spray. Small holes suddenly

appear in the sheet at a distance from the orifice that varies with the

viscosity of the liquid. They rapidly grow in size until the thickening

rims of adjacent holes coalesce to form ligaments of irregular shape that

finally break up into drops of varying size.

A reduction in the viscosity and surface tension of the emulsion,

produced by a dilution in the water content has an effect on the spray

break-up. Low viscosity makes the spray angle wider and easier for the

appearance of holes and formation of threads , which later break up into

droplets. Higher viscosity makes the sheet motion itself very much more

regular and inhibits the disintegration of the rims of the sheet (Fig. 4A)

.

Also, it resists the deformation of the ligaments into droplets so the

atomisation occurs far away from the nozzle. The sheet seems to be thicker

than the other one, and the spray takes longer to break up completely. This

highlights the importance of spreading the liquid into a very thin sheet to

achieve the finest atomisation [9].

Clark and Dombrowski [10] demonstrated that the drop size is

critically dependent upon the nature of the disintegration process; where

the disintegration occurs by perforations the size of the ligaments is

consequently reduced, resulting in relatively smaller drop size. Thus, the

water addition to the bitumen affects the break-up mechanism in such a way

that the formation of smaller droplets is very much enhanced despite the

high viscosity of the emulsion.

CONCLUSIONS

The atomisation of bitumen- in-water emulsions has some slight

differences compared with single phase fluids such as fuel oil or water in

terms of the break-up mechanism, the influence of the fluid's properties,

the interaction of fuel and air and atomiser design on the spray droplet

size distribution.

479
The main concern in making measurements with this type of fuel is the

selection of the best technique and the proper control of experimental

conditions which result in accurate and meaningful results in the analysis

of the sprays. Accurate determinations of drop size distributions are

important in attempting to relate spray characteristics to other parameters.

The use of the light scattering (Malvern) technique for the characterisation

of Orimulsion sprays demonstrated that the Model Independent distribution

function was the best to fit the experimental data. The Rosin-Rammler model

has been successfully applied to various two phase flows and sprays, but in

this case, the results were acceptable only at high air/fuel ratios.

Orimulsion and in general bitumen- in- water emulsion spray size distributions

are best made using the Model Independent distribution function.

Results obtained with Dantec equipment are highly dependent upon the

equipment measurement arrangement and the position in the spray where the

sampling is done. Although this equipment is ideally suited to detailed

spray modeling because of the velocity and size data obtained, two phase

fluids represent a complication for this technique because consideration has

to be made of the influence of the scattering angle, refractive index and

refraction or reflection pattern of the sample.

ACKNOWLEDGEMENTS

We want to thank INTEVEP, S.A. Venezuela for the samples of Orimulsion

and for financial support (NM) . Also, we are indebted to SERC for the loan

of the DANTEC phase doppler equipment and to Mr. R. Cowling and Mrs M.

Kershaw for their invaluable help.

REFERENCES

1. Lees, B. "Orimulsion review of combustion and handling tests", Energy

World, Vol. 179, pp. 8-10 (1990).

2. Olen, K.R. , Dooher, J. P. and Woodworth, T.H. "Emulsified heavy ends. A

new family of slurry fuels". Fifth International Workshop on Coal-Liquid

Fuel Technology . Technical University of Nova Scotia, October (1985).

3. Hidalgo, J., Jimenez, E., Rodriguez, D. , Alcantara, J., Carrizo, R. , Nava, N., England, G. and Kennedy, B. "ORIMULSION^'^ Demonstration Test

at New Brunswick, Canada: A new technological step towards bitumen

commercialisation" , 14th International Conference on Coal and Slurry

Technology . Florida, USA. April (1989).

4. Thambimuthu, K.V. , Stover, N.S.H. and Whaley, H. "The mechanism of

atomization of coal-water mixtures", IChem E Symposium Series, Vol. 107,

pp. 133-149 (1987).

5. Marcano, N., Pourkashanian, M. and Williams, A. "The combustion of

bitumen- in-water emulsions", FUEL (to be published).

6. Meyer, P. and Chigier, N. "Drop size measurements using a Malvern 2200

particle sizer", Atomisation and Spray Technology, Vol. 2, pp. 261-298

(1986) .

480
7. Felton, P.G., Hamidi, A. A. and Aigal, A.K. "Measurement of drop size

distribution in dense sprays by laser diffraction" , International

Conference on Liquid Atomization and Spray Systems . London , IVA/4

(1985)

.

8. Dodge, L.G. , Rhodes, D.J. and Reitz, R.D. "Drop size measurement

techniques for sprays: Comparison of Malvern laser diffraction and

aerome tries phase/doppler" , Applied Optics, Vol. 26, pp. 2144-2154

(1987).

9. Lefebvre, A. Atomization and Sprays . Hemisphere Publishing Corporation,

New York, NY (1989)

.

10. Clark, C.J. and Dombrowski, N. "An experimental study of the flow of

thin liquid sheets in hot atmospheres", Journal of Fluid Mechanic, Vol.

64, pp. 167-175 (1974).

Fig. 1. Variations of the SMD with the atomising air flow rate for

Orimulsion and water

481
Air <lo» ff » 100 (l/htl Fig. 2. Droplets collected from an Fig. 3. Theoretical SMD using

Orimulsion spray. Elkotb and Simmon's correlations.

A

Fig. 4. High speed photographs for A: Orimulsion, B: Bitumen- in-water

emulsion, 50% water. Flat spray, 3.4 bar.

482
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 5

1

THE INFLUENCE OF SPRAY PARAMETERS ON LOCAL MASS FLUXES

AND DEPOSIT GROWTH RATES DURING SPRAY COMPACTION PROCESS

V. Uhlenwinkel, U. Fritsching and K. Bauckhage

University of Bremen, Verfahrenstechnik

Bremen, Germany

ABSTRACT

The spray compaction process is a relative new method to produce near- net shaped

metallic preforms with better material properties and totaly new materials which are not

producible by conventional methods. But there are many parameters influencing the result.

Generally the aim is to predict local spray conditions in the compaction plane. Therefore a

strategy is developed for the determination of local mass fluxes or deposit growth rates. It is shown, that self similar profiles exist, but they depend on melt properties. The influence of the

atomization pressure is investigated, showing that the centerline value acts very different comparing water and steel atomization.

INTRODUCTION

Metal powder production by gas atomization of the melt is a well-known method. People

are mainly interested in mean particle sizes and size distributions of the overall product. The

equation from Lubanska /I/ is mainly used to predict particle sizes in dependence on atomization parameters, fluid properties and nozzle design. The spray compaction process (OspreyTM

process /2/ /3/, also called liquid dynamic compaction or spray casting) with its aim to produce

a near-net shaped, homogeneous product needs much more information to make a good

specification of the process possible. Therefore the knowledge of local values for mass flux or concentration, size- and velocity distributions are necessary for input data of model calculation

to predict other values of interest as cooling rates of the particles, shape of the deposit, heat

transfer in the deposit, microstructure of the deposit and so on. In this paper we try to develop a strategy to determine the local mass fluxes in the spray

in dependence on some influence parameters in order to find a correlation to calculate the shape

of the deposit. There are generally two methods to produce a desired shape of the deposit:

1) periodical movement of the gas atomizer ring in order to produce for instance flat

products /4/

2) moving the substrate in order to produce tubes or solid discs /5/.

The following list of parameters which may affect the local mass flux points out, how

complex this process can be.

I nozzle design (D, a); spray chamber design; distance from the atomizer (z)

II melt/liquid: properties of the melt {Pi^, a, rj^, ..); mass flow jJl^; initial velocity v^^;

initial temperature Tlq

483
III gas: properties of the gas (Pq, tJq, c^q ...); mass flow Mq, atomization pressure

Pq; velocity Vq

The parameters listed are partly interdependent, i.e. there is a relationship between

nozzle design, gas mass flow and atomization pressure. In this paper we present results showing

the influence of different liquid properties, gas mass flow and atomization pressure respectively.

As far as possible published results of other authors are included in our interpretation.

The first steo of our strategy is to verify whether the radial profiles of the mass fluxes

(or deposit growth rates) averaged by the centerline value (m^c; D,,) and the half-width radius

tq^ results in a self similar profile and if this is valid, to find an equation to determine this

profile. The second step is to describe the effect of various parameters on the centerline mass

flux m^g and the half-width radius Tq^. If the mass flow of the liquid is measured and a self

similar profile is known, the half- width radius can be calculated depending on the centerline

mass flux.

EXPERIMENTAL SET-UP

(1) induction coil

(2) crucible

(3) insulation

(4) heater

(5) tundish

(6) melt delivery tube

(7) primary gas ring

(8) atomization gas ring

(9) deposit

(10) substrate

(11) pressure governor

(12) fan

(13) control valve

(14) N 2" tank

Fig. 1 Experimental set up

484
primary gas(

)

atomizaHon

gas (Pg)

The experimental set-up for the metal

atomization and compaction is shown in Fig. 1. The

metal is heated by an induction coil and the melt is pulled into the pre-heated tundish when the desired

melt temperature is achieved. During the experiment the height of the melt in the crucible is kept

constant to guarantee a constant mass flow IvIl. The

melt leaves the cylindrical nozzle at the bottom of

the crucible and reaches the atomization zone in

free fall. The particles are accelerated and meet the

substrate at a distance z from the atomizer in order

to compact and to build up the deposit. (A method

to obtain particle probes is to collect the particles

in a patternator that consists of water filled tubes).

Fig. 2 shows the design of the nozzle in use

for the metal and water atomization. In general

atomizers are divided into two types: free fall and

confined atomizer. We use a free fall atomizer with

the option to use a primary gas to protect the

atomizer from back splashed melt particles.

geomelrical point

of atomization

Fig. 3 Example for the generation of a dimensionless profile

of the mass flux (b) from the absolute values (a)

SELF SIMILAR PROFILES

In the theory of free jets it is usual to utilize a dimensionless presentation comparing

radial velocity profiles for different distances from the nozzle and parameters of influence. We

use this method to proof, whether self similar profiles exist for the radial mass flux profiles

using different operation and boundary conditions. Fig. 3 shows the way how to make the

profiles dimensionless. Therefore the mass flux Al is divided by the centerline (maximum)

485
value itilj. and the radius is divided by the half- width radius Tq^. While the half- width radius is defined by the radius where the mass flux is just half as high as the centerline value.

Tab. 1 Experimental conditions and results from various authors

aattiod malt/ fluid

gaa

•

g/ca^/a

•

na/s

>^0,5 naz Pg

bar

daposit Pe 20 wt t Mn 5.4 18 400» 8,5

deposit AISI 1026 *,* 5,7 28 400» 8,5

N2 pattamator CuSn6 0,2S 25 400» 6,0

M2 calculation SStFe, lOtCu, ItC 1,03 1,47 115 (d - 50 980 3.0

1,33 1,90 88 (d - 100 1,24 1,77 75 (d - 200 la) deposit AISI 1020 0,36.. 1,09... 52... 800 2,5.. .4 0 "2 1,46 1,87 ...35

pattamator water 4,2... 4,2... 32... 1000 1,5.. .2 5 air . . .2,3 ...2,3 . . .49 pattamator AISI 1020 1,2... 1,5... 52... 40 800 2,5.. .4 0 "2 2,1 2,7

author typa of atoaizar

Mt,

Hathur at.al. /«/

Aiuiavarapu

/>/ Hartaann

/II/

Liu

/»/

Sauran at.al. /12/

this papar

this papar

confined

frea fall

free fall

frea fall

1080

860

860

1000

860

^) from poiol of impingemeal from lip of melt delivery lube

Tab. 1 summarises some results of other authors (including own results) and important

parameters as they are available. The parameters can be very different for example in type of

atomizer, measuring method and fluid properties.

Mathur et.al. /6/ and Annavarapu /7/ use nearly the same experimental conditions but

different melts (Fe 20 wt % Mn and AISI 1026) and this conducts in very different values of the

half- width radius (18 mm, 28 mm). So one can assume that the properties of the melt have an

influence on the spread of the spray.

The results from Liu /8/ are received from numerical calculations, assuming that

monodisperse particles (50, 100 or 200 /im) are injected into the flow field (the process of

atomization itself can not be taken into account). The most important boundary conditions and

assumptions are reported in /9/.

The results from Liu /8/ show - as to be expected - that for the same experimental

conditions, an increase of the particle diameter leads to a decrease of the half-width radius, because of a smaller dispersion of larger particles due to the turbulent fluctuations of the twophase flow. But the absolute values of the half-width radii seem to be too high, because the

value for the 50 /Xm particles is much higher than experimental values of the half- width radius

of the one-phase jet velocity. This may be an effect of the turbulence modelling by the K-e

model, which generally predicts the spread of free turbulent jets too strong /lO/. All the results can only be compared using the profiles of the mass flux or deposit growth rate averaged by the

centerline value in dependence on the radius averaged by the half-width radius. One can use

these findings in order to check whether these results can be described by self similar profiles.

Therefore we assume that the averaged profiles are a function of the type:

G (r*) = exp [-Ki (r*) K2]

• • *

* with G (r*) = mL/m^j. = D (r*)/Dc, r* = r/ro_5 and Kj, K2 = constant.

(1)

Fig. 4 shows the result of this comparison using = In (2) and K2 = 2 for equation (1).

In the range of /r*/< 1 only the result from Hartmann /U/ has a strong deviation from equation (1). This may be caused by the mass flow Ml, which is not constant during the experiment.

In the range /r*/> 1 the calculated values from Liu /8/ are below (d = 50 /Xm) or close to the curve of eq. (1), while the experimental results from Mathur /6/ and Annavarapu /!/ show

a strong deviation from eq. (1) with increasing radius /r*/.

486
Fig, 4 Averaged profiles of mass flux or deposit growth rate comparing eq.

(1) and results of other workers (compare Tab. 1)

Our results in atomizing water and steel (AISI =1020) are similar to the results from

Mathur /6/ and Annavarapu /7/ in their tendency. Fig. 5 shows the averaged results from water

atomization experiments for various atomization pressures, keeping constant the distance from

the atomizer and the mass flow. In comparison to eq. (1) the experimental results for /r*/>l

tend to higher values.

1

0,8 -

0,6

^ 0.4

0.2

fluid : water Ml : 1000 kg/h

z 1000 mm

varl.

equation (1)

-3 -2.5 -2 -1.5 -1 -0.5 0 0.5 1 1,5

dimensionless radius r*

2,5 3

Fig. 5 Averaged profiles of mass flux from water atomization,

using various atomization pressures; compared with eq. (1) (Kj = In (2), K2 = 2)

In Fig, 6 averaged profiles of the mass flux are presented, using various atomization

pressures, compared with equation (1). It is to be seen that the exponent K2 = 1,4 in the range

/r*/>l fits the results by far better than the exponent K2 = 2. Taking K2 = 1,4 in the range

/r*/>l also the calculated mass flow (from eq. (2)) delivers much better results compared with

the measured mass flow.

487
Ml= / 2 7rrmL (r) dr (2)

-4-3-2-10 1 2 3 4

dimensionless radius r*

Fig. 6 Averaged profiles of mass flux (patternator result)

INFLUENCE OF PROCESS PARAMETERS ON CENTERLINE MASS FLUX (OR DEPOSIT

GROWTH RATE) AND ON THE HALF-WIDTH RADIUS

If a self similar profile is found for a given fluid and nozzle system, one has to determine the averaging values (centerline value and half-width radius). We assume that the centerline mass flux itilc can be described by an equation of the type as follows:

= Ai • (Pg*)^ • z ^ • Ml^-* (3)

with Pq* = (pq - Pu)/Pu is the dimensionless atomization pressure. In the paper the influence of

the atomization pressure Pq* is investigated for water and steel atomization, keeping constant

the distance from atomizer z and the mass flow KIl- Therefore the parameters A and A2 of

equation (4) have to be determined.

niL,, = A [z, MJ . (pg*)A2

(4)

Fig. 7a presents the centerline mass flux and the half- width radius in dependence of the

atomization pressure Pq* for water atomization. The increase of the pressure leads to a decrease

of the mass flux and because of a constant mass flow to an increase of the half- width radius.

For the determination of the constants A and A2 Fig. 7b shows log (m^c) in dependence of log

(Pq*) also including the linear regression with the constants A = 2.82 and A2 = -0.55.

In the same way the results of steel atomization, using the same atomizer are evaluated

in Fig. 8a/b, comparing the results from the deposit growth rate and the mass flux. These

results differ because of the different experimental conditions. Three features are responsible

for this difference: 1. bounce of particles from the surface of the deposit, 2. disturbance of the

flow field in the vicinity of substrate and deposit respectively, 3. continuous variation of the

impact location (mainly in z direction) during the growth of the deposit.

488
Fig. 7a/b Centerline mass flux and half-width radius in dependence on the dimensionless

atomization pressure Pq* for water atomization

1 1,2 1,4 1,8 1,8 2 2,2 2,4 2,8 2,8 3 3,2 3,4 0 0,08 0,1 0,18 0.2 0,28 0,3 0.38 0,4 0,46 0,8 dimensionaless pressure pG« log (pG*)

Fig. 8a/b Centerline mass flux in dependence on the dimensionless atomization

pressure Pq* for steel atomization

Looking at the influence of the atomization pressure, both results show an increase of

the mass flux when increasing the atomization pressure and lead to nearly the same exponents

A2 = 0,79 (deposit growth rate) and A2 = 0,86 (mass flux). This behaviour is generally contrary

to the results from the water atomization and cannot be explained by turbulent transport

mechanism of free turbulent jets. We presume that this phenomena is caused by an interdependent influence of the flow in the nozzle near field and fluid properties. A better understanding

of the process in the nozzle near field is necessary to explain this behaviour, as it is partly

reported in /1 2/.

CONCLUSIONS

The spray compaction process is affected by various numbers of parameters. Aim of this

investigations is the prediction of local mass fluxes or deposit growth rates in the case of stationary atomization (static atomizer and static substrate), atomizing different fluids. It is found

that self similar profiles are created, but there is a dependence on fluid properties. Fluid

properties also affect the centerline value. The centerline mass flux is influenced by the

489
atomization pressure and leads to contrary behaviour comparing water and steel atomization.

While the centerline mass flux is decreasing with higher atomization pressure in the case of

water atomization, the centerline mass flux increases in the case of steel atomization.

NOMENCLATURE

Cp specific heat capacity

cf particle diameter

D diameter of the melt delivery tube

6 deposit growth rate

£)* dimensionless deposit growth rate = G gauss function (eq. (i))

m mass flux

m* dimensionless mass flux = "^/^c M mass flow

p pressure

Pq* dimensionless atomization pressure

r radius

rQ^ half- width radius

r* dimensionless radius = Vio^

T temperature

V velocity

z axial coordinate;

distance from the atomizer

a angle between centerline of the

atomizer holes and z-axis

dynamic viscosity

P density

a surface tension

Subscripts

c centerline

G gas

L liquid/melt

0 initial

u ambient

1 primary

REFERENCES

1. Lubanska, H., "Correlation of Spray Ring Data for Gas Atomization of Liquid Metals",

Journal of Metals, No.2, p.45-49 (1970).

2. Brooks, R.G.; Green, S., "Verfahren und Vorrichtung zur Herstellung von gespriihten

Metallkorpern", Patent Nr. P 25 37 103.3-24, 20.8.1975.

3. Brooks, R.G. et al, "The Osprey Process, a novel method for the production of

forgings", Metallurgia & Metal Forming, p. 157-163 (1977).

4. Reichelt, W.; Voss-Spilker, P., "Spray deposition - a new and original shaping process

for the metallurgical industry", Metallurgia, No.9, p.410-413 (1986).

5. Leatham, A.G.; Brooks, R.G.; Yaman, M., "The Osprey process for the production of

spray-deposited roll, disc, tube and billet preforms". Modern Developements in Powder

Metallurgy, Vol.15-17, p.157-173 (1985).

6. Mathur, P.; Apelian, D.; Lawley, A., "Analysis of Spray Deposition Process", Acta

metall., Vol.37, No.2, p.429-443 (1989).

7. Annavarapu, S., "Spray casting of steel strip: Modeling and experimental studies", Ph.D.

Drexel University, 1989.

8. Liu, H., "Berechnungsmodelle fiir die Geschwindigkeiten und die Abkiihlung von Tropfen im Spriihkegel einer Stahl-Zerst^ubungsanlage", Dissertation Uni. Bremen, 1990.

9. Fritsching, U.; Liu, H.; Bauckhage, K., "Numerical Modelling in the Metal Spary Compation Process" ICLASS-91 . Gaithersbury Maryland, (1991).

10. Leschziner, M.A.; Rodi, W., "Calculation of Annular and Twin Parallel Jets using various

Discretization Schemes and Turbulence-Model Variations", Journal of Fluids Engineering, Vol.103, No.6, p.352-360 (1981).

11. Hartmann, G.C., Die Erstarrung von Metallen im SoruhgieBprozeB am Beispiel der

Zinnbronze CuSn6 / Fortschritt-Berichte, Reihe 5: Grund- und Werkstoffe, Nr. 195,

VDI-Verlag Dusseldorf, 1990.

12. Uhlenwinkel, V. et al, "Strdmungsuntersuchungen im Diisennahbereich einer

Zweistoffduse - Modelluntersuchungen fur die Zerstaubung von Metallschmelzen",

Chem.-Ing.-Tech., Vol.62, No.3, p.228-229 (1990).

490
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 52

NUMERICAL MODELLING IN THE METAL SPRAY COMPACTION PROCESS

U. Fritsching, H. Liu and K. Bauckhage

Institut fur Werkstofftechnik

Bremen, Germany

ABSTRACT

Since about 20 years the producing of preformed metal products by means of the spray

compaction process has become a widely known and now quite well established process. But still work has to be done to predict the correct way for the controlling of the process and the influencing of the spray product properties, here the preform. This work is in progress and is at

time investigated by a variety of research groups. In this area the numerical modelling of the

spray process has become an important tool in order of the understanding of the heat and mass

transfer mechanisms in molten metal spraying. Following the subdivision of the spray process

into three different stages, as there are (a) the atomization area, (b) the transport area, and (c)

the compaction area, this paper gives some recent results of our group for the process controlling and optimization obtained by different computer calculations of the continous and two

phase flow field and the related thermal conditions of the dispersed phase. Eulerian and

Lagrangian calculations will be provided and compared as far as possible to measurements,

especially done with the phase- Doppler method. Possible ways for the optimization of the spray

nozzle design and the spray performance in terms of the mean droplet size and the mass flow

ratio gas/melt will be outlined.

INTRODUCTION

The spray compaction technology for molten metals is an industrial process for the basic

shaping and preforming of metals [1]. In the process, from above the melting point heated liquid

metal is poured into the nozzle and is atomized by an impinging inert gas jet of high kinetic

energy. The resulting spray of molten metal droplets will be compacted on the substrate surface

to the desired shape, dependent on its local mass flow rate. The application of the process covers an area of different metals like steel, copper, tin

and aluminium. Also several alloys and composite materials, which are difficult or impossible to produce by conventional metallurgical processes, are in the meantime object of industrial metal

spray processes. The advantages of the spray compaction process compared with conventional

casting techniques are therefore near net shaped products

better homogeneity and less impurities of the product

applicable to new alloys and composites

higher heat transfer coefficient

contribution to micro structural refinement in the pre products.

491
For the scientific analysis of the spray compaction process it is useful to subdivide the

total process in smaller parts, each of which is responsible for an individual aspect of the

process and product optimization. The unit operation "spraying" in an industrial context can be

devided into three parts [2]

the atomization area and the liquid disintegration

the transport process of the multiphase flow in enclosed cavities

the droplet impingement and the formation of the product.

primary gas (p-])

atomizer gas (pq)

geometrical

atomization

point

Fig. 1 Gas atomization of molten metals, principle sketch

The first area of investigations following the above list, to be reported here is the nozzle near field (Fig. 1). Particularly when atomizing sprinkling and/or solidifying fluids, special care has to be taken on to the spray nozzle design or selection and on the operation parameters of

the spray process. Due to back splashing droplets which glue and assemble at the nozzle (called

freezing), the atomization process will be influenced or sometimes even interrupted. In order to understand and control the flow field immediately in the atomization area, the gas flow field

was numerically modelled. Gas flow recirculations in the nozzle near field will be examined and

possible ways for a constructive improvement of the nozzle will be shown.

The second area of investigation is the two phase flow field of the gas jet and the dispersed metal droplets in the spray chamber. In order to enhance the understanding of the flow

field and the temperature distribution of the gas as well as of individual particles in the spray

chamber and to establish a control model for the regulation of the process, extensive experimental and theoretical studies are necessary. The purpose of this part of the paper is to present

model calculations based on an Eulerian approach for the momentum transport and exchange for

both phases under simultaneous consideration of the phase change kinetics of the dispersed

492
phase. This part of the study investigates the influence of the directly controllable input process

parameters on the stationary turbulent flow and temperature fields in the spray chamber.

Following the area division given above the third area of investigation is the compaction

area and the formation of the preform. This part will be presented as a separate paper in this

conference, based primary on experimental investigations in the metal atomization and in model

experiments with water as the dispersed fluid.

The three different areas of research are combined versus their initial and boundary

conditions.

Method of Solution

The numerical simulation of the gas flow field in the near nozzle region is based on the

approximation of the primary- and of the atomization nozzle as. coaxial ring gaps. Under this

assumption the system of partial differential equations describing the gas flow field are solved

using the standard features of the PHOENICS code, version 1.3 [3]. This multipurpose computational fluid dynamics code is based on a finite volume discretization of the governing equations

using the SIMPLEST formulation for the solution. [4]. It was assumed that the flow field can be

described as to be two dimensional, using polar co-ordinates and single phased. The behaviour

of the gas is described by the isentropic gas law and the turbulence modelling is based on the

standard K-€ model, together with log-low wall functions on all rigid walls.

The theoretical model for the two phase flow field in the spray chamber consists of the

equations of continuity, momentum and energy conservation for both phases as well as of the

phase change kinetic equation. The motion and temperature equations are derived by Reynolds

decomposition and time averaging of the corresponding instantaneous conservation and energy

equations, considering the weight and buoyancy of droplets as well as the friction between the

two phases to be the volume forces in the equations of motion, and the phase change heat and

the convective heat loss of the droplet-phase to be a heat source of the gas-phase and to be a heat sink of the droplet phase in the temperature equations. Closure of the time-mean equations

was achieved by modelling the turbulent correlations in the equations of mean motion according

to Elghobashi's method [5] and in the equations of mean temperature according to [6, 7]. The

corrected K-€ equations are similarly formulated. In deviation to the standard set of constants in

the K-€ model, the c^ coefficient was set to 0.07, according to [7]. For a first approach the turbulent diffusivity v^^ of the dispersed phase is related to the turbulent diffusivity of the gas

phase i/,g, according to Schonung [8] as t'tp/^tg ~ ^-^ because the ratio from i/,p to i/,g is between 0.3 and 0.6 according to [8] for tne conditions of this calculation. Similarly the extra

production and dissipation of the turbulent kinetic energy, which results from the correlations

of the fluctuations of gas velocity, are also treated simplified as a linear function of the droplet

concentration in the corresponding terms with the same approach as described by Danon [9].

A kinetic model for the phase change has to solve two problems: to what extent the

droplets can be undercooled and how quickly the phase change in the droplets occurs. On their

way downwards to the substrate the smallest droplets are subject of a rapid quenching process,

thus reaching the impact point as completely solidified particles. On the other hand the largest

droplets don't lose so much of their heat, thus reaching the substrate as entire liquid droplets.

As the microstructure of the deposit and also its porosity depend upon the temperature, the

fraction of solidification and physical structure of the impacting droplets, it is essential to develop a phase change kinetic model to investigate the influence of phase change on the temperature and cooling rate of droplets. In [10] a phase change kinetic model has been derived on

the base of the classical nucleation theory applying the Johnson-Mehl-Avrami transformation

for solid state phase change in metals and alloys. By the combination of this model with the

above mentioned temperature equation of the droplet phase, the nucleation temperature, i.e. the

maximal achievable undercooling can be estimated and the temperature dependence of the phase

change rate can be simulated.

The method of solution for the coupled system of partial differential equations is a modified version of the MAC-SOLA method with a second order upwind discretization [11].

493
The gas flow in the nozzle near field

The principal sketch of a twin-fluid atomizer for molten metal atomization is shown in

Fig. 1. In order to move the atomization area away from the melt outlet, the point where the

atomizer gas stream joins the melt stream has been brought downstream. In addition to that in

our special case we used a coaxial primary gas stream (treated as a coaxial free jet) in order to convey the undisturbed melt stream to the atomizing point, to raise the drag of back splashing

droplets and to transport these droplets downwards back to the atomization zone. The atomizing

gas meets the melt stream under an angle a, generating the atomization point in some distance

downwards from the nozzle. But as the photomontage from a model water nozzle shows, the

atomization of water in such a nozzle nevertheless produces back splashing droplets which cannot be avoided under wrong operating conditions. One can see the movement of individual

droplets upwards the atomization area. These droplets move in the inner circle vertically upwards against the main gas and melt flow direction and in the outer regions downwards.

Fig. 2 Gas flow in the near nozzle region, Pq (comp) = 1.89 bar; Pq (exp) = 2 bar

Fig. 2 illustrates on the left side the result of the computation for the nozzle sketched in

Fig. 1. This type of nozzle contains the typical central cylindrical flow regime of the melt

stream and an atomizing coaxial jet at a lower part of the flow system. This type of melt

atomization has been described in detail by [12] and [13]. The results plotted represent an

atomizer pressure level of 1.89 bar absolute, which equals the velocity of the incoming atomizer

gas flow to the velocity of sound, and without any primary gas application. The atomization gas

flow which enter the domain under an angle of 10° contracts in the same manner as a coaxial

jet, and hits the core region at a position in some distance upwards from the geometrical

atomizing point. One can see a big area of a closed recirculating vortex, the border of which is marked by the solid line. The upward pointing velocity component in this vortex has a maximum value of nearly 45 "*/^. Individual particles which enters this area will be accelerated upwards and some of them will hit the nozzle main body near by the fluid exit.

A9A
The comparison of this computational results with a picture of the water atomization in

Fig. 2 (right side) indicates a good agreement with the behaviour of single particles beyond the

atomization point, but the size of the standing vortex is underestimated in the simulation. One

has to notice, that the pressure level for the atomizer gas in the experiment is a little bit different from that used in the simulation.

(0 JQ

<0

3

%

20

154

10

5-

-3 i

•10

13

-20 -i

/

atomizer gas pressure

»- experiment pq = 1.0 bar

M- computation pQ = 0.89 bar

primary gas pressure p-j = 0.0 bar

-3 —I— -1 —

T"

10 1

1

nozzle distance Z /cm/

Fig. 3 Pressure on centerline - comparison computation/experiment

In Fig. 3 the comparison for the computed pressure on the core surface with pressure

measurements is drawn. (Z = 0 stands for the lower edge of the atomizer gas ring) The pressure

maximum, that is the main hit point of the atomizer gas stream on the core, is computed closer

to the gas exit than the experimental results indicate. In combination with this failure, the angle

of attack is overpredicted and therefor the radial velocity component is greater in the simulation

than in the experiment, resulting in a higher value for the pressure maximum and the underpressure. This deviation is mainly affected by the use of the standard li-€ model, which is known to overpredict the jet growth in the axial direction. Therefor the inner recirculation area of a coaxial jet will be underpredicted in size. The result of a constructive alteration is shown

in Fig. 4 for the case of 1.89 bar pressure for the atomizing gas and 1.4 bar pressure for the

primary gas. The figure indicates no vortices directly at the primary gas jet exit and the inner

area of the atomizer ring. A flow field without any closed recirculating vortex occurs.

The flow field in the spray chamber

The overall flow field of the continuous and the dispersed phase are shown in Fig, 5. Because of the great velocity gradients in the flow field, the vectors are scaled square-rooted.

The calculation was done for a mean droplet diameter of 100 jlm and a mass flow ratio gas/melt

of 0.8. The gas flow field consist of two parts. At first the gas jet which enters the spray chamber at the top. Due to the existence of the substrate the growth rate of the jet varies non linear

from top to bottom of the chamber. Secondary a huge amount of gas can be observed which

recirculates in the spray chamber. Therefor the velocity vectors of the dispersed phase at the

edge of the spray cone are directed radially inwards in the mid part of the spray cone, half between the inlet and the substrate, because of the vertical gas entrainment from the recirculating

gas into the jet. For this case no particle recirculation occurs into the spray chamber, all metal

droplets will hit the substrate.

495
In Fig. 6 the gas and droplet velocities on the centerline of the spray cone are illustrated

for three different calculations with mean droplet diameters of 50, 100 and 200 jlm. The direction of the momentum transfer is reversed at some distance between 300 and 600 mm from the

spray chamber inlet. While in the upper part of the chamber the droplets are accelerated by the

gas jet, thereby reaching a maximum kinetic energy dependent on the drop size, further

downstream the droplets are faster than the gas flow and the direction of the momentum transfer is from the disperse to the continuous phase. In this figure PDM-measured droplet velocities

at z = 450 mm are shown which fit well with the calculated data.

The radial profiles of the droplet velocities, shown in Fig. 7 as a function of the mass

flow ratio with values of 0.55, 0.80 and 1.08 give a non-Gaussian profile as to be expected. The

width of the spray cone increases with the augmentation of the mass flow ratio. It is not yet

clear, why this result does not fit the experimental observation of a narrower substrate material

distribution with increasing gas pressure. But it is obvious, that the assumption of a constant

mean diameter of the metal droplets with varying mass flow ratio is quite unrealistic, because

the mean diameter decreases with increasing atomizing gas pressures.

CONCLUSIONS

By means of the spray compaction process for molten metals, near net shaped industrial

pre-products can be manufactured with advanced properties and based on new alloys and composites. Process and product optimization can be based on the numerical modelling of the

process, which must be validated and interact with model and on-line experimental works.

By dividing the process into three stages in this paper the background and results for the

first two stages, the atomization area and the spay chamber, have shown that at first by the understanding of the gas flow field in the near nozzle area the geometry of the atomizer can be

improved in order to prevent recirculation of droplets in this area.

496
The simulation of the two phase flow in the spray chamber yields a basic understanding

of the transport process in the chamber and the exchange behaviour between both phases. The

important parameters are the mean droplet sizes and the mass flow ratio of gas and melt.

r [mm] = m(ptm)

M,/M, = 0,80

100 200 300 400 ao —

1

I

\S N N -

WWNWWNN^^-^""^ s ^

^ \\\\\\\\\\ \ ^

^

8

i 1

- i

- s

§

100

r [mm] <^ = ""'""•^

Af,/M, = 0,8

DO 900 400 900

l| X V

\ V

WW

]\\\\\w

\\\\\\\\\\\

s

h i

i

h i

s

i T

§

I- i

§

Fig. 5 Velocity vectors for continuous and dispersed phase

r = 0,Mj/M = 0.3

1 — <L, = 50(/xm)

2 — cL, = m{nm)

Z — dL, = 2G0(;im)

4,5,6 — Gas

• measured data

z (m)

Fig. 6 Gas and droplet velocities on centerline

497
z = 450(Tnm)

rfp = 100(/zm)

1 — Mj/M = 0.55

2 — iWj/M, = 0.80

3 — Mj/M, = 1.08

Fig. 7 Radial profiles of the gas and droplet velocities at z = 450 mm

REFERENCES

1. Apelian, D., Gillen, G. and Latham, A.: Near Net Shaped Manufacturing Via the Osprey

Process, ASM Meeting, Orlando, Florida, Oct. 6-9 (1986), Proc.

2. Bauckhage K.: Das Zerstauben als Grundverfahren, Chem.-Ing.-Tech. 62 (1990) 8; pp.613-625.

3. Rosten H.I., Spalding D.B.: The PHOENICS Reference Manual; CHAM Technical Report

TR/200; London 10 (1987).

4. Patankar S.V., Spalding D.B.: A calculation procedure for heat, mass and momentum

transfer in three-dimensional flows; Int. J. Heat Mass Transf. 15 (1972); pp. 1787 ff.

5. Elgobashi S.E., Abou-Arab T.W.: A two-equation turbulence model for two phase flows,

Phys. Fluids 26 (1983) 4, pp. 931-938.

6. Oliver A.J.: The prediction of turbulent flow and heat transfer over backward facing

steps, Computer Methods in Fluids, ed. by Morgan et al., Ondon Plymouth (1980)

7. Bradshaw P.: Turbulence, Springer- Verlag (1976).

8. Schonung, B.: Comparison of different dispersion models for particles in Lagrangian and

Eulerian prediction codes, Proc. of the 2. Int. Conf. on Fluids Mechanics . China (1987).

9. Danon H. et al.: Numercial calculations of two-phases turbulent round jet. Int. J. Multiphase Flow 3 (1977), pp. 223-234.

10. Liu H.: Berechnungsmodelle fiir die Geschwindigkeiten und die Abkiihlung von Tropfen

im Spruhkegel einer Stahl-Zerstaubungsanlage, Ph.D. thesis, Univ. Bremen (1990).

11. Hirt C.W. et al.: SOLA- A Numerical Solution Algorithm for Transient Fluid Flows,

LA-5852 (1975).

12. Mehrotra S.P.: Mathematical Modelling of Gas Atomization Process for Metal Powder

Production; Powder Metallurgy Int.; Vol 13 No. 2 (1981) 80-84.

13. Lubanska H.: Correlation of Spray Ring Data for Gas Atomization of Liquid Metals.

Journal of Metals; Feb. (1970) 45-49.

498
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 53

SIMULATION OF A SPRAY COMBUSTION BEHAVIOR FOR TWO KINDS

OF SLURRY FUELS (Effect of Two Stage Air Introduction and

Spray Characteristics on Combustion)

H. Aoki, T. Funihata, S. Tanno, T. Miura and M. Daikoku*

Dept. of Biochemistry and Engineering, Tohoku University, Sendai, Japan

*Dept. of Mechanical Engineering, Hachinohe Technical College, Hachinohe, Japan

ABSTRACT

The effect of two stage combust ion (TSC) air introduction and spray characteristics on combustion behavior was studied for two kinds of slurry

fuels (pitch-water, PWM, and coal-water mixture, CWM) spray combustion system.

Experiments were carried out on the isothermal flow in an acrylic duct and on

the combustion flow in two different furnaces. Velocity components and fluctuations for the isothermal flow field were measured with LDA. The gas temperature, oxygen and NO concentrations in the furnace and NO and soot concentrations in the exhausted gas were measured for combustion experiments.

Numerical analyses for the isothermal and combustion flow field were performed. As a result of the analysis, it was difficult to simulate the TSC

isothermal air flow pattern by two-dimensional axisymmetrical analysis. For

the combustion field, the NO concentration increased with the residence time

of the combustion gas in the high temperature region and the soot concentration decreased with an increase of the droplet residence time in the furnace

and of the oxygen concentration. For spray characteristics, the combustion

efficiency was enhanced by decreasing the initial droplet velocity even if the

droplet mean diameter was larger.

INTRODUCTION

A development of a high efficient combustion technology is urgently

necessary to prevent the global warming and the acid rain which are caused by

emission of CO2 and of NOx and SOx, respectively. The use of low rank fuels

such as pulverized coal, coke or pitch and the operation in a more severe

combustion condition are required from the view point of the energy cost and

the energy resources limit. The improvement of the design and operating

conditions on combustion apparatus is thus needed.

One solution for these problems is to examine the combustion characteristics with the simulation method. The simulation on the liquid spray combustion system cannot be however studied enough. In this study we examine the

TSC air introduction simulation method which is modeled as the axisymmetrical

flow. We then study the spray combustion behavior of pitch-water mixture(PWM)

and coal-water mixture(CWM) in two different furnaces and carry out the simulation of spray combustion on an axisymmetrical coordinate system using the

k- e turbulence model modified with the swirling flow.

499
THEORY

1. Govern in;; equations for gas phase

All time-mean transport equations in this study can be written for Cartesian geometry as;

dx (pU^) +

_1_ r a r (rp V* ) a X

a

a r

(

(rr *

a 4> ax

a » a r

)

) + S* + Sd* (1)

where (}> represents the dependent variables which denote the mass (continuity)

,

momentum(U, V, W), turbulence energy(k), eddy dissipation rate(£), enthalpy(h), radiative heat flux(Fj, j=l,2) and mass fractionCmi, i=soot, volatile, N2. O2. CO2. CO, NO and HCN). F 4> is the exchange coefficient, S is

the source term in the gas phase and Sd <l> is the source term between particle

and gas phases whose coupling is estimated using the PSl cell method. We use

the turbulence model modified to the swirling flow field proposed by Launder

+ 1 1

et al

The flow characteristics at the wall boundary are estimated using the

standard wall function modeP' . The wall temperature is given from the measurement value. The inlet conditions such as velocity components, temperatures, concentrations of gases are also given by the experimental condition.

2. Analysis of slurry droplet dynamics and reaction

' Atomizer

V2Dn

2-1. Analysis of droplet dynamics

Fig. 1 shows the model of the

spray jet near the burner. The

model of droplet dispersion is

assumed that Dthe position of

ejection is at the exit of the

burner, and 2)droplets are already spread out in the radial

direction at the tip of a core

region x=xo(xo=6Dn, Dn ; nozzle

diameter), beyond which there is

no break-up or coalescence of

droplets. For convenience computing the trajectories in Fig. 1,

let the locationCxo, r j ) be the initial point

fraction Pj ejected from the initial point in

pressed as;

Pj = 27rSjrj Ar/M,

core region

Xo=6xDnX

Dispersion

Fig. 1 Spray dispersion model.

of the computation. The number

the total droplet numbers is ex-

(2)

where Sj is the dispersion in the annular region between radius (rj-Ar/2) and

(rj + Ar/2). Sj is then given by;

Sj = S.axexp[-0.65(rj/b)Ml+0. 027 (rj/b)M] (3)

which is based on the correlation for the distribution of the concentration in

an ax isymmetr ical co-flow jet^'. Sma^i is a normalizing constant to be determined so that Sj satisfies Eqn. (4).

500
Ml = 2jSjrjdrj (4)

where b (= {12tan ( a ) +1} Dn /2) is the jet radius and Mi is the liquid mass flow

rate and aj is the injection angle. If we define N as the total injected

droplet number per unit time, the number rate na of droplet having the diameter Dpi ejected from the core region tip is?

nio = PiPjN (5)

where Pi is the droplet number fraction with Dpi, whose size is given by the

logarithmic normal distribution.

- 2-2. Droplet reaction model The

four-flux model**' is employed to

the radiative heat transfer. The

gas absorption coefficient

containing the soot is calculated

using the mixed gray gas assumption by Hottel^'. The formation

rate of NO is calculated from

both rates of thermal and fuel NO

formation models which are assumed to be produced by the extended Zeldovich

mechanism and the model proposed by Scott et al^' , respectively. The soot

formation and its burnout in gas phase are given by the model proposed by

Abbas et al" . The slurry droplet reaction model is shown in Fig. 2. The

droplet energy equation is solved including the convective and radiative heat

transfers, combustion heat release of particle and the latent heat of evaporation. The convective heat transfer rate is modified by the transfer number.

The devolat i 1 izat ion is computed by the single overall reaction model.

The rate coefficients are given by Arrhenius expressions (k = Aexp (-E/T) ) whose

constants are quoted from another experiment for coal*' and previously measured by a single droplet combustion experiment for pitch. Those values and

experimental conditions are shown in Table 1 with the proximate and ultimate

analyses of fuels.

Char burnout rate is limited by both rates of chemical reaction kc^* and

oxygen diffusion kd^°'. All carbons in the char are converted to CO with a heterogeneous reaction and then CO converts to CO2 with a homogeneous reaction

whose rate is controlled by the eddy dissipation model^^'.

3. Numerical solution

The Lagrangian droplet equations are simultaneously

Runge-Kutta-Fehlberg (RKF) method for each droplet which is

classes in the initial ejection angle and 15 classes in the

diameter. Partial differential equations for gas phase are

the method of Patankar^^'.

EXPERIMENTAL APPARATUS AND CONDITION

1. Isothermal flow experiment

The acrylic duct is a 500mm I.D. x 2,000mm length geometry. Velocity

components and fluctuations are measured by the two-color, three-beam type

laser Doppler anemometer (DANTEC Co.) having two channel counter processors.

The swirl burner is a co-axial flow type and both of primary and secondary

airs flow into the duct with a tangential velocity component produced by the

vanes at the burner tip. The vane angles of both air ports are fixed at 10

yHp /<VolatUes yrCO

, f ^ Devolatilization' Char

Heated tvap.of ^\\\^ combustion

"P /mfekwater ^^swelling

373K Coal

or Pitch O2

9}

Ash

and Soot

Fig. 2 Slurry droplet reaction model.

integrated with

divided into 7

initial droplet

calculated with

501
Table 1 Operating conditions for PWM and CWM combustion.

Species Pitch Coal

Proximate analysis [wt%]

Volatile 38.9 28.37

Ash 0.22 18.03

Ultimate analysis [wt%]

C 86. 22 68. 51

H 5. 69 0. 42

N 1.19 1.26

S 5.69 0.42

0 - 7.42

Water content [wt%] 30.0 32.0

Combustion parameters

Devolatilization: A Cs"'] 3.68x10^ 5.50x10'

E CK"'] 3.30x10' 9.47x10'

Char combustion: A [g/cm^s] 8710 8710

E Cr'] 17980 17980

Swelling factor 3.0 1.5

Operating conditions

Total fuel flow rate [kg/hr] 150 80

Primary air flow rate [Nm'/hr] 102 20

Secondary air flow rate CNm'/hr] 1140 362

Primary air temperature [K] 283 283

Secondary air temperature [K] 423 523

Air ratio [-] 1.32 1.2

and 20 degrees whose swirl numbers

are 0.12 and 0.32, respectively.

The volume flow rate is fixed at 40

and 260Nm'/hr for the primary and

secondary air, respectively. TSC

air flow is introduced into the duct

at the axial direction x/D=0. 8 to

let air flow tangentially whose rate

ratio to the secondary air is 30%.

To estimate the accuracy of calculation in an axisymmetr ical coordinate

system, two different TSC air introduction methods are examined in this

study, one for the practical introduction method as shown in Fig. 3- (a)

and another for method as modified

to be an axisymmetr ical coordinate

system in Fig. 3-(b).

I Fig. 3 Experimental apparatus for

isothermal flow measurement

on two stage combustion air

introduct ion.

2. PWM combustion experiment

The furnace is a 1, 400mm I.D. x 5,000mm length geometry with a single

burner. TSC air is introduced tangentially from the tube on the furnace side

wall. The temperature and the oxygen and NO concentrations in the furnace and

the NO and soot (residual carbon and soot in gas phase) concentration in the

exhausted gas are measured with variations of TSC air flow rate ratio.

502
3. CWM combustion experiment

The furnace is a GOOmm x 600mm rectangular cross section x 5,000mm

length geometry with a single burner. To simplify the calculation, the furnace is modified to be a cylindrical furnace though it has a rectangular cross

section. The sootd.e. residual carbon and soot in gas phase) concentration

at the furnace center line is measured using the isokinetic sampling method

with variations of axial distance.

RESULT AND DISCUSSION

1. Application of TSC air introduction model

As the TSC air is usually introduced into the furnace from one or more

inlet ports, it is necessary to calculate as a three-dimensional cylindrical

coordinate system. However, it takes more CPU time and memory in the calculation than the two dimensional analysis. We then examine the applicability of

the two-dimensional analysis to the usual TSC system.

Fig. 4 shows the axial and tangential velocity profiles in the duct where

the solid point, solid line and dotted line correspond to results of the

measurement (method (b) shown in Fig. 3) and the predictions with and without

TSC air, respectively. Both measured and calculated results with TSC air are

in good agreement at two different axial distances.

Fig. 5 shows experimental results at two different axial distances. Measured velocity components, solid circle marks (a), tend to have non-ax isymmetric velocity profiles compared with the solid line (b). It is thus necessary

to apply the three-dimensional analysis to the usual TSC process flow into the

furnace from one tube.

0.00

0.50 -

0.00

0.50

I

c^0.00 ^ 0.50

0.00

• Measurement

With TSC air

Without TSC air

==« "" —, , m ~

\ \^ -

1 1 1 1

2.00

1 .50

0.50

0.00

O.SO

0.00

0.50

I

1 .00 Q ^ 0.00 TSC ^

0.50

0.00

• Measurement

With TSC air

Without TSC air

/As •

1

1 1 1 1

2.00

1 .50

1.00 Q

TSC^

0.50

0.00 0.10 0.20 0.20 0.40 0.50

r/D [-]

0.00

0.00 0.10 0.20 0.30 0.40 0.50

r/D [-]

Fig. 4 Measured and predicted velocity profiles in the duct

for the introduction method (b).

503
X / D = 0.5 D = 0.5 [m]

Uin=38.8 Cm/s]

. • (a)

,1 — (b) .

^V. •

•

• . .

- -

• *

1 * * y . •

•

-0.3-0.5

Fig. 5

cO.2 5

^0.1

-0.1

0.2

J. 0.1

c5

0

r/ D

'- -0.1

-02

X / D= 1.5 D=0.5 Cm]

Uin=38.8[nn/s]

• (a) — (b) .

• / ~ • » / • y

• • •

r • • •

\ • * / "\ •

05 -0.5 [-]0

r/D

0.5

Measured velocity profiles in

ducL for the two different

introduction methods.

[-]

the

5 200

o

to

Fig. 6

20 AO 60

TSC air ratio C7o]

Effect of TSC air flow rate

ratio on NO and soot concentration in the exhausted

gas.

2. Effect of two stage air flow rate ratio on combustion

Fig. 6 shows the effect of TSC air flow rate ratio to the total combustion

air flow rate on NO and soot concentrations in the exhausted gas. Predicted

results are in good agreement with experimental results in both NO and soot

concentrations. The predicted NO concentration takes the maximum at 30% in

TSC air ratio and the predicted soot concentration increases with TSC air

ratio.

To confirm numerically this tendency, the gas temperature and oxygen

concentration at the center line of the furnace, the droplet weight and NO

concentration which is integrated over each cross section of the furnace are

shown in Fig, 7. In the case of a small TSC air ratio, NO is produced downstream in the furnace as the residence time of combustion gas is short. On

the contrary, in the case of a large TSC air ratio, the gas temperature decreases with an increase of the TSC air inflow rate and then the NO concentration in the exhausted gas decreases, though NO concentration increases in the

high temperature region near the burner.

Soot production rate in exhausted gas nearly equals to the droplet weight

since most of the soot (about 90%) are derived from residual carbon. In the

case of small TSC air ratio, the oxygen near the center line of the furnace

does not steeply decrease and the high temperature region is formed as the

residence time of combustion gas and droplets is short, and the combustion gas

is mixing well. With an increase of TSC air ratio, there is little oxygen

content near the center line of the furnace, though the combustion delay time

is short.

504
2000.00

1500.00 u

3

o 1000.00 v

a.

500 . 00

co

c

«>o

coo

co

c

0)o

co

Xo

25.00

20.00

15.00

10.00

100.00

o.o

r-. 1000.00

750.00

500.00

250.00

0.00

0.00 r 1.00 2.00 3.00 4.00

TSC oi r x/D [-]

Fig. 7 Predicted gas temperature, oxygen

concentration, droplet weight and

NO concentrat ion (gas temperature

and oxygen concentration are shown

at the center line, droplet weight

and NO concentration are integrated

over each radial plane.

1.0 2D 3.0 40

Axial distance x Cm]

Fig. 8 Comparison of predicted

measured unburned ratio

the exhausted gas with

variations of Ma/Mi.

5.0

and

in

3. Effect of spray characteristics on combustion

Fig. 8 shows the predicted

and experimental unburned ratio

at the furnace center line with

variations of Ma/Mi. Both predicted and experimental results

show that the unburned ratio is

the largest in the case of the

maximum Ma /Mi ( in it ial droplet

velocity is the highest) though

the SMD is the smallest.

Fig. 9 shows the predicted

temperature distribution in the

furnace with variations of Ma/Mi.

In the case of small Ma/Mi (initial droplet velocity is the

lowest), the increase of gas

temperature delays, however the

droplet weight decreases and the

high temperature region expands

down-stream because of long

residence time of droplets though

high temperature region

Fig. 10 shows the

furnace with variations

air is recirculated by

reaction of droplets and

droplets flow densely near the

center line of the furnace. In

the case of large Ma /Mi, the gas

temperature decreases toward the

down-stream due to short residence time of droplets though the

is formed near the burner.

predicted oxygen concentration distribution in the

of Ma/Ml. In the case of small Ma/Mi, the combustion

the large recirculation flow zone which promotes the

inflammable gas with oxygen. As Ma/Mi increases, the

zone of low oxygen concentration expands down-stream so that the droplet and

505
inflammable gas continue reacting

even in the outlet of the furnace.

The combustion efficiency can be

thus enhanced by decreasing the

initial droplet velocity even if the

droplet is larger.

CONCLUSIONS

A theoretical model for spray

combustion of PWM and CWM on an

axisymmetr ical coordinate system has

been described and effects of twostage air flow rate ratio and spray

characteristics on combustion behavior are studied. Summary is;

(1) lt is necessary to apply the

three dimensional analysis to the

usual TSC process flow into the

furnace from one tube.

(2) N0 formation rate is mainly

influenced by the combustion gas

residence time in the high temperature region and the soot is mainly

influenced by the gas temperature

and oxygen concentration.

(3) The combustion efficiency is

enhanced by decreasing the initial

droplet velocity even if the droplet

is larger.

O.M

o

c

0.50

o.so

a

oc

0.00

0.00 2.00 i.OO 6.00 X/0 [-J

Fig. 9 Predicted gas temperature

distributions in the furnace

with variations of Ma/Mi.

a

c

0.50

o

0.50

REFERENCES

Priddin and

Eng. , 99, pp.

B. I, 0.00

Fig. 10

x/0 t-1 Predicted oxygen concentration

distributions in the furnace

with variations of Ma/Mi.

Pub. Co. (1976)

on Comb. , pp. 661-671 (1973).

1) Launder, B. E. , C. H

Sharma: J. Fluids

231-239(1977).

2) Launder, B. E. and D. B. Spalding^Comp

Methods in Appl. Mech. and Eng.,

3, pp. 269-289(1974).

3) Rajaratnam, N.

: "Turbulent Jets", Elsevier Sci

4) Gosman, A. D. and F. C. Lockwood : 14th Symp. (Int. ) 5) Hottel, H. C. : "Sect ion in Heat Transmission", McGraw-Hill (1954)

.

6) Scott, C. H. , L. D. Smoot and P. J. Sm ith: 20th Symp. (Int. ) on Comb. , pp. 1391-

1400(1984).

7) Abbas, A. S. and F. C. Lockwood: J. Inst it. of Energy, 58. PP. 112-115(1985)

.

8) Goldberg, P. M. and Essenh igh, R. H. : 17th Symp. (Int. ) on Comb. , pp. 145-154 ( 1978)

.

9) Field. M. A. :Comb. Flame, 13. pp. 237-252(1969).

10) Mulcahy, M. F. R. and I. W. Smith: Rev. Pure and Appl. Chem. . 19. pp. 81-108 (1969)

.

11) Magnussen. B. F. and B. H. H jertager : 16th Symp. (Int. ) on Comb. , pp. 719-729 (1976)

.

12) Patankar, S. V,

: "Numer ical Heat transfer and Fluid Flow", Hemisphere Pub. Co.

(1980).

506
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 54

CRYOGENIC LIQUID-JET BREAKUP IN TWO-FLUID ATOMIZERS

R.D. Ingebo

National Aeronautics and Space Administration

Lewis Research Center

Cleveland, OH, U.S.A.

ABSTRACT

A two-fluid atomizer was used to study the breakup of liquid-nitrogen jets

in nitrogen, argon and helium atomizing-gas flows. A scattered-light scanner

particle sizing instrument previously developed at NASA Lewis Research Center

was further developed and used to determine characteristic drop diameters for

the cryogenic sprays. In the breakup regime of aerodynamic-stripping, i.e.,

sonic-velocity conditions, the following correlation of the reciprocal Sauter

mean diameter, , with the atomizing-gas flowrate, Wg, was obtained: D~|

= k where k is a proportionality constant evaluated for each atomizc ^ c ing gas. Values of k^ = 120, 220 and 1100 were obtained for argon, notrogen

and helium gasflows respectively. The reciprocal Sauter mean diameter D~l

and gas flowrate, W^, have the units of cm~l and g/sec, respectively.

In the regime of capillary-wave breakup, or subsonic conditions, it was

found that; = kgWO-75, where k = 270, 390 and 880 for argon, nitrogen and

helium gasflows , respectively.

INTRODUCTION

The drop size of disintegrating liquid jets in low velocity airflow is

controlled primarily by a capillary-wave mechanism and characteristic particle

size is generally relatively large. However when a liquid jet is injected into

high velocity or sonic flows, an aerodynamic stripping mechanism occurs in

which the jet is disintegrated into a large number of small liquid particles,

thereby producing a large liquid surface area. In this case, disintegration

occurs before waves have time to form on the liquid- jet surface. The emphasis

of the present study was on the mechanism of aerodynamic-stripping, in which

two-fluid nozzles were used to impact liquid-nitrogen jets with argon, nitrogen

and helium atomizing-gas flows.

Cryogenic liquid-jets with atomizing-gas flows injected into rocket combustors are quickly disintegrated into small-droplet sprays. In order to calculate vaporization or burning rates, it is necessary to characterize the spray

in terms of drop size distribution and mean drop size such as the Sauter mean

diameter. Once characteristic drop sizes are known, mathematical expressions

can be derived to adequately describe a two-fluid atomization process in which

507
various liquid propel lants and atomizing-gas combinations are used to enhance

spray combustion and yield high combustor performance over a wide range of

operating conditions.

Knowledge of the characteristic size of drops in sprays are also needed in

experimental studies of fuel-spray combustion in diesel and other internal combustion engines and gas turbine combustors. In addition, knowledge of water

sprays is needed to study the formation of ice on airfoils in icing wind tunnel

experiments. In the task of characterizing sprays, detailed knowledge of the

mechanics of liquid-jet disintegration is especially needed at the point of

initial spray formation close to the atomizer orifice. This is especially true

in the study of highly volatile cryogenic liquids. From this knowledge, accurate initial conditions can be established for modeling a fuel spray combustion

process or modelling the process of ice formation on airfoils.

Numerous investigators have obtained experimental drop size data and correlated it with relative velocity, i.e., gas velocity relative to liquid velocity and also with liquid properties, as given in Refs. 1 to 7. Some of the

correlations do not agree very well with atomization theory; which is generally

attributed to the fact that measurement techniques and drop sizing instruments

have yet to be developed and standardized to the extent that good agreement

might be expected. The fluid property having the greatest effect on the drop

size of sprays produced with two-fluid nozzles is the gas velocity. Although

numerous investigators have studied this phenomena, the effect of gas mass

flux on the drop size of cryogenic sprays has not been established in the

spray literature. Therefore, the main objective of the present study was to

determine the effect of atomizing-gas mass flux on liquid-nitrogen spray

characteristics.

Prior to the present study, an investigation of water sprays was made with

two-fluid nozzles and good agreement of experimental results with atomization

theory was obtained, as discussed in Ref . 1. It was found that the Sauter mean

diameter, D32 , could be correlated with nitrogen gas flowrate, W^, raised to

the -1.33 power, which agrees well with theoretical expressions for liquid jet

breakup in high velocity gasflow. As a continuation of this study, the present

investigation was initiated to extend experimental conditions to include cryogenic liquid-jet breakup in helium, argon, and nitrogen gas flows.

Two-phase flow in which transfer of momentum from low velocity and sonic

gasflows to the surface of liquid-nitrogen jets was experimentally investigated

by using three different atomizing gases to produce clouds of liquid-nitrogen

droplets. Tests were conducted in the cap i 1 lary-wave and aerodynamic-stripping

regimes of disintegrating liquid jets. A scattered-light scanning instrument

developed at NASA Lewis Research Center was used to measure characteristic drop

diameters of the cryogenic- liquid sprays. By correcting for gas turbulence and

thermal gradient effects on measurements, reproducible data was obtained with

the scattered-light scanner. To avoid the loss of small liquid-nitrogen droplets due to their high vaporization rate, measurements were taken close to the

nozzle orifice. The entire spray cross section was sampled with a 4.4 by

1.9 cm rectangular laser beam, at an axial distance of 1.3 cm downstream of the

atomizer

.

APPARATUS AND PROCEDURE

A two-fluid nozzle was mounted in the test section as shown in Fig. 1,

which also shows the optical path of the scattered-light scanner. Air supplied

at ambient temperature, 293 K, passed through the 15.24 cm inside diameter test

section and exhausted to the atmosphere. The test section was 1 m in length

508
AIR

. 2« CM .

Fig. 1 Atmospheric pressure test section and optical path

of scattered-light scanner.

NITROGEN If^^^

0.561 CHLIOUID NITROGEN DIAIETER £

/ 0.663 CH--'

Fig. 2 Diagram of pneumatic two-fluid atomizer.

and a 5.08 cm diameter orifice was used to measure the air flowrate in the test

section. A flow of dry air at a velocity of 5 m/sec was maintained in the test

section to aid in transporting small droplets through the laser beam. This

prevented ambient humid air from contacting the spray and thereby avoided the

formation of ice particles in the laser beam.

A detailed diagram of the two-fluid nozzle is shown in Fig. 2. It was

mounted in the center of the test section and operated over pressure ranges of

0.2 to 1.0 MPa for both liquid nitrogen and the atomizing gases. Liquid nitrogen sprays were injected downstream into the airflow just upstream of the duct

exit. The spray was sampled at a distance of 1.3 cm downstream of the atomizer

orifice with a 4.4 by 1.9 cm rectangular laser beam produced by the 0.01 cm

diam aperture shown in Fig. 1.

Liquid nitrogen, LN2 , at a temperature of 77 K measured with an I.C. thermocouple, was axially injected into the airstream at a flowrate of 27.5 g/sec,

as indicated by a turbine flow meter. The atomizing gas was then turned on to

break up the LN2 jet and weight flow rate was measured with a 0.51 cm diameter

sharp-edge orifice. After the air, atomizing-gas and LN2 flow rates were set,

characteristic drop diameters were determined from measurements made with the

scattered-light scanner.

509
The optical system of the scattered-light scanner is shown in Fig. 1.

The instrument measures scattered-light as a function of scattering angle by

repeatedly sweeping a variable- length slit in the focal plane of the collecting lens. The data obtained is scattered-light energy as a function of the

scattering angle relative to the laser-beam axis. This method of particle size

measurement is similar to that given in Ref. 8 and it is described in detail in

Ref . 9. Calibration was accomplished with five sets of monosized polystyrene

spheres having diameters of 8, 12, 25, 50 and 100 \xm. Since the sprays were

sampled very close to the atomizer orifice, they contained a relatively high

number of very small drops. As a result, the light-scattering measurements

required correction for multiple scattering as described in Refs. 9 and 10 for

the case of high concentrations of droplets. Drop size measurements were also

corrected as described in Ref. 9 to include Mie scattering theory when very

small characteristic droplet diameters, i.e., <10 \m, were measured.

Atomizing-gas temperatures were approximately 293 K and surface temperatures of the liquid-nitrogen jets were always near the boiling point of liquidnitrogen, i.e., approximately 77 K. This created large temperature gradients

that deflected the laser beam and caused beam steering to occur. However, by

carefully adjusting the laser beam away from the slit in the scattered-light

scanner optical system, the effect of beam steering on measurements was

negl igible

.

EXPERIMENTAL RESULTS

As shownin Fig. 1, the entire spray cross section was sampled at an axial

distance of x = 1.3 cm downstream of the two-fluid atomizer orifice. Values

of the Sauter mean diameter, . and reciprocal SMD, , are plotted against

atomizing-gas flowrate as shown in Fig. 3. Tests were conducted over a relatively wide range of atomizing-gas flowrates that included both subsonic and

sonic gas-velocity conditions.

At relatively low or subsonic atomizing-gas flow conditions, breakup of

the LN2 jets occurred primarily in the capillary-wave regime of atoraization

ATOHI ZING-GAS O HELIUM O NITROGEN A ARGON

ATOMIZING-GAS FLOHRATE, g/sec

Fig. 3 Correlation of Sauter mean diameter with atomizing-gas flowrate.

510
Table 1 Proportionality constant, k, and gas flowrate

exponent, x, for correlating expression:

Breakup regime Argon Ni trogen Hel ium

Capi 1 lary-wave

:

kg, constant

X, exponent

Aerodynamic-stripping:

k^, constant

X, exponent

270

0.75

120

1.33

390

0.75

220

1.33

880

0.75

1100

1.33

cind the following general correlation of D~l with gas flowrate, W , was

obtained:

p-l _ , ^0.75

32 - Yg

where the reciprocal SMD and gas flowrate are expressed as cm~^ and g/sec,

respectively. The correlating constant kg = 270, 390 and 880 when argon,

nitrogen and helium, respectively, were used to atomize the LN2 jets. The

exponent 0.75 for Wg is approximately 10 percent higher than the 2/3 exponent given for gas velocity in the theoretical expression reported in Ref. 11

In regime of aerodynamic-stripping or sonic-velocity atomizing-gas flow

conditions, the following expression was obtained:

"32 - c^g

where the correlating constant kc = 110, 220 and 1100 when argon, nitrogen and

helium, respectively were used as atomizing gases. The exponent 1.33 for Wg

is the same at the 4/3 exponent given for gas velocity in Ref. 11. Thus, good

agreement with atomization theory was obtained in the regime of aerodynamic

stripping conditions.

Transition from the regime of capillary-wave breakup to that of

aerodynamic-stripping occurred at values of D~| = 700 cm~l or = 15 lim,

approximately, for the atomizing gases used in the present study. Values of

kg and kg are given in Table 1.

CONCLUDING REMARKS

The fluid mechanics of liquid-nitrogen jet breakup is much more difficult

to study than the atomization of fuel or water jets. This is primarily due to

the fact that surface temperatures of the liquid-nitrogen jets used in the

present study were always near the boiling point of liquid nitrogen, i.e.,

approximately 77 K. Since the atomizing gases were at room temperature,

approximately 293 K, this created large temperature gradients that deflected

the laser beam and caused beam steering to occur as characteristic drop diameter was measured with the scattered-light scanner. This problem was minimized

in the present investigation by moving the laser beam away from the slit in the

scattered-light scanner optical system.

511
REFERENCES

1. Ingebo, R.D., "Experimental and Theoretical Effects of Nitrogen Gas Flowrate on Liquid-Jet Atomization, " Journal of Propulsion and Power , Vol. 4,

No. 6, Nov. -Dec, 1988, pp. 406-411.

2. Kim, K.Y. and Marshall, W.R., Jr., "Drop Size Distributions from Pneumatic

Atomizers," AIChE Journal , Vol. 17, No. 3, May 1971, pp. 575-584.

3. Lorenzetto, G.E. and Lefebvre, A.H., "Measurements of Drop Size on a Plain

Jet Airblast Atomizer," AIAA Journal , Vol. 15, No. 7, July 1977,

pp. 1006-1010.

4. Nukiyama, S. and Tanasawa, Y., "Experiments on the Atomization of Liquids

by Means of a Air Stream, Parts III-IV," Transactions of the Society of

Mechanical Engineers, Japan , Vol. 5, No. 18, Feb. 1939, pp. 63-75.

5. Weiss, M.A. and Worsham, C.H., "Atomization in High Velocity Airstreams,"

American Rocket Society Journal , Vol. 29, No. 4, Apr. 1959, pp. 252-259.

6. Wolf, H.E. and Andersen, W.H., "Aerodynamic Break-up of Liquid Drops,"

Proceedings of the 5th International Shock Tube Symposium , Z.I. Slawasky,

J.F. Moultin, Jr., and W.S. Filler, eds., Naval Ordnance Lab., White Oak,

MD, 1965, pp. 1145-1169. (Avail. NTIS, AD-638011.)

7. Ingebo, R.D., "Fuel Spray Simulation With Two-Fluid Nozzles," AIAA Paper

89-0053, July 1989 (also, NASA TM-101367, 1988).

8. Swithenbank, J., Beer, J.M., Taylor, D.S., Abbot, D., and McCreath, G.C.,

"A Laser Diagnostic Technique for the Measurement of Droplet and Particle

Size Distribution," in Experimental Diagnostics in Gas Phase Combustion

Systems , " B.T. Zinn and G.T. Boulman, eds.. Progress in Astronautics and

Aeronautics, Vol. 53, AIAA, New York, 1977, pp. 421-447.

9. Buchele, D.R., "Particle Sizing by Weighted Measurements of Scattered

Light," NASA TM-100968.

10. Felton, P.G., Hainidi, A. A., and Aigail, A.K., "Measurement of Drop Size

Distribution in Dense Sprays by Laser Diffraction," in ICLASS-85

;

Proceedings of the Third International Conference on Liquid Atomisation

and Spray Systems , P. Eisenklam and A. Yule, eds.. Institute of Energy,

London, 1985, Vol. 2, pp. IVA/4/1-IVA/4/11

.

11. Adelberg, M., "Mean Drop Size Resulting from the Injection of a Liquid Jet

Into a High-Speed Gas Stream," AIAA Journal , Vol. 6, June 1968,

p. 1143-1147.

512
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 55

EFFECT OF AIR SWIRLERS ON ATOMIZATION

AND SPRAY FLAME STABILITY

C.-P. Mao

Delavan Inc. West Des Moines, lA, U.S.A.

ABSTRACT

This study investigated the effect of outer air swirlers on atomization and stability performance of spray

flames. In airblast atomizers, the outer swirler controls the air effective area and spray angle. Thirteen different

air swirlers were used to examine the importance of air effective area and spray angle. It was found that a wider

spray angle produced smaller mean drop sizes and better stability. For a given spray angle, an increase in air

effective area resulted in further improvement in spray quality and stability limits. However, spray angle had

a more significant effect on stability performance than the air effective area.

INTRODUCTION

The advantages of swirl flow in gas turbine engine systems have long been recognized. The proper use

of swirling flow in combustors provides better fuel/air mixing, improves combustion efficiency, stabilizes the spray

flame, and reduces pollutant emissions.

The flow characteristics produced by swirler systems in combustors have been studied by many

researchers. Gupta et al [1] summarized various aspects and applications of swirl aerodynamics. However, there

is insufficient information about the effect of air swirlers used in the fuel injectors on atomization and spray

flames.

In advanced gas turbine engines, the fuel injectors are required to operate over a wide range of

conditions. Their performance is especially critical for aircraft engines which must function in harsh and

abnormal environments. Air swirler design can play a very important role in determining the performance of

fuel injectors. Several investigators [2-4] have reported the effect of airblast nozzle swirlers on atomization. Mao

[2] concluded that they have a significant effect on mean drop size and spray angle. Generally, a wider spray

angle is associated with finer droplets. In a later study [5], it was found that the air swirler design in atomizers can also have a significant impact on spray flame stability.

Turbine combustors often encounter flame-out problems during engine acceleration because of the rich

mixture produced by the fuel injectors. Therefore, one of the primary requirements in aircraft engines is to ensure that spray flames can be sustained over a wide range of fuel and air flows. The air swirler helps control

the structure of the spray flame. It was noted that an optimized swirler not only provides good flame stability,

but also improves performance during ignition and altitude relight conditions.

The present study examines the effect of air swirler design on atomization and spray flame stability. The

structure of the sprays produced by different swirlers will be analyzed to determine its influence on the flame.

TEST METHODS

A conventional method [6] of evaluating spray flame stability is to gradually increase or reduce the fuel

flowrate at a fixed air flow condition to establish the flame extinction limits. This process is repeated at increasing levels of air flow, until the complete stability loop can be determined.

Our study selected a slightly different approach. The air flow was gradually increased or decreased at a fixed fuel flow condition to determine the lean and rich limits. The process was then repeated at several

different fuel flowrates. This approach was chosen simply because our experimental setup allowed us to change

the air flow more easily than the fuel flow.

513

J
A prefilming airblast atomizer was selected for our study. The fuel sheet produced by the atomizer was

sandwiched between a non-rotating central air jet and a swirling outer air stream. The outer swirler and the air

gap were the primary design variables of interest. There are a number of parameters which designers can

manipulate to provide different swirl flow characteristics. For a typical air swirler with heUcal vanes, these

parameters include vane angle, slot depth and cutter width. By varying these design parameters, wide ranges

of spray angle and air effective area could be attained.

Since the air effective area and the spray angle are two of the primary dependent variables in the swirler

design for a given nozzle envelope, it was only necessary to vary the air gap and the slot dimensions in the outer

swirler. Table I summarizes the swirler parameters.

During the study, photos were taken to measure the spray angles. An Aerometrics Phase/Doppler

Particle Analyzer was used to make detailed measurements of drop size and velocity. Combustion tests were

also conducted to determine the stability limits of the spray flames.

RESULTS AND DISCUSSION

Effect of Outer Svtirlers on Air EfTective Area and Spray Anple.

Air flowrate measurements were made for all swirlers to find their effective areas. The air pressure drop

across the nozzle was varied within the range between 1.2 and 55 kPa. Spray angle measurements were obtained

at fuel pressures of 69, 138 and 276 kPa.

From Table I, we can perceive systematic variations in the air effective area, CdA, and the spray cone

angle when the air gap and the slot dimensions are changed. For the same swirler slot width and depth, the

effective area and spray angle increased with an increase in the air gap. It also appears that, for a fixed air gap, a smaller slot dimension would generally produce a smaller CdA and a wider spray angle.

Both the air effective area and spray angle play a very important role in determining the spray structure

in the combustors. From an aerodynamic standpoint, they not only affects fuel atomization, but also control the

characteristics of the spray flames.

Effect of CdA and Spray Angle on Spray Structure

An Aerometrics Phase/Doppler Particle Analyzer was used to make detailed measurements in sprays.

This unit was recalibrated using a single-drop generator to obtain new separation distances between the three

detectors in the instrument. The original distances specified by Aerometrics were found inappropriate for drops

larger than 200 microns.

The primary parameters used for examining the structural differences in sprays included distributions

of drop Sauter mean diameter (SMD), axial velocity, and volume flux. All measurements were made at a 5-cm

distance downstream from the nozzle face. The nozzle was traversed in 0.5-cm increments from one boimdary

to the other along the major diameter. The radial distributions of the parameters were then plotted.

For a given swirler, the spray structure gradually changed as the flow operating conditions were varied.

Using Swirler No. 11, the effect of increasing the fuel flowrate on SMD and axial velocity at a fixed air pressure

drop of 6.9 kPa is shown in Figs. 1-2. In airblast atomizers, it is typical that the SMD increases and the spray angle decreases with an increase

in the fuel pressure when the air flow is maintained at a constant value. The increase in fuel pressure reduces

the relative velocity between the liquid and air phases. Hence, atomization becomes poorer and drop size

increases. Meanwhile, the higher fuel flow also increases the mean axial velocity of the drops, resulting in a narrower spray angle.

Shown in Fig. 2, there was a significant change in the velocity profile as the fuel pressure increased from

69 to 276 kPa. At 276 kPa, the droplets located in the center core of the spray had the maximum axial velocity.

However, at 69 and 138 kPa, the velocity profiles were characterized by a double peak. Two apparent maximum

velocities appeared at the center area of the sprays. It is very common in a swirling flow that the velocity profile

undergoes a transition from a double-peak to a single-peak mode as the fuel flow increases. This is primarily

associated with the change of spray angle and relative magnitude between axial and swirling velocities. This type

of structural change has a serious impact on spray flames. The next section will discuss what type of spray

structure is most beneficial in engine combustors.

514
120

100 -

Sauter Mean Dia. (um)

-8 -6 -4 -2 0 2 4

Radial Distance (cm)

69 kPa 138 kPa 276 kPa

Fig. 1 Effect of increasing fuei flowrate on

tiie radial distribution of Sauter mean

diameters for air swirler No. 11. Air

pressure drop was at 6.9 l<Pa.

Fu«l Prvtture

-B-%9 kPa

3°CP O

OGGOC^

•• -« •« -I • t 4 • « RMIal Olstaiie* (an)

Fig. 2 Effect of increasing fuel flowrate on

drop axial velocity distribution for

air swirler No. 11. Air pressure drop

was at 6.9 kPa.

To observe the effect of spray angle on spray structure, it is necessary to mmntain a constant air effective area (CdA). Swirlers 4, 7 and 13 were selected for a comparative study. These three swirlers have about the

same CdA (17.4 mm^, but different spray angles.

Fig. 3 shows a comparison of the radial SMD and drop axial velocity distributions for the three swirlers

at a fuel pressure of 276 kPa and an air pressure of 6.9 kPa. At the same CdA and radial distance, the SMD

decreased with an increase in spray angle. Swirler No. 13 had the widest spray angle (124°), and the smallest

mean drop size at most of the sampling locations.

The shape of the velocity profile changed from a double-peak to a single-peak mode as the spray angle

decreased. With the same CdA, it was surprising to note that the maximiun drop velocity could be tripled by

reducing the spray angle from 124° to 81°. This clearly demonstrates that the design of the outer air swirler can

control the spray structure.

It was also of interest to examine the effect of the air effective area, CdA, on spray structure when the

spray angle remained constant. Swirlers 5 and 12 were selected for comparison. These two swirlers produced

about the same spray angle at a fuel pressure of 276 kPa and air pressure of 6.9 kPa. Swirlers 5 and 12 have

effective areas of 18.7 and 14.8 square mm, respectively.

Fig. 4 shows the radial distributions of SMD and axial velocity for the two swirlers. The increase in the

effective area had a very minimal effect on the mean drop size and only increased the velocity in the center area. The shape of the velocity profile remained about the same, primarily because the spray angle did not change.

In some other cases, (e.g., Swirlers 3 and 11), similar results were observed.

For a given air swirler design, Fig. 5 demonstrates the effect of the air gap on spray drop size and

velocity. Swirlers No. 1, 2 and 3 were chosen to examine this effect. For the same slot dimension, a larger air

gap increased both CdA and the spray angle. As indicated in Fig. 5, a larger gap produced a wider spray angle

and a smaller mean drop size.

The drop velocity was affected differently by the air effective area and spray angle. Although an increase

in CdA generally increased the drop velocity, an increase in the spray angle reduced the velocity. In Fig. 5, the

spray angle seems to have a dominant effect on the velocity profile. A wider spray angle produced a double-peak

velocity profile, whereas a narrow spray angle produced a single-peak velocity profile.

515
Fig. 3 Effect of spray angle on radial SMD and drop mean axial velocity distributions as the spray angle increases. All swirlers have the same CdA.

Fig. 4 Effect of air effective area CdA on the radial SMD and drop axial velocity

distributions. Spray angle maintained the same value.

516
As the air gap increased, the shape of the velocity profile shifted from a single-peak to a double-peak

mode, and the maximimi velocity decreased. It will be demonstrated later that a double-peak velocity profile

is desirable in combustors. This type of spray structure can sustain the flame over a wider range of air flows.

After comparing all of the swirler designs, an important result was derived from the spray structure

measurements. That is, the structure appears to be the same if the air effective area and spray angle remain

constant. What remains to be examined is whether quality of the combustion will also be similar? It will also

be of great interest to determine whether the spray angle or the CdA plays the greater role in affecting the

stability spray flames.

Effect of Spray Angle and Air Effective Area on Sprav Flame Stability

One of the primary goals of the present study was to optimize the air swirler designs for better stability

performance. From the preceding analysis, spray angle and air effective area appeared to be the significant

variables. For a given CdA and spray angle, however, the spray structure seemed to be independent of the air

swirler dimensions. Therefore, the flame stability study will be focused on CdA and spray angle rather than

dimensional differences.

Flame stability was evaluated from the lean and rich limits observed for several different fuel flowrates.

As in the spray structiu'e analysis, spray angle and CdA were varied.

The rich and lean limits were determined for each swirler using the test method described earlier. At

a flxed fuel pressure, the rich limit was determined by reducing air flow to the extent when flame extinguishes.

Because the air flowrates were uncertain near the lean limits, the extinction criterion was that the spray flame

has to stay alight for at least one minute. The air flow was further increased until it met the above criterion to And the actual lean limit. After establishing the limits, the results were plotted as fuel/air ratio vs. air flowrate.

At a flxed fuel flow, an increase of air flow will typically reduce the spray angle and push the flame front

away from the nozzle face. Because the flame is conflned within the spray boundary, blow-out ultimately occurs when the flow velocity exceeds the flame speed.

Fig. 6 shows typical flame stability ranges at several fuel flow conditions. As the fuel pressure increased,

there was a greater separation between the rich and lean limits. This indicates that fuel nozzles may operate

over a wider range of air flows without flame-out problems at higher flowrates. For all swirlers, the lean limits

vary signiflcantly, depending upon the spray angle and CdA. However, the rich limits do not vary signiflcantly.

SMD (urn)

-4 -2 0 2 4

Radial Distance (cm)

swirler 3

Spray AngI* 94*

Air Oap O.aa mm

wirtor 2

89'

0.B1 mm

awirlar 1 74* ass mm

Axial Velocity (m/s)

4 -2 0 2 4

Radial Distance (cm)

swlrlsr 3

Spray Angia 94*

Air Qap O.OS ma

wirlsr 2

89' Ml am

swirlar 1 74* ass ma

Fig. 5 Effect of Increasing air gap on radial SMD and drop axial velocity distributions.

Fuel pressure was 276 kPa and air pressure was 6.9 kPa.

517
FUEL/AIR RATIO FUEL/AIR RATIO

0 5 10 IS 20

AIR FLOWRATE (kg/hr)

O 69 kPa ^ 138 kPa ° 276 kPa Fuel

Fig. 6 Spray flame stability limits for air swirler

No. 8 at a fixed fuel flowrate of 69, 138

and 276 kPa.

Rich Limit Fuel Flowrate 9.8 kg/hr

\

1

Lean Limit

I 1 1 0 5 10 15 20 25

AIR FLOWRATE (kg/hr)

O swirler 4 ( 81*) ^ swirler 7 (111*) ° swirler 13 (124*)

Fig. 7 Comparison of flame stability limits at increasing spray angle under the same air

effective area.

It would be desirable to determine those spray characteristics which will produce a stable flame over a wide range of air flow conditions.

The range of air flowrates required for good flame stability will be determined for several different spray

angles. Fig. 7 shows a comparison of stability limits for several swirlers at a fixed CdA. At the same fuel flow

rate, the swirler that produced the widest spray angle showed the widest range of stability limits. That is, the

spray flame sustained without extinction over the widest range of air flow conditions. Swirler 13 had the widest

spray angle and the best performance in flame stability.

The above results can be easily explained using the spray structure shown in Fig. 3. Because of its wide

spray angle, Swirler No. 13 produced the smallest mean drop size. In addition, its velocity profile showed a

favorable double peak where the low velocity drops reside in the center core of the spray. This type of velocity

profile favors a central recirculation zone for anchoring the spray flames.

On the other hand, Swirler No. 4 had a very narrow spray angle. Its velocity profile is chcu^acterized by

a single-peak where the maximum velocity is located near the center of the spray. This maximum velocity can

easily exceed the speed of the flame and cause its extinction.

The importance of the air effective area on the performance of combustors has long been recognized

by engine manufacturers. Because variations in CdA are usually associated with chzmges in spray angle, it is

difficult to identify which parameter is more important. Hence, the effect of CdA was examined using air

swirlers that produced the same spray angle. In the following case, Swirlers No. 5 and 12 were compared in Fig. 8. The data indicate that the air

flowrate range required for a stable flame is slightly wider for the swirler with higher CdA. This result seems

to contradict what is assumed by many engine designers; namely, that a lower CdA will produce better

performance from the standpoint of flame stability limits. In fact, for a fixed spray angle, a decrease of CdA

further reduced the range of stability limits. Hence, the observed performance may actually be due to the wider

spray cone angle. As indicated in Table I, for a given air gap, a reduction of CdA resulted in an increase in

spray angle.

According to Fig. 4, the primary characteristics of the spray structure for swirlers 5 and 12 are similar

because the spray angle was about the same. The effect of a larger CdA was a higher average drop velocity and

a slightly reduced mean drop size. The improvement in stability limits is probably caused by the reduction in

drop size rather than the increased drop velocity. In addition, the local fuel concentrations have increased so

as to favor the sustention of flames.

518
FUEL/AIR RATIO

2

0

•wirier 0 cdA is.r mm^ o •wirlar 12 CdA t4.a mm^

- \ 270 kP« Fuel

H

\%

\* Q \t

\ \*

- \ \ \ \ \ * \ \

»

V \* V V

V V

Y V V X V ^

1S8 kPmV X.^^

1 1

-a

1 1 0 S 10 16 20 26

AIR FLOWRATE (kg/hr)

Fig. 8 Comparison of flame stability limits by

increasing CdA at several fuel flow

conditions under the same spray angle.

The above results indicate that the spray angle has a significant effect on the stability limits of the flame.

An increase in spray angle and CdA would result in improved stability. However, these two parameters have

a contrary effect on drop velocity. That is, an increase in spray angle decreases drop velocity but an increase

in CdA increases velocity.

In terms of spray structure, a wide spray angle provides a favorable velocity profile characterized by a double-peak mode. A higher CdA produces smaller drops but an unfavorable drop velocity profile. In addition

to drop size and velocity, it is also important to take into account the volume flux distribution. A higher volume

flux generally extends the range of flame stability limits.

SUMMARY

The following major conclusions can be drawn from this study:

1. The air swirler design affects the air effective area (CdA) and the spray angle. These two parameters

determine the structure and stability of the spray flames.

2. For any given CdA and spray angle, the air swirler design has little effect on spray structure and flame

stability.

3. A wide spray angle appears to allow the spray flame to stay alight over a wider range of air flow

conditions.

4. The air effective area has a less significant effect on spray flame stability limits than that of the spray

angle. However, for a given spray angle, an increase in CdA will result in a further improvement in

stability limits.

5. Although smaller drops and low velocity produce a more stable flame, the local drop concentration also

plays important role in determining the stability limits. A higher drop concentration means a more

stable flame.

519
REFERENCES

1. Gupta, A. K., Lilly, D. G. and Syred, N., Swirl Flows, Abacus Press, 1984.

2. Mao, Chien-Pei, "Effect of Air and Fuel Swirlers on Atomization", Delavan Inc, Technical Report 562,

August 1990.

3. Martin, C. A., "Air Flow Performance of Air Swirlers for Gas Turbine Fuel Nozzles", ASME Paper No.

88-GT-108, 1998.

4. Martin, C. A., "Leading Edge Vortex Generation and Its Application to Twin Fluid Atomization

Nozzles", Ph.D. Thesis, University of Detroit, 1982.

5. Mao, Chien-Pei, "Effect of Air Swirlers on Atomization and Spray Flame Stability", Delavan Inc,

Technical Report 563, September 1990.

6. Lefebvre, A. H., Gas Turbine Combustion, Hemisphere Publishing Corporation, 1983.

TABLE 1

A SUMMARY OF AIR SWIRLER PARAMETERS

Swirler Air Gap

(mm)

Slot Width

(mm)

Slot Depth

(mm)

GdA

(mm^

Spray

Angle*

(Degree)

1 .38 2.39 1.27 17.4 74

2 .51 2.39 1.27 20.0 89

3 .66 2.39 1.27 20.6 94

4 .36 2.39 1.09 17.4 81

5 .48 2.39 1.09 18.7 96

6 .38 2.39 0.86 16.1 106

7 .46 2.39 0.86 17.4 111

8 .64 2.39 0.86 18.1 128

9 .51 1.98 1.27 18.7 103

10 .66 1.98 1.27 20.6 116

11 .51 1.98 1.09 18.1 97

12 .36 1.57 1.09 14.8 102

13 .53 1.57 1.09 17.4 124

All swirlers have twelve vanes with a 45* vane angle

* Spray angle was measured at air aP of 6.9 kPa and fuel pressure of 276 kPa

520
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 56

EFFECT OF DODECANOL CONTENT ON THE COMBUSTION OF

METHANOL SPRAY FLAMES

C. Presser*, A.K. Gupta^, C.T. Avedisian* and H.G. Semerjian*

*Chem. Sci. and Tech. Lab, National Institute of Standards and Technology, Gaithersburg, MD, U.S.A.

^Dept. of Mechanical Engineering, University of Maryland, College Park, MD, U.S.A.

*Sibley School of Mechanical and Aerospace Engineering, Cornell University, Ithaca, NY, U.S.A.

ABSTRACT

The structure of a swirl-stabilized spray flame, fueled by methanol and two methanol/dodecanol mixtures,

has been examined. The main objective of this study was to obtain evidence of the occurrence of microexplosions

in spray flames. Measurement of droplet size and velocity distributions were carried out under burning conditions

using a single-component phase/Doppler interferometry system. Initially, experiments were carried out with

methanol and 50/50 methanol/dodecanol mixture spray flames. Some evidence was found to indicate the occurrence of microexplosions in the mixture flame. In this study, it was observed that the occurrence of microexplosions is reduced for 75/25 methanol/dodecanol mixture flames. These results indicate that the tendency for microexplosions

to occur appears to strengthen with higher dodecanol content.

1. INTRODUCTION

An investigation of the combustion of methanol/dodecanol mixtures offers the possibility of examining the

phenomenon of microexplosions. Microexplosions or disruptive burning is the process whereby a droplet literally

explodes during its combustion into smaller droplets due to internal bubble formation [1,2]. Microexplosions of

particular mixtures of methanol and dodecanol are thought to occur because of large difference in volatility of the

two fuels and preferential vaporization of the two components of the miscible mixture. An equivolume mixture

of methanol and 1-dodecanol has been used in studies of burning single droplets. Wang and Law [1] have shown

that, for monodispersed droplet streams, microexplosions occur in a 50/50 mixture (by volume) of methanol and

1-dodecanol. Recent evidence by Yang et al. [2] has indicated that microexplosions for this mixture were not

observed for stationary isolated and unsupported droplets under spherically symmetric burning conditions; however,

microexplosions were observed for mixtures more diluted by methanol (in particular, 75/25 methanol/dodecanol

mixtures). These results have suggested that in a convective flow enviroimient the internal liquid circulation

(resulting in concentration gradients [3]) may distribute the mixture components within the droplet in a manner

favorable to creating the necessary superheat conditions for microexplosions. Properties that may influence the

tendency for microexplosions to occur (e.g., droplet mixture composition, residence time, temperature, etc.), of

course, require further examination. The motivation of this work was therefore to determine the extent to which

microexplosions occur ui 75/25 methanol/dodecanol mixture flames.

In previous investigations [4,5], a comparative study of methanol and 50/50 methanol/dodecanol mixture

spray flames was carried out. Spatially resolved information on droplet size and velocity distributions was obtained

using a phase/Doppler interferometry system [6]. Some evidence for the occurrence of microexplosions was found

ia the methanol/dodecanol mixture flame. This evidence was in the form of a sudden decrease in droplet size and

velocity, and an increase in number density when compared with the methanol flame or with data obtained at different spatial positions of the mixture flame. In addition, photographic evidence in the form of particle streaks

in random directions was observed which was also suggestive of the presence of microexplosions.

521
2. EXPERIMENTAL ARRANGEMENT

Experiments were carried out in a spray combustion facility which was designed to simulate practical

combustion systems. The facility includes a swirl burner with a moveable 12-vane swirl cascade; the vanes are

rotated simultaneously to impart the desired degree of swirl intensity to the coflowing combustion air stream. The

spray facility permits examination of the effects of combustion air swirl, atomizer design, fuel type, and air preheat on spray structure and combustion characteristics. The swirl strength introduced to the combustion air is given in

terms of the swirl number, S, which is determined according to the theory outlined in Ref. [7]. Experiments were

carried out for S = 0.53 under burning conditions.

A simplex pressure-jet nozzle is located along the centerline of the burner. The fuel nozzle was operated

at total air and fuel flow rates of 64.3 kg/hr and 3.2 kg/hr, respectively. The spray is injected vertically upwards

from the nozzle, located at the burner exit. The pressure-jet nozzle provides a nominal 60° hollow-cone spray. A

stepper-motor-driven three-dimensional traversing arrangement is used to translate the burner assembly in the vertical

(Z) and both horizontal (X and Y) directions. All optical diagnostics are fixed in position about the burner assembly

so that the burner can translate independently of the optical equipment. This arrangement permits precise alignment

of the measurement volume within the spray (see Fig. 1). Measurement of radial profiles of the spray flame

properties, e.g., droplet size and velocity distributions, can then be carried out at different axial positions. Further

details on the experimental arrangement are given elsewhere [8].

Swirl Burner

loosed

3 404'

20^

Micro- PDI Computer Processor

100-

804 ^ 604

3 40i

20^

I I

Z=76.2 mm I^0

= 50.8 mm

(AAA

Z = 25.4 mm

Fuel Type

-S- methanol a- 75/25 meth/dodec

-A- 50/50 tneth/dod'

Z = 10 mm

I I I 60 -40 -20

I

I

I I T

I

I r

20 40

100

-100

^0

^60^e

[-20

60

r (mm)

Fig. 1. Schematic of experimental spray combustion

facility. BS - Beam stop; LH - He-Ne laser; PR - PDI receiving optics; PT - PDI

transmitting optics.

Fig. 2. Variation of droplet Sauter mean diameter

(D32) with radial position (r) at different

axial positions (Z) for the methanol and two

methanol/dodecanol mixture flames.

522
Droplet size and velocity distributions were measured using a single-channel phase/Doppler interferometer [6]. The instrument provides in-situ nonintrusive measurements at different positions within the spray. For

these experiments, an off-axis light collection system was used, with the optics positioned at a scattering angle of

30° (see Fig. 1). The focal length of the transmitting and receiving optics was 495 and 500 mm, respectively. A

10 mW He-Ne laser, operating at a wavelength of 632.8 nm, provided the light source for the methanol and 50/50

methanol/dodecanol mixture flames. An upgraded system, incorporating a 4W cw Ar-ion laser and operating at a wavelength of 514.5 nm, was used for the 75/25 mixture flame. The laser was operated at 400 mW for the

measurements presented in this paper. The signal data rate was kept at a relatively low value (approximately

200 Hz) by adjusting the supply voltage to the scattered light detector. Data were also acquired by optimizing the

system gain to detect the maximum number of droplets without saturating the photomultiplier tube (PMT) detectors

and thus affecting the signal-to-noise level. This was accomplished by setting the PMT voltage (for a specified laser

beam intensity) at a level commensurate with the maximum value of volume flux [9,10]. This value generally

corresponded to the instrument operating procedures recommended by the manufacturer.

3. RESULTS AND DISCUSSION

3.1 Droplet Mean Diameter, Number Density and Velocity

In this study, data for a 75/25 methanol/dodecanol mixture flame were obtained since recent single droplet

experiments under spherically symmetric burning conditions have indicated an enhanced tendency for the occurrence of microexplosions. The Sauter mean diameter (D32) and number density (N) are presented in Figs. 2 and 3 for

methanol and the two methanol/dodecanol mixture flames, described earlier, at axial positions of Z = 10, 25.4, 50.

8

and 76.2 mm. The solid boxes along the abscissa indicate the position of the burner passage walls, with the fuel

nozzle located at the centerline. Limited data are presented for Z = 50.8 and 76.2 mm because of the very low data

rate in the central region of the spray flame. The variation of D32 with spatial i>osition is typical of hollow-cone

pressure-atomized spray flames. The droplets near the spray boimdary (defined by the peaks in the D32 distributions) are larger than those near the spray centerline in the upstream portion of the spray. The profiles become more

uniform with increasing axial position. At Z = 10 mm the values of D32 are similar for all three flames near the

spray boundary and toward the center of the spray. At upstream positions, microexplosions are not expected to occur since the droplets have not experienced the necessary period of rapid heating (i.e., sufficient residence time

in a high temperature environment). Further downstream (at Z ^ 50.8 mm), the droplet mean diameter for the

50/50 mixture flame becomes significantly smaller near the spray boimdary than the methanol flame; as discussed

earlier, the reduced size for the 50/50 mixture flame is indicative of the occurrence of microexplosions. The droplet

mean size and number density (see Fig. 3) for the 75/25 mixture flame remain similar to that of methanol. These

results suggest that there is little or no evidence of the presence of microexplosions for the 75/25 mixture flame.

Radial profiles of droplet mean axial velocity (U) for the methanol and two mixture flames are presented

in Fig. 4 at Z = 10, 25.4, 50.8 and 76.2 mm. The mean velocity profiles are also indicative of the overall structure of hollow-cone spray flames, viz., the highest droplet velocities are foimd near the spray boundary. The results

indicate that the value of U decreases with higher dodecanol content near the spray boimdary. This trend may be

expected since denser fuels will create larger droplets with lower velocities for a given initial droplet momentum.

However, since the droplet mean size decreases in regions where microexplosions are expected to occur, the

decrease in droplet velocity may be attributed to rapid droplet entrainment into the air flow field and subsequent

deceleration. These trends may be further substantiated by examining changes in the size and velocity distributions,

and size/velocity correlations with spatial position, as discussed in the next section.

3.2 Size and Velocity Distributions

The droplet size and velocity distributions have been found to vary significantly with spatial position in the

investigated spray flames. A typical set of size and velocity distributions, and velocity/size correlations obtained

for the three spray flames is presented in Fig. 5 at a radial position (r) of 33.0 mm and an axial position (Z) of

50.8 mm. This radial position was chosen since it represents a position where the occurrence of microexplosions

is expected (see Figs. 2-4). The solid bars of the distributions represent the measured data on a flux (or temporal)

basis, while the open bars represent the data corrected for variations in the effective measurement volume [6]. Also

note the change in axes scales between distributions. The results indicate that the size distributions for the methanol

and 50/50 methanol/dodecanol mixture flames generally remain monomodal (see Figs. 5A and 5C); however, a
i

5 -

3S

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

1

Z = 76.2 mm

Z = 50.8 mm

Z = 25.4 mm

Fuel Type

-0- mellionol "B- 75/25 melh/dodec

-A- 50/50 meth/dodec Z = 10 mm

I I

I

I I I

I T

60 -40 -20 20

I I

I

I r

40

30-:

2(H

-10

30-

20-

10-

0^

10-

I 'I I I I I I I I I I I I I I I I I I I I I

Z = 76.2 mm

-30

-20

-10;

Z = 50.8 mm

Z = 25.4 mm

Fuel Type

-0- melhonol

-a- 75/25 meth/dodec

A- 50/50 melh/dod«c Z=10 mm

I I I

I

I I I

I T

-60 -40 -20

I I I r

r30

r-20

HO;

"I

I I I

I

I I I

20 40 60

r (mm) (mm)

Fig. 3 . Variation of droplet number density (N) with

radial position (r) at different axial positions

(Z) for the methanol and two

methanol/dodecanol mixture flames.

Fig. 4. Variation of droplet mean axial velocity (U)

with radial position (r) at different axial

positions (Z) for the methanol and two

methanol/dodecanol mixture flames.

B

35 68 DIAMETER (^m) 32 61 DIAMETER (^m)

IIH1H.I..Y—

35 68 1on DIAMETER (iim)

477-

Z

no 239- o

0- 1 lllii,,.. r

I 15 VELOCITY (m/s)

35 68 DIAMETER (^m) METHANOL

32 61 DIAMETER ((im)

35 68 DIAMETER (|im)

75/25 METHANOL/DODECANOL 50/50 METHANOL/DODECANOL

Fig. 5. Distribution of droplet size, velocity and velocity/diameter correlations for the a) methanol, b) 75/25

methanol/dodecanol and c) 50/50 methanol/dodecanol flames at r = 33.0 mm and Z = 50.8 mm.

524
significant increase is observed in the number of detected smaller droplets and a decrease in the larger sizes for the

mixture. The 75/25 mixture flame appears to be an intermediate case, i.e., abimodal distribution that corresponds

to a combination of the other two flames (see Fig. 5B).

The velocity distributions for the methanol and 50/50 mixture flames indicate a dramatic increase in the

number of lower velocity droplets for the mixture. The bimodal velocity distribution for the 50/50 mixture flame

(see Fig. 5C) may result from a combination of faster-moving larger droplets originating from the nozzle and slowermoving smaller droplets which form during microexplosions. The velocity/size correlation indicates the presence

of two separate regions, with the formation of the lower velocity region as a result of microexplosions. Further

details are presented in Refs. [4] and [5]. Again, the 75/25 mixture flame appears to be a combination of the

methanol and 50/50 mixture flames, resulting in a trimodal velocity distribution (see Fig. 5B). The velocity/size

correlation shows the initial formation of this second lower velocity region. These trends are also present at Z =

76.2 mm. The results indicate that the 75/25 methanol/dodecanol mixture flame does not appear to be a mixture

which enhances microexplosions but an intermediate case, perhaps indicating incipient formation of microexplosions.

Evidence for the occurrence of microexplosions appears to be strengthened with higher dodecanol content; the

optimum quantity of dodecanol remains to be determined. Other properties to consider in future studies include

examination of mixture composition (volatility of the mixture components), residence time, and temperature (air and

fuel preheating).

3.3 Effect of System Gain

Spray flames generally contain a wide range of droplet sizes, ranging from submicron droplets to droplet

diameters of several hundred microns. Detection of all droplets encountered in the probe volume of the

phase/Doppler interferometer can be a formidable task in dense regions of the spray. To optimize detection of

droplets in these regions the system gain has to be set such that the PMT detectors are not saturated by the larger

droplets and remain sensitive enough to detect the smaller sizes. Alternatively, multiple measurements must be

recorded and combined during postprocessing of the data to determine the appropriate statistics.

The system gain of the phase/Doppler system can be affected by the PMT voltage setting, laser power and

other parameters. It is therefore of interest to determine how the system gain influences the aforementioned results,

especially as regards the occurrence of microexplosions. Data were recorded at two different operating conditions

of the phase/Doppler instrument. In the first case, the data rate was kept low at approximately 200 Hz in order

to keep the sample duration time relatively constant across the entire radial profile. The results presented in

Figs. 2-4 were recorded for this condition. In the second case, the PMT voltage was changed to optimize detection

of the droplets as recommended by McDonell and Samuelsen; this meant adjusting the PMT voltage at each point

to achieve maximum volume flux [9]. This criterion appeared to parallel in most regions of spray flames the

procedures recommended by the manufacturer (i.e., when the PMT detectors begin to saturate).

Figures 6-8 present the results obtained for droplet mean size, number density and velocity, respectively,

at Z = 10, 25.4 and 76.2 mm. Data were recorded in the 75/25 mixture flame for the two above-mentioned

instrument operating conditions. The results indicate that, at Z < 25.4 mm, there is a significant difference in the

values of D32, N and U measured across the profiles. This is to be expected since the "optimum gain" case detects

more smaller droplets (resulting in a strong signal with a high data rate); the threshold level for the "low data rate" case is too high to detect smaller droplets. For example. Fig. 9 presents the size and velocity distributions at r =

15.2 mm and Z = 25.4 mm for the two cases. The optimum gain case (see Fig 9B) shows the detection of many

smaller droplets and development of a second lower-velocity peak in the distribution as compared to the low data

rate case (see Fig. 9A). As a result, the mean value for D32 decreases from 58.5 to 46.3 fxia, U decreases from

19.3 to 14.2 m/s, and N increases from 4.2 x 10^ to 2.2 x lO'' particles/cc for the optimum gain setting. At Z >

50.8 mm, the results indicate negligible difference between the two readings. This is as expected because of droplet

vaporization and dispersion with increasing axial position. The measurement therefore requires a high PMT voltage

to maintain a low data rate, i.e., the two instrument settings are essentially the same at downstream positions. The

importance of these findings is that at Z = 50.8 and 76.2 mm, where possible evidence for microexplosions has

been found, the instrument system gain has negligible influence upon the measured trends.

525
10(H

s

20H

I I I I I I I I I I I I I I I I I

Z = 76.2 mm

Z = 25.4 mm -

-9- Low Data Rale Q- Optimum Gain

Z = 10 mm

n I I r "T I I

I

I I r I

I I I

I

-100

E-80

|-50-

20 N

-50 -40 -20 0 20 40 60

r (mm)

Fig. 6. Variation of droplet Sauter mean diameter

(D32) with radial position (r) at different

axial positions (Z) at two different instrument

operating conditions for the 75/25

methanol/dodecanol mixture flame.

30-

20^

10^

30-

20-

10-

0-3

I I I I I I I I I I I I I I I I I I I I I I I

Z = 76.2 mm

-9- Low Oato Rate

-Q- Optimum Coin

Z = 25.4 mm

Z=10 mm

-60 -40 -20 20 40

-30

-20

H0|

_ us

--10

60

r mm

Fig. 8. Variation of droplet mean axial velocity (U)

with radial position (r) at different axial

positions (Z) at two different instrument

operating conditions for the 75/25

methanol/dodecanol mixture flame.

I I I I I

'

I I I I

I I I I ' I I I I I

Z = 75.2 mm

Z = 25.4 mm

-9- Low Data Rote W

-Q- Optimum Gain Z=10 mm

I I I

I

I

I I

I ~r I I

I

1 I P

I

I I I

I

I l~l—

I

-60 -40 -20 0 20 40 60

r (mm)

Fig. 7. Variation of droplet number density (N) with

radial position (r) at different axial positions

(Z) at two different instrument operating

conditions for the 75/25 methanol/dodecanol

mixture flame.

890-

28 54 DIAMETER dun)

VELOCmr (m/s) Fig. 9. Distribution of droplet size and velocity for

the a) low data rate and b) optimum gain

operating conditions of the phase/Doppler

system with the 75/25 methanol/dodecanol

mixture flame at r = 15.2 mm and Z =

25.4 mm.

526
The change in the mean droplet properties (i.e. , D32, N and U) with axial position is, in general, dependent

upon the instrument operating conditions. For the low data rate case, the maximum value of D32 decreases with

axial position (see Fig. 6). This can be attributed to detection of primarily larger droplets as they undergo

vaporization. This trend is contrary to that indicated by the optimum gain case (since the rapid vaporization of the

smaller droplets tends to increase the mean diameter) and to observations generally reported in the literature for

nonbuming sprays [8]. The maximum mean velocity decreases with axial position for both cases. The maximum

number density decreases with axial position for the optimum gain case (as expected) and remains relatively

unchanged for the low data rate case. The importance of these trends is that at downstream positions there are primarily larger vaporizing droplets. If there is a sudden abundance of smaller droplets, it may then be attributed

to microexplosions.

It is also important to ascertain that the laser power did not influence the results since two different

phase/Doppler systems (with lasers of different power) were used. Figure 10 presents the effect of PMT voltage

upon the droplet mean axial velocity at four different laser beam intensities of 10, 125, 200 and 400 mW. The data were acquired for the 75/25 mixture flame at r = 0 and Z = 10 mm. The results indicate that the PMT voltage

has a significant effect upon the values of U for a preset laser power. This effect is the same for the droplet mean

size and number density. It is also observed that the number density (and volume flux) does not reach a maximum

value at each measured position (see Fig. 10). A maximum value is generally achieved in regions near the spray

boundary where the data rate is high. When no maximum value is found the PMT voltage is set according to the

recommended procedures of the manufacturer. Figure 10 indicates that as the laser power increases, lower PMT

voltages are required to maintain a particular value of droplet mean velocity (or size). The use of two different

lasers with the phase/Doppler system therefore should have negligible influence on the results, provided the detector

voltage is appropriately adjusted.

in

-I—

o

o

"o) 1 >

- 5

125 \ \ 10 mW

0 -

1

\—I—I—I—I—I—I—I—1—I—I—r—I—I—I—1—I—I—I—I—I

\—I—I—I—I—I—I—I — 1

- 3

- 2

200 250 300 350 400 450 500

Voltage (v)

Fig. 10. Variation of mean droplet velocity with PMT voltage at different laser powers for the 75/25

methanol/dodecanol mixture flame at r = 0 and Z = 10 mm.

4. SUMMARY

Measurement of droplet mean size, velocity and number density have been obtained using a phase/Doppler

interferometry system. The measurements were carried out in a methanol and two methanol/dodecanol mixture

flames. The study focused on detection of microexplosions in methanol/dodecanol mixture flames. The results

obtained in these three flames indicate that the tendency for microexplosions is strengthened with higher dodecanol

content. The effect of system gain (i.e., PMT voltage and laser power) on the measurements has been found to influence dense regions of spray flames, but has a negligible effect at downstream positions where microexplosions

are expected to occur.

527
ACKNOWLEDGEMENTS

The authors would like to acknowledge the support of this work by the U.S. Department of Energy,

Conservation and Renewable Energy, Office of Industrial Processes, Advanced Industrial Concepts Division.

Messrs. Marvin Gunn and John Larson are the project monitors. One of the authors (CTA) would like to acknowledge the support of the New York State Center for Hazardous Waste Management (Dr. Ralph R. Rumer,

project manager). The technical support provided by Mr. J.D. Allen is also gratefully acknowledged.

REFERENCES

1. Wang., C.H. and Law, C.K., "Microexplosion of Fuel Droplets Under High Pressure," Combustion and

Flame, Vol 59, pp. 53-62 (1985).

2. Yang, J. C, Jackson, G.S. and Avedisian, C.T., "Combustion ofUnsupportedMethanol/Dodecanol Mixture

Droplets at Low Gravity", Twenty-Third Svmp. (Int'l) on Combustion. The Combustion Institute, Orleans,

France (1990).

3. Lara-Urbaneja, P. and Sirignano, W.A., "Theory of Transient Multicomponent Droplet Vaporization in a Convective Field", Eighteenth Svmp. (Int'l) on Combustion, pp. 1365-1374, The Combustion Institute,

Pittsburgh, PA (1981).

4. Presser, C, Gupta, A.K., Avedisian, C.T. and Semerjian, H.G., "Combustion of Methanol and

Methanol/Dodecanol Spray Flames", AIAA 90-2446, AIAA/SAE/ASME/ASEE 26th Joint Propulsion

Conf. . Orlando, FL (1990).

5. Presser, C, Gupta, A.K., Avedisian, C.T. and Semerjian, H.G., "Fuel Property Effects on the Structure

of Spray Flames", Twenty-Third Svmp. (Int'l) on Combustion. The Combustion Institute, Orleans, France

(1990).

6. Bachalo, W.D., Houser, M.J. and Smith, J.N., "Behavior of Sprays Produced by Pressure Atomizers as Measured Using a Phase/Doppler Instrument", Atomization and Spray Technology, Vol 3, pp. 53-72

(1987).

7. Gupta, A.K., Lilley, D.G. and Syred, N., Swirl Flows. Abacus, Tunbridge Wells, England (1984).

8. Presser, C, Gupta, A.K., Semerjian, H.G., and Santoro, R.J., " Application of Laser Diagnostic

Techniques for the Examination of Liquid Fuel Spray Structure", Chemical Engineering Communications,

Vol 90, pp. 75-102 (1990).

9. McDonell, V.G. and Samuelsen, S., "Sensitivity Assessment of a Phase-Doppler Interferometer to UserControlled Settings", in Liquid Particle Size Measurement Techniques: 2nd Volume ( E.D. Hirleman,

W.D. Bachalo, and P.G. Felton, Eds.), ASTM STP 1083, pp. 170-189, American Society for Testing and

Materials, Philadelphia, PA (1990).

10. Chao, J., Hwang, K. and Lee, T., "Sensitivity of User Parameter Settings of Phase Doppler Particle

Analyzer in Combusting Spray", AIAA 90-1553, AIAA 21st Fluid Dynamics. Plasma Dynamics and Lasers

Conf. . Seattle, WA (1990).

528
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 57

SPRAY CHARACTERIZATION IN AXISYMMETRIC,

SWIRLING COMBUSTOR FLOWS

M.J. Ward, J. A. Bossard, R.E. Peck, and E.D. Hirleman

Department of Mechanical and Aerospace Engineering

Arizona State University

Tempe, AZ, U.S.A.

ABSTRACT

Droplet size and velocity characteristics of burning fuel sprays in a laboratory combustor have been measured

with a single-component Phase-Doppler Particle Analyzer. Tests were conducted to investigate the effects of inlet

velocity, fuel flow, swirl quantities, and nozzle air-assist on the spray dynamic processes. The results indicate that

the air-assist atomizer produced hollow-cone sprays having a Sauter mean diameter (SMD) of about 20-25}im. The

atomization improved (smaller SMD) with increased air-assist flow or decreased fuel flow. Increasing swirl was

found to centrifuge the larger droplets comprising most of the liquid volume flux further away from the centerline. In certain cases these droplets could escape the recirculation zone and impinge the combustor wall causing poor

combustion efficiency. Size-segregated velocity measurements revealed areas of recirculation and further showed the

droplet acceleration due to the accompanying heat release.

INTRODUCTION

The development of advanced gas-turbine engines is always striving to improve combustion stability, efficiency,

and emissions. Because the fuel spray distribution strongly affects combustor performance, knowledge of spray

dynamics in reacting flows is a prerequisite to design improvements. Combustion-chamber design and performance

evaluation has heretofore been accomplished by component testing. The advent of advanced optical diagnostic

methods permits engineers to study non-intiusively spray characteristics such as droplet size and velocity.

During the 1960's and 1970's research began on the application of lasers for combustion diagnostics. The

advantages of using laser-based methods are that they are non-intrusive and enable real-time data acquisition and

processing. Bachalo and Houser [1] developed the Phase Doppler Particle Analyzer (PDPA) for measuring

simultaneously the drop size and velocity of spherical particles. The instrument utilizes an optical system

resembling a conventional laser-Doppler velocimeter (LDV) except that three detectors are located at selected spacings

behind a receiver aperture. Droplets passing through the intersection of two laser beams scatter light which produces

a far field interference fringe pattern. The spacing between these fiinges is proportional to the drop diameter. The

PDPA obtains measurements of this fringe spacing image with the use of a photomultiplier tube, which converts

the scattered light into an electronic signal. The fringes move past the detectors at the Doppler difference frequency

and produce identical sigruils but with a phase shift i^oportional to the fringe spacing.

McDonell et al. [2] made measurements with a phase-Doppler instrument in an isothermal chamber (nonreacting fuel), and a swirl-stabilized combuslCM". As expected, they found significant differences in the drop size from

reacting to non-reacting flow. McDonell and Samuelsen [3] investigated symmetry of an air-blast atomizer using a two-component instrument They measured velocities in both the axial and azimuthal directions inside a combustion

chamber. Brefia de la Rosa et al. [4] studied the effect of swirl on the dynamic behavior of drops and on the velocity

and turbulence fields of an isothermal spray using a two-component phase-Doppler particle analyzer (PDPA).

Recently, Hardalupas et al. [5], used a similar technique to study spray characteristics of unconfined swirl-stabilized

flames.

The purpose of the present study is to apply the PDPA to obtain some basic data on droplet ballistics in

complex flows found in practical gas-turbine combustors. This information can then be used to clarify the relevant

spray combustion processes and assist in the development of computational tools for combustor design.

529
EXPERIMENTAL

The continuous-combustion test facility, shown schematically in Fig. 1, was used to conduct the present

expwiments. The combustor consists of a 78-mm dia, x 350-mm long stainless-steel, flanged pipe with 305mm x 25mm x 5mm quartz glass side-panels for optical access. A Delavan siphon-type nozzle (P/N 30609-2) was used for

on-axis fuel injection into a co-flowing air stream (up to 0.2kg/s) entering at aunospheric pressure and 300K. Two

51 -mm o.d. x 24-mm i.d. vane-type swirlers (60°; 45°) or a 38-mm o.d. disk were mounted to the nozzle for flame

stabilization.

Particle diagnostics included an Aerometrics PDPA system consisting of a transmitter Model XMT-1 100-4S

with a lOmW He-Ne lasCT, motor controller Model MCB-7100-1, receiver Model RCV-2100, signal processor Model

PDP-3100, and data management system Model DMS-4000-5 including a Compaq 386 computer. Instrument

accuracy was assessed in an earlier study [6] where the results for two different PDPA optical configurations were

found to be within a total experimental uncertainty of 3.5% in sizing monodisperse droplets. In order to measure the

axial component of velocity, which coincides with the horizontal spray axis, the transmitter and receiver are each

mounted at 15° on a movable optical table that enables mapping the flowfield. The PDPA configuration used

throughout these experiments used a 300-mm coUimating lens and 500-mm transmitting lens, with three tracks

providing the following size ranges: 0.9 to 127.8 ^im, 2.7 to 258 ^im, and 3.8 to 520 |im. The velocity ranges were

first chosen so that there were no velocity over rejections; then the velocity offset was increased to eliminate the

velocity under rejections. In some earlier work [6] it was found that the SMD increased dramatically as the PMT voltage is reduced. The

explanation was that the larger drops are seen at the lower voltage settings while it becomes increasingly more

difficult to detect the small drops. Figure 2 shows the SMD's versus PMT voltage with maximum diameters of

100-150 [un, 250 ^lm, and 520 jim. The 100-150 case found the SMD decreasing to about 13 im at 400V then

increasing to 32 at 600V showing that the larger the PMT voltage, the larger the mean diameters. This

discrepancy may be caused by the presence of non-spherical droplets or possibly due to detection of stray light

reflected by small droplets on the edge of the measurement volume at higher voltage settings. These results suggest

that measurements within 10 mm of the nozzle exit are subject to increasing uncertainty, apparently due to droplet

nonsph^icity or multiple scattering, that the PMT voltage will effect the results of all measurements, and that the

data acquisition process must include adjustments for the on-line indications of out-of-range size and velocity

measurements.

RESULTS AND DISCUSSION

This study is designed to determine the structure of burning fuel sprays in a laboratory combustor. Experiments were conducted for various inlet air velocities (3-17 m/s), heat inputs, swirl numbers (0.77 and 1.33), and atomizer

air flow. The combustor was operated on Jet-A fuel at equivalence ratios ranging from 0.04-0.20. The test matrix is presented in Table 1. All valid measurements along a radius are reported at four axial stations. The PMT voltage was initially set at 400V and the size range at Dmax of 125 [im. Minor adjustments to these settings were made

during the course of the test as needed.

Effect of Fuel Flow

Figure 3 shows the Sauter mean diameter (D32) as a function of the radial and axial locations for cases 1 and 2

with two different fuel flow rates (equivalence ratios). The higher fuel flow rate produced largw SMD's indicating

reduced atomization effectiveness at constant air-assist flow. At 20.32 mm from the atomizer tip, the higher

equivalence ratio droplet SMD's are larger near the center of the combustor then become smaller than the SMD for

the low equivalence ratio spray further from the centerline of the combustor. Because the higher fuel flow has higher

flame temperatures the smaller droplets are consumed faster. The peaks in the curves are indicative of the rapid

ccMisumption of smaller droplets in the flame zone. Figure 4 presents size-specific droplet velocity profiles. The lower equivalence ratio spray has a wider

recirculation zone, the extent of which is determined partly by the relative momentum of the liquid jet and the

surroimding swirling air. At the centerline and near the wall all drops have about the same velocity. In the swirl

region the centrifuge effect separates the particles according to size with the larger droplets having a greater velocity

at a given radius.

Similar trends to the above were observed wh«i varying the atomizer air flow (cases 2 and 5). Higher nozzle air

flows produced finer droplets with higher velocities and a wider recirculation zone.

530
Effect of Swirl

Cases 1 and 3 are identical operating conditions but the aerodynamics were changed by using different swirlers.

Case 1 was run with a 60° swirler (S=1.33); case 3 with a 45° swirler (S=0.77). Figure 5 shows the effect of swirl

number on the Sauter mean diameter. The 60° swirler produces a larger circumferential velocity component which

tends to bend the spray away from the centerline. There is a local peak of 20-30 \im SMD in the radial profile

between radial locations of 5-10 mm for both swirlers. At large radii, the 60° swirler yields larger SMD's. The

SMD gradually increases at increasing axial distance from the nozzle due to preferential consumption of smaller

droplets. Figure 6 shows the particle velocities for droplets within 5 and 30 pim size bands. The results show that

the recirculation zone is much stronger with the 60° swirler.

Reacting and Non-reacting measurements

The Sauter mean diameters measured as a function of axial and radial locations are shown in Fig. 7. The SMD

for the non-reacting spray (Case 6b) stays nearly constant at about 25-30 jim at most measurement locations, but

shows a slight tendency to increase near the spray periphery. The SMD measured in the reacting flow (Case 6a)

condition is about the same but the spray boundaries are much narrower, the non-reacting case having the larger

droplets. Because the small droplets are being consumed during reaction, the effect is to raise the SMD.

The mean axial velocity of the 5 and 30 \im droplets are shown versus radial location in Fig. 8. The reacting

flow is accelerating faster than the non-reacting flow due to the heat release from combustion. The 5 \im droplets

show much wider velocity fluctuations and give a better indication of the gas-phase velocity distribution. The

reacting case shows an expanded recirculation zone extending from 10 mm to 16 mm. Near the combusior wall both

the 5 and 30 ^im droplets approach the average main stream velocity (8 m/s) for these cases. The non-reacting case remains fairly flat at the more distant radial locations. The 5 [im droplets are slower at the centerline reflecting the

effect of the recirculation and their velocity increases as the drops move further out from the centerline.

The volume flux measurements are shown in Fig. 9. The non-reacting case had a larger volume flux at all locations as would be expected, since the reacting droplets are being consumed and therefore are not being measured.

The data indicates that about 80% of the incoming liquid (spray) is already consumed at the 20.32 mm axial location.

This trend must be considered when evaluating the size and velocity measurements previously discussed. The

number density is reduced for the reacting spray, therefore the SMD and velocity measurements represent fewer

droplets. The reacting flow measurements are detecting only large droplets that pass through the reaction zone. Non-reacting droplet data at the 50.8-mm axial location are quite scattered. The integrated liquid volume flux for the

non-reacting case compares favorably to the measured fuel flowrate at axial stations beyond one pipe radius

downstream of the nozzle exit

CONCLUSIONS

A Phase Doppler Particle Analyzer has been used to obtain real-time, point-by-point data on droplet size,

velocity and liquid volume flux in a confined, turbulent spray flame. The Sauter mean diameter for these hollow- cone sprays produced by an air-assist atomizer was typically 20-25 jim. The effect of increased swirl air was to shift

the volume flux radially outward. Increased fuel flow or decreased atomize air flow produced sprays having larger

droplets. The velocity data was grouped by droplet size and the small (5 jim) droplets were found to track the gas

velocity even in reverse flow regions while the larger droplets were centrifuged away from the centerline for the

swirling flow conditions. The volume flux measurements in the non-reacting spray were integrated and compared to the fuel flow in the system. At one pipe radius downstream of the nozzle, the measured mass flows were within 9%

of the known fuel input

The Phase Doppler Particle Analyzer is a useful tool for spray diagnostics in reacting and non-reacting flows.

Measurements too close to the nozzle exit introduced errors in the droplet size and volume flux due to the high

number density of droplets, and possibly from measuring the break-up of the fuel ligaments which are non-spherical.

The instrument output was sensitive to the User settings, and a procedure to determine the appropriate values was

developed and used throughout the experiments.

Acknowledgments

This research is sponsored in part by a grant from the Allied-Signal Aerospace Company (Garrett Engine

Div./Auxiliary Power Div.). MJW was also the recipient of a Garrett Fellowship to supjK)rt graduate studies.

Thanks are also extended to Tony Bruner, Steve Kenney, Eric Krinke, Makarand Kulkami, Keith Parsons, and Steve

Thomas for their assistance in hardware design, instrumentation and software development and data acquisition.

531
REFERENCES

1. Bachalo, W.D., Houser, MJ., "Phase/Doppler Spray Analyzer for Simultaneous Measurements of Drop Size

and Velocity Distributions", Optical Engineering , Vol. 23, pp. 583-590 (1984).

2. McDonell, V.G., Wood, CP., and Samuelsen, G.S., "A Comparison of Spatially-Resolved Drop Size and Drop

Velocity Measurements in an Isothermal Chamber and a Swirl-Stabilized Combustor", Twenty-First

Symposium (International) on Combustion, pp. 685-694, The Combustion Institute, Pittsbiu'gh, (1986).

3. McDonell, V.G., Samuelsen, G.S., "Influence of the Continuous and Dispersed Phases on the Symmetry of a Gas Turbine Air-Blast Atomizer", ASME Paper No. 89-GT-303, presented at the Gas Turbine and Aeroengine

Congress and Exposition, Toronto. Ontario, Canada, (1989).

4. Brefla de la Rosa, A., Wang, G., and Bachalo, W.D., "The Effect of Swirl on the Velocity and Turbulence Fields

of a Liquid Spray," ASME P^r No. 90-GT-32, presented at the Gas Turbine and Aeroengine Congress and

Exposition, Brussels, Belgium, (1990).

5. Hardalupas, Y., Taylor, A.M.K.P., and Whitelaw, J.H., "Velocity and Size Characteristics of Liquid-Fuelled

Flames Stabilized by a Swirl Burner", Proc. R. Soc. Lond. A, Vol. 428, pp. 129-155 (1990).

6. Parsons, K.D., "Characterization of the Performance of the ACTometrics Phase Doppler Particle Analyze", M.S.

Thesis, Arizona State University, Tempe (1990).

Table 1 Test Matrix

Case

No.

Avg. Inlet

Air Velocity

[m/sec]

Equivalence

Ratio, <I>

Swirl

Angle

[degrees]

Air flow in

Atomizer

Comments

[kg/sec]xl03

12

3

4

5

6

6a

6b

2.9

2.9

2.9

8.0

2.9

2.9

8.0

8.0

0.14

0.20

0.14

0.05

0.19

0.16

0.04

0.04

60

60

45

60

60

0

0

0

0.37

0.37

0.37

0.47

0.47

0.25

0.30

0.30

reacting

non-reacL

532
533
E

E,o

u

c

«

v> 5

"5

^5

CO

40

35

30

25

20

15

10

5

0

f I'"-

0.2t

1 A A

Qt

0

m

3A I D

<;

A

A

A c )>-

0.39

f A I

ID C

OO

AA

cjF

i A/iL o

t

i .oa— 0—

1/0 O.Si

o

CK5

Q]

It \ 00 A

c r o

1 i/D>o.es -H

A CiimI 1

>m2,(

i i O Ci 0 /

k ( A

A A o

? C

.o

; -i

DO

A

AC

—t >—

O

1 0 20 30 1 0 20 30 10 20 30 1 0 20 30 40

Sauter Mean Diameter [^m]

Fig. 3 Effect of fuel flow on SMD

40

35

•g- 30

0 25 o1 20 5

« 15

(0 CC 10

5

0

x/0>0. A

26

cro

aacx- nO

lA (

3 —

1

OD

AO

A ( ++

-es>—

ICD A

D (o aI

•fCS Afi ——1-DH-

-1 0 10 20 30

i/o>o. 3«

(a

(DD

i

est«-

1 cm

A

t & »

•*

4- a i A

n

CD

t. O

4

t

n: M1 ^

ri (EDA

H—

c

OA

a:1S|

AicOl —1—Lm—

1 0 10 20 30

1 I/D.0.S2a

i BTT*

fm

C L O

{ m

n

iOa

tz

4-

1 —

< f T K mm.

C2k.

o <Z> A

AA1 0

i/D.O.

1 ' -

55 A

1 1

1 1

1 1 04.20, Sfim

£3f CD

0+

CCD

/HD o

a-02 3,30)un Q - #o

D O

£ Q flCD

44- n ( 3A

- + n n a

CD

^

10 20 30 1 0 1 0 20 30 40

Velocity (m/s]

Fig. 4 Effect offuel flow on droplet velocities

334
40

35 ^ 30

E

E" 25 e

u

« 20

D

rt 10

5

f 0

i/D •0.21 A A

O

A<

h

g>A

AJ

I/O •O.J ) A

0

i

o

c AA

o A

(

c A

n In

x/D •O.Si

O

It

oo

(

Ai AA o JL

\

i a

iO

A

o

ola...

—A

o

—

C

t—

1 0 20 30 10 20 30 1 0 20 30

t/D

1 I .0.6$

1 A

.0

Cut I.VmI

1 1 kA

A

\ A

A

£ A3

O

kO

k

< Q

i

A aO

TO

Tt1 0 •20 30

Sauter Mean Diameter [nm]

Fig. 5 Effect of swirl on SMD

40

35

E 30

£

e> 25 u

c

nw 20

1 15

eg ^=10

i/D>0.2(

_£4

At D

O

AA

5

0

• 1 0 10 20 30

z/D'O. s»o

- +

H

«^ n

d

+

¥

J ^ «3

A.

> C

nrw- A A U 1

o-

>+-

M3

COA

i/D>0. S2

•

- -a

«ft- C

+

to

A O

a

o

& 0

&

trncrw-

—ftii

AO")

o

J-*—

Tn +

T-l 1 1

z/D.O.SS

-fib

o

A O

I A tiI I

' O 4S* w<rt. X)iffl

I I

• I S9* wnrt. X)im

4- DkTO

A

HO

ITH-O

D10 20 30 '1 0 10 20 30 1 0

ceo —Q-CD-t

1 0 20 30 40

Velocity (m/s]

Fig. 6 Effect of swill on droplet velocities

535
E

e.o

o

cn

CO

rt •6n

cc

40

35

30

25

20

15

10

5

0

1 I/O.

' 1 0.2t

o

O

CO

c

i

t

A

o

-cf

cx

-o3 -o-

' 1 ilD-

' 1 0.39

O

rt o O

A

\,

i ~£? >

ea —i r-O

0L

1 i/D>1 O.Si

>

o D

Ai a

0

a

.

ea ex

tD( )

3

\

-6-

)O-

'IT x/o-o.es

1

1 A C«Mta.n

1 1 M, inn-iwi OQ

AO

A

•QC

C

(T

O A

& C

33V

O

AC c —

0 1 0 20 30 40 50 60 70 0 1 0 20 30 40 50 60 70 0 1 0 20 30 40 50 60 70 0 1 0 20 30 40 50 60 70 80

Sauter Mean Diameter Qim]

Fig. 7 Effect of combustion reaction on SMD

40

35

•g- 30 E

uI 20 «

n

15

10

x/D«0 .26

A ran- A

+

30

>

H—

1

O

+

pjA3O— •

< o

>o—

A

it—

^ 1

1

1 AS

po

h

p

AS—O- air- +f—

c

0

-CK

A —tr-

'1 0 0 1 0 20 30 40

i/o>o .52

C

C

xp o

A O

£. 4r3An +D

-*a

A C

<i>o- + D A

o A

x/D« I.E5

1 1

'

1 1

1 1 tn

(

o

•c

[2 Han

4a X)

>0-A- <2 A

-O—

'1 0 0 1 0 20 30 40 .1 0 0 1 0 20 30 40 - 1 0 0 1 0 20 30 40 50

Velocity [m/s]

Rg. 8 Effect of combustion reacticxi on droplet velocities

536
E

£

a>u

c

aw

TO«

n

40

35

30

25

20

16

10

5

1 1 i/0> |—

1

I.2S

y

6

i

Q>

ko

a—

o

a ,. —

D

3—

O

A

ino— —c: 0—

O

0^ 0.00

][

).3t

i>

oa..- A O

a m_

o —

Ik —cx 3

30

A CD

j

1

1

).S2

o—e;

0

oo

i/D><

1 ts

1 1 1 h C*M U. nmetrf

1 1 1 O Cam

1

iooo

r

' >opar rs CO 0

) o—

c

oc

-o— —a> 0

0.02 0.04 0.00 0.01 0.02 0.00 0.01 0.02 0.00 0.01 0.02 0.03

Volume flux [cc/cm2]

Fig. 9 Effect of combustion reaction on volume flux

537

ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 58

SPARK-IGNITED SPHERICAL FLAMES

PROPAGATING IN A SUSPENDED DROPLET CLOUD

K. Nakabe*, Y. Mizutani*, F. Akamatsu',

M. Fuchihata^ and S.H. El-Emam*

Faculty of Engineering, Osaka University, Osaka, Japan

^Mazda Motor Corporation, Hiroshima, Japan

*Mansoura University, El-Mansoura, Egypt

ABSTRACT

A droplet suspension of liquid fuel produced by an ultrasonic atomizer

was spark-ignited and the flame ball propagating outwards was observed in

detail in order to elucidate the mechanism of flame propagation and the

complicated group structure of spray flames. The chemiluminescence of OH

radicals and the flame luminosity in C^-radical emission band were monitored

at a fixed point during the passage of a flame front. The flame structure was

visualized by processing the flame images taken by a water-cooled CCD camera

system which had an extremely high sensitivity, a wide dynamic range and a

sufficiently wide spectral range between 200 nm and 1000 nm in wavelength. In

addition, the statistical behavior of the luminous flame contour was

visualized by processing the images taken from the successive frames of highspeed photography. It was found that a nonluminous flame was propagating

continuously through a gas-phase mixture followed by a luminous flame

frequently having a discontinuous geometry. Due to the inherent spatial

nonuniformity of the droplet cloud, a number of small-scaled droplet clusters

burning brightly in diffusion combustion mode were observed behind the flame

front propagating in premixed combustion mode. This fact implies that the

group structure of a spray flame is not simple but complex.

INTRODUCTION

The process of flame propagation in a liquid-fuel spray is apparently

statistical and discontinuous as far as the behavior of the luminous portions

of a flame is observed by a photographic method, in paticular, by high-speed

photography [1 - 4]. The mechanism of flame propagation, however, should be

elucidated not only from the behavior of luminous portions of the flame but

also from the one of nonluminous portions. The authors have found that the

instantaneous images of a spray flame at the OH-radical emission band

(corresponding to the nonluminous portions) and at the C^-radical emission

band (corresponding to the luminous portions) are significantly different in

shape from each other [3,53, and have successfully developed a technique to

determine the burning velocity of a spray from the time-averaged image of OHchemiluminescence [3]. Unfortunately, no information has been obtained on the

detailed dynamic behavior of the nonluminous portions of a spray flame. In

539
addition, the authors previously observed the detailed structure of spray

flames using an optical technique [6], and found that a spray flame had a

group structure, not simple one but complex one.

In the present study, a flame ball propagating outwards through a droplet

suspension was selected as the object of optical observation for the

elucidation both of the mechanism of flame propagation and of the complex

group structure of spray flames, in expectation of the easiness of detailed

observation due to its slow movement. In this case, the instantaneous image

of a flame ball gives us some information on the detailed structure of the

flame at least in the initial stage of its growth.

A droplet suspension of liquid fuel produced by an ultrasonic atomizer

was spark-ignited and the flame ball propagating outwards was observed in

detail. The chemiluminescence of OH radicals and the flame luminosity in C^-

radical emission band were monitored at a fixed point during the passage of

the flame front. The flame structure was visualized by processing the flame

images taken by a water-cooled CCD camera system which had an extremely high

sensitivity, a wide dynamic range and a sufficiently wide spectral range

between 200 nm and 1000 nm in wavelength. In addition, the statistical

behavior of the luminous flame contour was visualized by processing the images

taken from the successive frames of high-speed photography.

It was found that a nonluminous flame was propagating continuously

through a gas-phase mixture followed by a luminous flame having a ragged

discontinuous geometry. Due to the inherent spatial nonuniformity of a

droplet cloud, a number of small-scaled droplet clusters burning brightly in

diffusion combustion mode were observed behind the flame front propagating in

premixed combustion mode. This fact implies that the group structure of a

spray flame is not simple but always complex.

EXPERIMENTAL APPARATUS AND PROCEDURE

amp.

Fuel

CC

F

L

PC

S

TC

The experimental apparatus is shown in Fig.l. A kerosene spray was

produced using an ultrasonic atomizer (the resonance frequency: 18.5 kHz)

installed at the top center of a vertical square duct of 280 mm x 280 mm and

was made to fall down freely and

to mix with the air induced

through the top end around the

atomizer. The Sauter mean

diameter of spray droplets

varied from 57 to 66 /im as

the fuel injection rate was

varied from 2.13 cm /min to 4.44

cm /min. A spark gap of 4.0 mm

wide was located at 400 mm

directly below the atomizer tip,

and the spray was ignited by an

intermittent spark of ca. 2

duration.

The optical system

flame imaging is also shown

Fig.l. After the preset delay

period from the start of

discharge, the mechanical iris

shutter S was opened for 1 ms by

the timing circuit TC. The

ms

for

in

Spark gap

S F

Camera controller

Optical filter

Lens

Personal computer

Shutter

Timing circuit

CCD

Electrodes

TC a

Fig.l Experimental apparatus

and CCD camera system.

540
Bragg cell Beam spliMer

Dichroic mirror

Filter Lens

Mirror

Pinhole

Photomultiplier

Collection lens

Transmission lens

PMC2

Fig. 2 Optical system for point monitoring.

flame image was focused onto

a CCD array by a UV Nikkor

lens L after passing through

an exchangeable interference

filter F (being removed in

the present study). The CCD

camera system (Photometries,

CH220, CE200, CC200) was of

water-cooled highsensitivity type, having an

extremely wide spectral

range from 200 nm to 1000 nm

in wavelength and a

sufficiently wide dynamic

range of 14 bits. The range

of sight was confined to

55 mm(H) x 80 mm(W) by the

quartz window installed in

the duct wall.

The optical system for

local continuous monitoring

of the velocity and diameter

of droplets and the light

emissions in OH- and C^-

radical emission bands is

shown in Fig. 2. The LDA

with a He-Ne laser was of

dual-beam, off-axis forwardscattering type, the offset

angle of which was 5

degrees. The light of a HeNe laser scattered by

droplets was transmitted

through a dichroic mirror

converted into the droplet velocity

The light of an Ar laser scattered

was reflected by DM^ , detected by a photomultiplier PI^^^,* ^J^d converted

the droplet diameter signal by a low-pass filter and a root circuit.

As illustrated in Fig. 3, the control volume of the He-Ne LDA

pseudo-ellipsoid of 0.182 mm diam. and 1.87 mm long. The beam from

laser was once expanded up to 7 nim diam. and made to pass through a pinhole of

1.5 nrni diam. to realize a uniform intensity distribution. The axis of the Ar

laser beam was made coincident with the major axis of the pseudo-ellipsoid.

DM

Fig. 3 Arrangement of laser beams,

detected by a photomultiplier PM

He'

and

signal by a counter-type signal processor,

by the same droplets, on the other hand,

into

was a

the Ar*

ijor axis of

If, therefore, only the scattered signals of the Ar laser coincident with the

velocity signal were picked up,

be obtained simultaneously.

The light emitted by the

beam axis was collected by a

the diameters and velocities of droplets could

flame in the direction at 5 degrees from the

quartz lens CL . The range of

pseudo-cylinder of 0.1 mm throat diameter

OH-chemiluminescence signal was

detected by a photomultiplier PM

the

respectively, whereas the light emission

detected by P'^q2 through an interference

sight was a

with^apex angle of 2 x 3.6 degrees,

reflected by a dichroic mirror DM^ and

through an optical interference filter F,

mirror DM^

OH — filter . peak wavelength and half-value width of which were 308.5 nm and 18.0 nm.

in C_-band

filter

C2'

was transmitted by DM^ and

the peak wavelength and

541
half-value width of which were 516.9 nm and 0.9 nm, respectively. The latter

signal was biased by the solid emission from soot having a continuous

spectrum.

EXPERIMENTAL RESULTS AND DISCUSSIONS

_ Figure 4 shows the distributions of mean diameter (SMD) d, mean velocity

U and velocity fluctuation (RMS) u' of the droplets passing through the

horizontal cross section at the spark gap_for the injection rates of liquid

fuel of (a) 4.44 cm /min and (b) 2.13 cm /min, where r denotes the radial

distance from the axis of the atomizer. Both the mean diameter and the mean

velocity are high around the axis decreasing outwards, whereas the velocity

E

a.

IT)

ine

inE

80.0

70.0

60.0

50.0

AO.O

0.6

0.5

0.4

0.3 0.02

0.01

0.00

E

IT3

E

80.0 r

70.0

60.0

50.0

40.0

0.6

0.5

0.4

0.3 0.02

U)

0.01

0.00

-A—

10 15 mm

20 10 15 mm

20

o 3

(a) Injection rate of 4.44 cm /min. (b) Injection rate of 2.13 cm /min.

Fig. 4 Distributions of Sauter mean diameter, mean velocity and velocity

fluctuation of droplets.

50 100

d jjm

(a) Injection rate of 4.44 cm /min. (b) Injection rate of 2.13 cm /min.

Fig. 5 Size distribution of droplets at spark-gap position.

542
* Table 1 Properties of droplet clouds

Injectj^on rate Sauter mean diameter Population density Average interval

cm' /min /zm particles/cm mm

4 44 65.2 653.5 1.43

2 13 57.0 385.0 1.71

* Measured at the position of spark gap.

10mm

(a) Low-contrast image. (b) High-contrast image.

Fig. 6 Two aspects of a flame ball with different

tones

.

fluctuation is as low as

ca. 2 % around the axis

slightly increasing

outwards

.

Typical examples of

the size distribution of

droplets are shown in

Fig. 5, where Jn/Ad/n

denotes the number

fraction of droplets in

a unit size span (1 jum)

.

It is noticed that the

size range is rather

narrow and the mass

fraction of the droplets

smaller than 30 ^wm is

extremely small. This

is the characteristic of

ultrasonic atomizers.

The properties of the

droplet clouds are shown in Table 1.

Typical examples of flame image taken by the CCD camera system are shown

in Fig. 6. (a) and (b) are the images of the same flame ball, but the tone

(the contrast between light and shade) has been changed on displaying. Due to

the extremely wide dynamic and spectral ranges of this camera system, the

image changes its appearance completely as the displaying tone is changed. The

peripheral flame front burning in premixed combustion mode with intense OHradical chemiluminescence is visualized in Fig. 6(a). The contour is slightly

vague but continuous. Droplets illuminated by the luminous flame are seen

around the vapor flame front. A number of droplet clusters various in size

and liiminously burning in diffusion combustion mode are involved within the

closed contour of the continuous vapor flame front as seen in Fig. 6(b), where

the horizontal dark zones seen on both sides in the upper portion of the flame

ball are the electrodes. The interval between the vapor flame front and the

luminous flame zone is considerably shorter in the upper portion of the flame

ball, where droplets are falling into the flame ball, than in the

portion.

Then, the light emission signals in OH- and C^-radical emission

were monitored at 25 mm below the spark gap while "the bottom of the flame

front was passing by the position. Typical examples of the signal patterns

are shown in Fig. 7 for fuel injection rates of (a) 4.44 cm /min and (b)

2.13cm /min. The emission in OH-radical band is almost pure chemiluminescence

of OH radicals, whereas the major fraction of the emission in C„-radical band

is the solid emission from soot with a continuous spectrum and the

chemiluminescence of C^ radicals is almost hidden behind the background noise.

lower

bands

543
nA

IC2

50

0

/

'V/

I OH

5

0

- J. SlV^l*^ ."•A.-4iw*»Vv.....l . .. . . A .

1 1

1 1 1 0 20 AO 60 80 100 120

t ms

(a) Injection rate of 4,44 cm /min.

nA

Ic2

50

0

/'uj

,w,^.i,^„.,...„^'*'^"''''''-v"'''AA.,_„

I OH

5

0 . >. .... A ../... s JmM: k'h

1 1 1 1

1 0 20 AO 60 80 100 120

t ms

(b) Injection rate of 2.13 cm /min.

Fig. 7 Typical light emission signal patterns.

As compared the emission signals in OH- and C -radical bands with each

other, the interval between the rise-up points or both signals is a few

milliseconds regardless of the fuel injection rate. Peaks appear in the C„-

band signal intermittently whereas the OH-band signal has usually a single

peak gradually descending thereafter. This fact implies that a nonluminous

flame propagates downwards in the forefront and that droplet clusters burn

behind the front in diffusion combustion mode with soot formation. No

nonluminous flame burning in premixed combustion mode exists in the central

region of the flame ball with some exceptions. This is coincident with the

appearances of flame image shown in Fig. 6.

Finally, the statistical behavior of the luminous flame burning in

diffusion combustion mode was observed by high-speed photography combined with

image processing. For that purpose, the images taken from a few successive

frames were binarized with the threshold adjusted at the luminous flame

contour level.

Typical examples of the„binarized images are^shown in Fig. 8 for fuel

injection rates of (a) 4.44 cm /min and (b) 2.13 cm /min, where bold and fine

solid curves represent the flame contours at 40 ms and 46 ms, respectively,

after the start of discharge. The luminous flame expands principally

downwards due to the movement of droplets and air. Here, the indicated

portions of 11 mm x 15 mm, where the typical behavior of luminous flame was

observed, were cut out from Fig. 8, and the movement of luminous flame contour

was observed from frame to frame as partially shown in Fig. 9.

In case of a large fuel injection rate shown in Fig. 9(a), the luminous

flame contour is rather continuous and smooth and it moves slowly. In case of

a small fuel injection rate shown in Fig. 9(b), on the other hand, the contour

is discontinuous and ragged, and it moves rapidly and discontinuously . The

difference between the both cases may be ascribed to the difference in the

population density of droplets (see Table 1). Some investigators observed

similar phenomena [2 J, and called it "relay type flame propagation". It has

544
1 0 mm

J

t=AO ms

t=A6 ms

(a) Injection rate of 4.44 cm /min. (b) Injection rate of 2.13 cm /min.

Fig. 8 Binarized flame images and their movements.

1 0 mm

t = 4 0 ms 42 ms

3

(a) Injection rate of 4.44 cm /min.

4-A ms

1 0 mm

t = 40 ms 42 ms 4-4 ms

(b) Injection rate of 2.13 cm /min.

Fig. 9 Successive frames of a luminous flame image.

545
been found, however, by the present and previous studies [5] of the authors,

that this is not real flame propagation but it is only the

random/discontinuous expansion of the luminous flame zone behind the

nonluminous flame front.

CONCLUSIONS

A droplet suspension of liquid fuel produced by an ultrasonic atomizer

was spark-ignited and the flame ball propagating outwards was observed in

detail in order to elucidate the mechanism of flame propagation and the

complicated group structure of spray flames. The results obtained are

summarized as follows:

(1) A nonluminous flame front burning in premixed combustion mode

propagated continuously through the vapor-air mixture formed in front of the

flame followed by a luminous flame burning in diffusion combustion mode. The

time interval between the passages of both flames was a few milliseconds

regardless of the fuel injection rate.

(2) A number of droplet clusters various in size and luminously burning

in diffusion combustion mode were involved within the closed contour of the

continuous vapor flame front. No nonluminous flame burning in premixed

combustion mode existed in the central region of the flame ball with some

exceptions

.

(3) The apparent relay transfer mechanism of luminous flame contour was

not a real flame propagation but it was only the random/discontinuous

expansion of the luminous flame zone behind the nonluminous flame front.

(4) The group structure of a spray flame is not simple but complex.

REFERENCES

1. Polymeropoulos , C.E. and Das, S., "The Effect of Droplet Size on the

Burning Velocity of Kerosene-Air Sprays", Combustion &. Flame, Vol. 25, pp.

247-257 (1975).

2. Ashgriz, N. and Yao, S.C., "Combustion Studies in Fuel Rich Idealized

Sprays", PhysicoChemical Hydrodynamics, Vol. 8, pp. 335-347 (1987).

3. Nakabe, K, Mizutani, Y. , Hirao, T. and Tanimura, S., "Burning

Characteristics of Premixed Sprays and Gas-Liquid Coburning Mixtures,

Combustion & Flame, Vol. 74, pp. 39-51 (1988).

4. Queiroz, M. and Yao, S.-C, "A Parametric Exploration of the Dynamic

Behavior of Flame Propagation in Planar Sprays, Combustion & Flame, Vol.

76, pp. 351-368 (I989).

5. Nakabe, K. , Mizutani, Y. , Akamatsu, F. and Fujioka, H. , "Simultaneous

Visualization of Droplet Clusters, Chemiluminescence and Spectral

Luminosity in Spray Flames", ASME-JSME Thermal Engineering Joint

Conference, Vol. 5, pp. 71-78, The American Society of Mechanical

Engineers, New York, N.Y. (1991).

6. Nakabe, K. , Mizutani, Y. , Hirao, T. and Fujioka, H. , "An Experimental

Study on Detailed Flame Structure of Liquid Fuel Sprays with and without

Gaseous Fuel", to be published in Combustion & Flame.

546
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 59

SOME STUDIES ON FLAME INTERFERENCE EFFECTS OF BURNING

METHANOL DROPLETS

B. Roy, R. Chandrasekhar, S. Bhattacharyya and A.K. Ghosh

Department of Aerospace Engineering

Indian Institute of Technology

Kharagpur, India

ABSTRACT

The present study deals with experimental investigations on flame

interference effects involving two burning droplets placed at different interdrop

distances. The two drops of identical diameter have been simulated by a porous

sphere technique with independent fuel feed control. Provision was also made to

rotate one of the spheres about its horizontal axis up to 120 rpm for studies on

drop rotation effects. The data generated have been compared with those obtained

from single non-rotating droplet combustion experiments under identical

conditions. The experimental measurements include fuel consumption rates, flame

height, flame temperature and composition. The results do reveal some effects of

flame interference on burning behaviour of droplets.

INTRODUCTION

Drop rotation and flame interference effects seem to have received little

attention in the numerous studies carried out so far on droplet combustion

(Ref. 1) While drop rotation has been discussed in Ref. 2, some limited studies

on interference effects have been reported in Refs. 3, 4 and 5.

The porous sphere technique has been used in the present study with

methanol as fuel to study drop rotation and flame interference effects. This

study considers interference effects on flame height, flame temperature and

composition in addition to the burning rate. In earlier studies, attention was

mostly devoted to interference effects on burning rate.

EXPERIMENTAL SET-UP AND MEASUREMENTS

The experimental facility is based on the well-known porous sphere

technique used by several investigators earlier (e.g., Ref. 6) and need not be

discussed in detail. Only some specific aspects of the facility will be

stressed. The consumption rate for a stable flame has been measured as

volumetric rate in cc/sec. instead of mass consumption rate. For flame

interference experiments , two identical spheres of 6 mm diameter with independent

feed control have been used with a provision to vary interdrop distance (IDD)

.

IDD is the distance between the centres of the two spheres and expressed in terms

547
X

yi2.0

Top sphere,! QOzLO

single sphere

( 0,60,120 rpm)

^ 6.0 o

I

3.0

^ 0

18.0

Top sphere ai - IDDi2D H

To p sphere

Bottom sphere

AIRFLOV ^ DIRECTION 2

o

a >

\Lef t sphere,! DD- 2 D

Left sphere, IDD--l-5D\_^.^-_.

'"^eft sphere, I DD= 1 D

-^-Ri ght sphere,

IDD= 1-5 D

Right sphere,! DD- 2D

AIR FLOW

DIRECTION

n ISO

o

X 1 2.0 \

u

9.0

a

6.0

3.0

6

0

20 ^0 6 0

A I r ve locit y in c m /s

Bot to m sphere,-

8 0 2 0 ^ 0 60

Air ve locit y m c m /s

80

Bottom sphere,-

1 DD= 2D

RotQt'ion

_ imparted to bottom sphere

. Bottom spher

0 0 r pm

X 1 2 0 r p m

Right spherc,-

1DD= 1-5D

Rl ght sphere, 1DD-- 1 D

^eft sphere. ID D= 1

• 5 D

Left sphere,! DD^ 10

20 ^0 60

Air VG !oc 1 1 y i n cm / s 80 0 20 ^0 60

Air velocity In cm /s

Fig. 1(a) Methanol consumption rates vs. air velocity for different flame

configurations

.

h_D u

0

ENVELOPE REGION

4

0 rpm::0

X r pm =60

A rpm z120

AIR FLOW

20 Rotation given to bottom

sphere only.

DiO

Flame

WAKE

RE G !0N

IDD RPM SYMB.

0 0

50 X 120

2 D 0 •

5 0 @

120 -o20 ^0

Air velocity in c m /s

6 0 80 0 20 40 60 80

Air ve loci t y in c m / s

Fig. 1(b). Nondimensional flame height vs. air velocity for methanol flames

548
of sphere diameter. Sphere rotation was imparted with a separate motor pulley

arrangement up to a speed of 120 rpm. A Pt-Pt+10%Rh thermocouple of 0.4 mm

diameter has been used to measure flame temperature. The bead was coated with

Si02 to minimize catalytic effect and usual conduction and radiation corrections

were applied. Gas samples have been drawn from the flame by means of a gas tight

syringe and introduced into a dual column chromatograph with a microprocessor

controlled integrator which recorded the flame composition. Molecular sieve 5A

and Porapak-N were used as the column material and helium was the carrier gas

.

The flame height has been measured from flame photographs taken with a Nikkon

camera with close-up arrangements.

RESULTS AND DISCUSSION

Burning rate

:

Figure 1(a) shows the measured burning rate plotted against

air velocity. The principal observations are: (a) There is an increase in fuel

consumption rate with increasing air velocity, (b) Drop rotation up to 120 rpm

does not appear to affect consumption rate significantly, (c) For vertical

arrangement (one drop placed above the other) in upward air flow, IDD=4D produces

a burning rate comparable with that of a single drop whereas IDD=2D produces a

burning rate less than that of a single drop, (d) For the same arrangement in

downward air flow, flame interference increases burning rate for IDD=4D. But for

IDD=2D, flame interference increases burning rate below an air velocity of 59

cm/s and decreases the rate above 59 cm/s . (e) For drops in tandem arrangement,

the left drop has a slightly higher burning rate than the right one. The reverse

trend is observed for downward airflow.

Flame height: Figure 1(b) shows non-dimensional flame height vs. air

velocity. It is observed that: (a) Flame height decreases with air velocity until

a critical point is reached where the envelope flame changes over to a wake

flame, (b) There is a slight reduction in flame height with rotational speed as

compared with a non-rotating drop, (c) For flame interference experiments, the

flame height seems to reach a peak at an air velocity of 20 cm/s.

Temperature profiles : Figures 2a and 2b show axial and radial temperature

profiles in upward airflow. The following observations are noted: (a) Compared

with a single drop, axial profiles are flatter for double drop studies, (b) Flame

interference produces slightly higher maximum axial temperature, (c) Drop in

temperature after reaching the maximum value is slower for double drop

experiments, (d) Increasing IDD produces flatter axial temperature profiles, (e)

The point of maximum axial temperature shifts downstream for interference

experiments, (f) The maximum radial temperature for a double drop is higher than

a single drop, (g) The rise in radial temperature is rapid in the case of a

single drop when compared with interference studies.

Figure 3 shows some typical temperature profiles for downward airflow. The

main observations are: (a) Comparing Figs. 3b and 3c, it is seen that the main

effect of drop rotation seems to make temperature profiles a little broader with

no appreciable difference in maximum temperature, (b) In general, radial

temperature profiles show peaks at the interface between the flame boundary and

ambient air. (c) Comparing Figs. 3a and 3d, the shape of a single drop profile

is different from that of a double drop, (d) Comparing Figs. 3c and 3f, the

effect of increasing air velocity seems to make the radial temperature profiles

somewhat steeper with a slight increase in maximum temperature.

549
Axial distance Axial distance Axial distance

Fig. 2(a). Axial temperature profile for methanol droplets in upward air flow.

Radial distance in mm Radial distance in mm Radial distance m mm

Fig. 2(b). Radial temperature profile for methanol droplets in upward air flow.

(Radial distance is measured from the central axis)

.

550
Radial distance in 'mm' Radial distance in "mm'

1200

r

1000

800

600

1200 -

1000

800

600

Radial at 2 D

1 200r At 1 D

1000

Air velo.'.Slcm

sphere rpm 800

^120

vVG k e flame

Air velo.'-SScm/s

sphere rpm-0

I D D AD

J i_ 1 U 6 8 10 12 U

Radial dis tance i n "mm'

600

U 8 12 16 20 U 28

Radial distance in'mm'

1 2 00

-At

midpoint

v,between

sphe res

Air velo:SScm/s

sphere rpm-.O^ ID D-- 2D

1 U 6 8 10 12

Radiai distance m'mm'

1000

T.K

800

600

A ir velo;.7 5cm/s

sphere rpm-0

I DD= 2D

At

midpoint

between

spheres

2 ^ 6 8 10 12 R ad lal d I s t once in'mm

Fig. 3. Typical radial temperature profiles for methanol in downward air

flow. Inset photo shows flame configuration.

551
Flame composition: The composition data for upward airflow are shown in

Table 1. It shows that the principal components are CO, and fuel

vapour. The fuel decomposes into CO and due to pyrolysis (Ref .7). CO finally

gets oxidized to CO2. Fuel vapour concentration is maximum near the drop surface

and progressively diminishes downstream. This trend does not appear to change

due to flame interference

.

Table 2 shows flame composition data for downward airflow. Again near the

drop surface, fuel concentration is high and decreases downstream progressively.

Again CO is the principal pyrolysed product which gets oxidized to CO2.

ACKNOWLEDGEMENT

One of the authors (AKG) gratefully acknowledges financial assistance

received from the Aeronautics Research and Development Board, Ministry of

Defence, for part of the work reported in this paper.

REFERENCES

1. Williams, A., Oxidation and Combustion Review, Vol. 3, pp. 1 (1968).

2. NatarajanR., Can. J. Chem. Engg. , Vol. 52, pp. 834 (1974).

3. Rex, J.F., Fuhs , A.E. and Penner, S.S., Jet Propulsion, Vol. 26, pp. 174

(1956) . 4. Nuruzzaman, A.S.M., Headley, A.B. and Beer, J., J. Inst. Fuel, Vol. 39,

pp. 248 (1966).

5. Brzustowski, T.A., Twardus , E.M., Woiceiki, S. and Sobiesiak, A., AIAA J

.

Vol. 17, No. 11, pp. 1234 (1974).

6. Gollahalli, S.R. and Brzustowski, T.A., Fourteenth S3rmp . (Intl.) on

Combustion, pp. 1333, The Combustion Institute, Pittsburg, PA (1973).

7. Gollahalli, S.R. and Brzustowski , T.A., Combustion and Flame , Vol. 22, pp.

313 (1974).

552
553
554
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 60

ATOMIZATION AND COMBUSTION OF LIQUID FUELS IN COUNTER-FLOW

B.N. Raghunandan, G. Jagannath and D.S. Mehra

Department of Aerospace Engineering

Indian Institute of Science

Bangalore, India

ABSTRACT

Counter-flow configurations have certain inherent advantages

over the normal co-flowing types like flame-holding capability

and compactness of the combustion chamber. The study aims to

generate some basic data on the flame-holding capability of

pressure as well as air assisted jets in the counter-flow mode.

It is shown that even in cold flow, the opposing air stream

exerts a considerable influence on the atomization from a

pressure jet. The trends of stability boundaries in both the

cases are similar to each other and to gaseous fuel jets. Within

the acceptable range of fuel penetration levels, a finer spray

provides better quality of combustion, but has a narrower

stability bounds than a coarse spray.

INTRODUCTION

There are many recent developments in burner configurations

both in aeroengine field and in industrial applications where the

injection of liquid fuel is not necessarily at the head-end as in

conventional burners. In such non-cof lowing configurations, an

obvious advantage is that the fuel finds a longer residence time

in the combustion chamber [1]. In addition to making the chamber

more compact, this mode is envisaged to accommodate a coarse

spray as would result in off -design conditions. Even from the

point of view of safety of igniter, the counter-flowing system

offers some advantages. The flame after ignition can be easily

manipulated to anchor away from the igniter which remains in a

cooler air stream [2].

Burners, in general, contain either a swirl flow or a bluff

body for the purpose of flame stabilization. The ability of tho

counter-flowing jets to act as aerodynamic flame holders has been

a subject of investigation since long [3,4]. The studies so far

have been restricted to gaseous fuels only whereas it is more

relevant to determine such blow-off characteristics for liquid

fuel sprays which are more commonly used in practical burners.

Even basic information like stability of jet flames of liquid

555
fuel in counter-flow configuration does not seem to be available.

Towards this objective of generating some basic data and

understanding, this study on the combustion stability of liquid

fuel jets- both pressure jets as well as air-assisted jets in a

counter-flow of air has been undertaken. Some of the interesting

observations on atomization in cold-flow conditions are expected

to find industrial applications.

EXPERIMENTAL

Experiments have been conducted in a combustion tunnel as

shown in Fig. 1. The air from a centrifugal blower is channelled

through a control valve, flow straightner and metering instrument

into a settling chamber. A nozzle of high contraction ratio

(18:1) and having an exit section of 200mm x 200 mm provides a

uniform outlet. While some of the initial experiments were

conducted in a test section attached to the nozzle-exit later

ones have been conducted without any confinement. However, a

large shroud had been used for the sake of safety when the flame

sizes were too big for a laboratory experiment. The water-head

measured at the venturimeter is calibrated against the velocity

in the test zone.

Fig.l Schematic of the experimental set up: (1. Blower

2. Control v^alve 3. Venturimeter 4. Settling chamber

5. Test Section 6. Atomizer 7. Pressure bottle

8. Pressure gauge 9. Orifice meter 10. Fuel Tank)

The fuel injection line consists of a tank, pressurising

system and finally the atomizer. While this suffices for a

pressure jet, an additional high pressure air attachment is

necessary (Fig. 1) for the studies with air-assisted atomizer.

The atomizing air flow rate is measured with a choked-orif ice

meter. The liquid fuel (kerosine) flow rate is obtained by

direct calibration of the flow against pressure drop. While

checking the flow rates in the presence of counter-flow, a slight

556
variation has been adopted namely one of discharging a known

amount of fuel over a measured time period. The details of the

atomizer characteristics are included in the following section.

RESULTS AND DISCUSSION

Pressure Atomizer

;

The pressure atomizers employed are all conventional swirl

chamber type with wide range of swirl parameter, A from 3 to 12.

In terms of the parameter Ap/dgdQ (see Ref 1), this corresponds

to a range of 0.08 to 0.35. Out of the several variations tried,

three representative ones are described below.

Table 1 Pressure Jet Characteristics

Parameter PJ-1 PJ-2 PJ-3

Orifice diameter (mm)^ 0.8 0.8 0.8

Swirl Parameter, A 3.0 6.0 12.0

Coefficient of Discharge 0.55 0.28 0.175

Cone Angle (Degrees) in 45-50 50-55 45-50

Quiescent surroundings

Penetration High Low Low

Spray Coarse Very good Good

Effect of counter-flow Unaffected Significant Significant

* A = Aodc/Atdo

** Nearly independent of Ap and counter-flow velocity (within 5%)

Table 1 describes some of the major characteristics of the

pressure jets PJ-1, PJ-2 and PJ-3. The case PJ-1 and similar

ones having high values tend to produce high penetrations both

due to coarse particle sizes and high axial component of the

velocity. Since such large penetrations annul the advantage of

compactness and also that in the present experimental set up pose

some hazard by way of fuel entering the settling chamber, thoy

are considered unsuitable. However, some tests were conducted to

determine if the counter-flowing air alters the spray

characteristics. But no favourable changes were observed in the

experimental range of velocity.

The injectors PJ-2 and PJ-3 with relatively low cone angle,

finer spray and low penetration have been tested for their

behaviour in the presence of counter-flowing air. With the air

velocity, though the C^ values are not altered much (within about

3%), the cone angle variation shows an interesting trend. As

seen in Fig. 2, for both types, the cone angle initially shows a

very weak increasing trend with air velocity followed by a sudden

increase in slope. In this expanded state, the atoraization level

A = AQdg/ApdQ where and Aq are diameter and area of the

orifice and dg the diameter and Ap inlet area of swirl chamber.

557
becomes much finer even for those cases which can otherwise be

termed as poor sprays. This trend can be seen for all the

values of pressure-drops studied.

u

M

E

>«o

o

Si

10

8

6

4

2

0

Injector PJ2

A on a

a

AODCa

A OD

-A - oQ-Ua

50 55

Aoa a

ODa

APi^j(p5i)

A 44

O 56

64

a 76

10

« 8

\

S 6

I 4

a>>5

2

0

Injector PJ3 aO 3

A o a

AOD a'

- AO a

AO a

Ao n dFig.

45 50 55 6 0 6 5 70

2 Injector characteristics with counter-flow (cold flow)

45 50 55 60

Cone angle (deg)

Further experimentation to understand

the transition process

lead to the observation

that there is a

transition region (see

Fig. 3) where the

variation of cone angle

is not really monotonic,

but involves some

oscillations. The cone

at the exit of the

orifice fluctuates

between the normal state

and the expanded state

aperiodically . Though

this can be attributed

to the influence of

counter-flow on the aircore formation and its

size in the vicinity of

the orifices, it

warrants further work to

12

^ 8

u

-2 4

>

O Injector PJ 2

A Injector PJ3

Expanded ^^^""^^

state ^..^J^-'-''^!^^'^

'A —1 -A

Normal state

Transition zone

1 1 20 25 30

Fuel injection velocity (m /sec)

35

Fig. 3 Transition in cold flow

ascertain the exact mechanism.

The blow-off characteristics of the flame generated by

counter-flowing system is shown in Fig. 4. Extinction of the

flame occurs for both PJ2 and PJ3 at air velocities slightly

higher than that causing complete transition described above.

Interestingly, at a given fuel injection velocity, the counterflow air velocity for blow-off is in the same range as for

gaseous fuels (see Ref 2) though the actual fuel mass flow rates

in the present case are very much higher. Like with the gaseous

fuels, even in the present set of experiments, the onset of

558
extinction is marked by the formation of a dark zone along the

axis of the injector.

Air-assisted Atomizer

;

At the outset, it

may be noted that this

type of atomizers

provide cooling to the

injector inside the

enveloping flame and at

the same time, provide

air for combustion in

the vicinity of the jet

orifice. The atomizer

consists of a fuel

nozzle concentric with

which is an annular

passage for atomizing

air [see Ref 5 for

details]. The atomizer

has the provision of

moving the fuel orifice

relative to the

atomizing air outlet in

such a way that the same

unit can be made to act

either as an internal

mixing type or as an

external mixing type.

Three of the

representative atomizer

configurations are

described in Table 2.

Like in pressure jets,

here also narrower cone

angles show greater

ability of flame hoi

for the experiments,

experimental injector yi

position, but in a rather

retracted deeper into the

10

8

6

4

1 2

0

Extinction

boundary

Cold flow transition

Injector P J 2

o 20 25 30

Fuel injection velocity (m /sec ) 35

>

|I2

1 n jection PJ3

•

^

Cold-fiowtransition

Extinction

boundary

j 1 20 25 3 0 35

Fuel injection velocity (m/sec)

Fig. 4 Combustion stability characteristics (Pressure

Atomizer )

.

ding and hence are adopted

Cold flow tests have shown that the

elds fine spray in external mixing

coarse spray when the fuel orifice is

air nozzle. As is to be expected, the

Table 2 Some Characteristics of Air-assisted Jets

Parameter AJ-1 AJ-2 AJ-3

Fuel Orifice(mm) 0.4 0.4 0.4

Air Nozzle dia (mm) 3.0 3.0 4.5

Configuration (relative In-plane Internal External

position of the orifices) ext .mixing mixing mixing

Spray Fine Coarse Finer than AJ-1

Cone angle (deg) 12-19 7-10 12-14

559
fineness of the spray improves as the ratio of atomiser air , flow

(m^ ) to fuel flow (mf ) is increased.

Unlike the case of pressure jets, here the counter-flowing

air does not seem to alter the spray in the vicinity of the

injector, i.e., the cone angle does not get affected. This is

perhaps due to the fact that the atomizing air velocity has to be

necessarily very large ( " 50 m/sec) and also there is no lowpressure air core in this type of atomizer.

Combustion of this type of spray yields a flame which has a

hollow conical region which is blue in colour and at the far end

a mush-room like flange. The conical region is due to combustion

of fuel in the presence of atomizing air. With increasing

counter-flow velocity, the flange bends towards the injector

before being blown off. Increasing atomizing air flow tends to

shift the flame away from the injector and at a given counterflow velocity, the flame can be blown-out with atomizing air

alone although the atomizer air-fuel ratios are well within the

values used in such air-assisted atomizers [6].

In this configuration,

the atomizing air flow is an

additional parameter in the

stability diagram. The data

of blow-off as obtained in

the experiment are shown in

Fig. 5 for two different

configurations AJ-1 and AJ3. For the case of AJ-2

also one gets similar plots

with the curves shifted

towards higher counter-flow

air velocity, but the flame

tends to be sooty due to

coarseness of the drops.

The nature of curves hint at

a possible explanation for

extinction in terms of the

effective equivalence ratio

at the critical zone around

the stagnation point much as

the one invoked by Schaffer

and Cambell [4] in their

work on flame holding. But

a precise description in

terms of accurate numbers

has so far proved elusive.

Shown in Fig. 6 is the

stability plot for AJ-1

redrawn from the data in the

073 05 07

^

Atomising air flow rate(rhQ) 10' kg/s

Atomising airflow rate (mo) IO"^kg/s

Fig. 5 Combustion stability

plots for air-assisted

jet in counter-flow.

560
previous figure. On this plane with (mf/m^) as a variable the

different curves merge into a narrow band. Identical behaviour

is observed with the fine spray of AJ-3, but for AJ-2 substantial

dispersion exists. Noting that such a merging is obtained for a

range of both mf and m^ , one can view the abscissa as the fuel

flow variation at a constant atomizer air flow rate. From this

point of view the trend in Fig. 6 is comparable with that of

pressure atomizer in Fig. 4. This argument suggests that the

mechanism of extinction may be common to both the cases discussed

here and to the gaseous counter flow system described in Ref. 2.

12

8

u

«n \

J

o

o

a> > 4

0

Injector AJ-I

O A

•

• •

O

o CD

A

gm/sec • 0.71

O 0.60

A 0.55

0.50

0.2 0.6 0.8 1.0 1.2 1.4

Fig. 6 Flame-holding ability of air-assisted jet

CONCLUDING REMARKS

The atomization from a pressure jet is considerably

influenced by the presence of a counter-flowing air stream. An

air-assisted fuel jet however is quite resistant to the counterflow. Fuel sprays of either type can act as flame-holders over a

range of air velocities without any swirl motion. One can expect

the stability to improve if recirculation is also imposed.

Coarse sprays tend to have higher penetration and accommodate

higher counter-flow velocity than the fine-sprays. The

combustion stability characteristics of both the types of jets

are similar and seem to share some features with the gaseous fuel

jets. Unlike the co-flowing streams where the flame anchoring is

governed by the processes near the fuel outlet, in these flames

the region around the stagnation point is critical . This is

despite the fact that the sprays fan out leaving low fuel

concentrations along the axis.

561
ACKNOWLEDGEMENTS

The authors thank Mr. S. Nagaraj for his ready services in

photography.

REFERENCES

1. Lefebvre, A.H., Gas Turbine Combustion

,

Hemisphere,

Washington (1983).

2. Raghunandan, B.N., Gopakumar, B.R. and Jagannath, G. , "Arc

Ignition of Gaseous Fuel Jets in a Counter-flow of Air", Jl

.

Inst. Energy, Vol. 60, pp 132-137 (1987).

3. Schaffer, A. and Cambel, A.B., "The Effect of an Opposing

Jet on Flame Stability", Jet Prop., Vol. 25, pp 284-287

(1955)

.

4. Schaffer, A. and Cambel, A.B., "Continued Investigations of

the Opposing Jet Flame holder". Jet Prop., Vol. 26, pp 576-

578 (1956).

5. Jagannath, G. , "Combustion Stability Studies of Plain-jet

Atomizer in a Counter-flowing Air Stream", ME Thesis, KREC

Surathkal (1988).

6. Rizk, N.K. and Lefebvre, A.H., " Influence of Atomizer

Design Features on Mean Drop Size", AIAA Jl . , Vol. 21, pp

1139-1142 (1983).

362
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 61

SIMILARITY BETWEEN THE BREAK-UP LENGTHS

OF A HIGH SPEED LIQUID JET

IN ATMOSPHERIC AND PRESSURIZED CONDITIONS

M. Arai*, M. Shimizu^ and H. Hiroyasu*

* Dept. of Mechanical Systems Engineering, Gunma University, Kiryu, Gunma, Japan

^ Dept. of Mechanical Engineering, Kinki University, Osaka, Japan

* Dept. of Mechanical Engineering, University of Hiroshima, Hiroshima, Japan

ABSTRACT

It is well known that the break-up length of a high speed jet such as a diesel spray is greatly affected by a

structure of an injection nozzle. In this paper, to reveal the relationship between the break-up length and the

structure, an internal flow in a transparent nozzle of 3 mm diameter was observed. The breakup lengths of jets

injected through nozzles of diameters 0.3 mm and 3 mm, were measured by an electric resistance method. They

were compared and suitable nozzle structures were selected to keep phenomenological analogies of break-up lengths

in nozzles of different scales and in different surrounding conditions. Similarities of the break-up lengths in

atmospheric and pressurized conditions were discussed about an internal flow effect on the break-up mechanism. A

reattachment point of a separated flow in a nozzle was investigated as a characteristic of the internal flow. The

behavior of break-up length in the spray flow region could be explained by a movement of this point.

INTRODUCTION

The main causes for the disintegration of a high speed liquid jet have been suggested by vibration of an

injection pressure velocity distribution in a nozzlc'^^, injection velocity'^\ internal and shear forces between

a jet and gaseous environment"^'^ ^. However, disintcgraiion mechanisms just after a nozzle exit are not well

understood, although several studies^'^'^^ have attcmpied to explain them. Liquid turbulence, cavitation

and aerodynamic interaction between the jet and surroundings iire thought to be important factors^\

The disintegration of the jet is also affected by physical properties of the injected liquid, structure of the

o nozzle, injection velocity and ambient gas pressure or density. It have been already suggested in previous studies

that there is a strong dependence between the break-up length and an interfacial force due to the gaseous

environment. Therefore, as a further study of the bre^ik-up length, the effect of the internal flow in a nozzle on the break-up length was investigated and discussed in this study. The break-up mechanism of the liquid jet

injected in a pressurized surroundings was discussed by a phcnimcnological analogy of the break-up lengths of the

jets in atmospheric and pressurized surroundings.

EXPERIMENTAL APPARATUS AND METHOD

The apparatus shown in Fig. 1 was constructed to study the break-up phenomena of a high speed jet in a high pressure surroundings. The apparatus consisted of a spray chamber, a liquid injection system, an

injection nozzle and a wire net detector for measuring a break-up length.

Water was continuously injected by the injection system into the spray chamber. The injection velocity of

a jet was calculated as the sectional average velocity in the nozzle. The Reynolds number Re of the jet was

563
calculated with kinematic viscosity of water, nozzle diameter and this velocity. An electrical circuit was

used to determine the break-up length. Definitions of the break-up length and spray angle were reported with

O Q\

measurement methods in the previous papers ' ^ A fundamental structure of a nozzle used in this investigation is also shown in the figure. It had a

long inlet pipe to obtain a smooth inlet flow to the nozzle. The entrance of the nozzle was made in a sharp

edge, so that a consU'icted flow was easily generated at the entrance. Hole diameter D, length L and inlet pipe

diameter D„ specified a structure of the nozzle. Lengih-io-diameter ratio L/D and sectional area ratio of the hole

diameter to the inlet pipe diameter, M = (D/D^r were the non-dimensional parameters used to represent the

structure.

Accumulater Pressure Gauge

Chamber

Do

[7

-D Nozzle

QOinA

Fig. 1 Experimental apparatus and nozzle

BREAK-UP LENGTHS IN VARIOUS LIQUID JETS

The disintegrating behavior and break-up length of a high speed liquid jet are frequently a technical matter

of great importance for diesel injection nozzle performance. Therefore, a hole diameter 0.3 mm of a nozzle and

an ambient pressure 3.0 MPa into which a jet was injected were selected to reproduce a spray in a combustion

chamber of a diesel engine.

Effects of the length-to-diameter ratio L/D on the break-up length are summarized in Fig. 2. Since the ambient

pressure was high enough to produce a strong intcrfacial force between the jet and surroundings, fine atomization

could be observed when an injection velocity was higher than 70 m/s and this was termed spray. The breakup length in this region decreased with an increase in the injection velocity. For the cases of L/D = 4 and 10, it passed through a minimum, and then reached an almost constant value. When the L/D ratio was 1 or far larger

than 10, there was no minimum point of the break-up length in the spray region.

When the ambient pressure was reduced from 3.0 MPa to 0.1 MPa, the break-up length was extended as shown in Fig. 3. Further, a kind of hysteresis appeared in a low velocity region. If the velocity increased

from lower value than where two possible break-up lengths existed, the actual break-up length was the shorter of

the two. However, if the velocity decreased from the higher value, it was the longer break-up length produced by

the separated flow issuing from the nozzle This phenomena was more evident in the break-up length

produced by 3.0 mm diameter nozzle, that is also shown in the figure. From the behavior described above, it

is concluded that no analogy in the break-up length could be maintained between the 0.3 mm and 3.0 mm nozzles,

using in the different pressures.

The sectional area ratio M = (D/D^) could sometimes affect the break-up length. Figure 4 shows this effect on the break-up lengths of jets injected by 3.0 mm nozzles into atmospheric environment. The hysteresis

mentioned above was only observed with the nozzle which had 0.0036 as the value of M. The break-up lengths

in the spray region that is shown in higher velocity than 10 m/s, were different corresponding the values of M.

564
3000r

10 10 Injection Velocity Vi m/s

0.15 1.0 10 10^ Injection Velocity Vi m/s

10=

Fig. 2 Effect of L/D on the break-up length

in 3.0 MPa surroundings

Fig. 3 Effect of the hole diameter on the break-up length

2.5x

1,0 10

Injection Velocity Vi m/s

Fig. 4 Effect of sectional area ratio of the inlet pipe on the break-up length

ANALOGY OF THE BREAK-UP LENGTHS IN DIFFERENT PRESSURE SURROUNDINGS

Figure 5 illustrates the two different forms of the break-up behavior recognized in Fig. 2 as a function of the

injection velocity. The spray region was divided into two different sub-regions, such as the pseudo spray and the

complete spray which are shown by E, G and F. In this study, a new analogical method was introduced to explain

the break-up behaviors in the E - F and E - G - F processes, using an internal flow structure in a nozzle

injecting into atmospheric surroundings.

As mentioned before, the break-up length was greatly affected by D, L/D and M of a nozzle as well as the

velocity and the pressure of the surroundings. Thus, the L/D and M of a transparent nozzle of 3.0 mm

diameter at the atmospheric condition, could be selected to produce similar behavior in break-up length that

produced by a 0.3 mm diameter nozzle in a pressure of 3.0 MPa.

Figure 6 shows the comparison of the non-dimensional break-up length Lj^/D with Reynolds number under the

conditions of high and atmospheric pressures. The horizontal axis indicates the Reynolds number because the

injection velocities and the nozzle diameters were different. In general, when the nozzles of different diameter were used in different pressure surroundings, the behavior of the break-up lengths were different owing to

565
different break-up mechanisms. However if a suitable value of the sectional area ratio M and the length-to-hole

diameter ratio L/D of a nozzle were selected. The same behavior of the break-up length with Reynolds number

could be approximately obtained as shown in this figure.

In the case of the nozzle in which a minimum break-up length was produced within the pseudo spray region

in 3.0 MPa surroundings, 10 times larger nozzle that had different M but had the same break-up behavior in

atmospheric surroundings could be found out as shown in Fig. 7. The Reynolds numbers were different

where the non-dimensional break-up length L^/D indicated the minimum in the pseudo spray region.

However, the trends of the break-up lengths were similar to each other if the difference in the Reynolds numbers

was discounted.

The reason why this analogy was attained could be explained by discharge coefficients shown in these

figures. The decreasing tendency of the discharge coefficient meant an increasing tendency of a drag force

that produced turbulence in the nozzle. The rapid decreases appearing in Fig. 6 at Re = 4 x 10^ to 10^

produced the pseudo sprays in both nozzles, so that the analogy of the break-up length was obtained by the

analogy of the internal flow in the nozzles. Tabic 1 shows the summarized specifications of the nozzles used in

Figs. 6 and 7. The break-up behavior of the jet injected by Nozzle-c into atmospheric surroundings was

similar to that of Nozzle-a in 3.0 MPa surroundings, and the break-up length of Nozzle-d consists with that

of Nozzle-b. Further investigation on the relationship between internal flow and break-up length was

therefore carried out under atmospheric conditions using transparent nozzles specified by Nozzle-c and -d.

Injection Velocity Vi

Fig. 5 General behavior of the break-up length of a jet in a high pressure surroundings

Reynolds Number Re Reynolds Number Re

Fig. 6 Analogy of the break-up lengths and discharge Fig. 7 Analogy of the break-up lengths and discharge

coefficients between Nozzle-a and Nozzle-c coefficients between Nozzle-b and Nozzle-d

566
Table 1 Specifications of nozzles

K

1

^\ m ml A

iNozzie

D

(mm)

L

(mm)

1 /n L/U

Do

(mm) Comments

a 0.3 15.0 50 3.0 0.01 E-F in 3.0 MPa.

b 0.3 1.2 3.0 0.01

E-G-F in 3.0 MPa,

Analogous to a diesel nozzle.

c 3.0 60.0 20 50.0 0.0036

E-F in 0.1 MPa,

Analogous to nozzle -a.

d 3.0 12.0 40 0.56

E-G-F in 0.1 MPa,

Analogous to nozzle -b.

CORRELATION OF BREAK-UP LENGTH AND SPRAY ANGLE WITH AN INTERNAL

FLOW IN A NOZZLE

Photographs of the internal flows of Nozzle-c and the jets under conditions specified in Fig. 6 are shown in

Fig. 8. As the injection velocity increased from the pseudo spray at 13.5 m/s ( Re = 5x10^ ) to the complete

spray at 27.0 m/s ( Re = 10^ ), a constriction flow and a cavitation with bubbles were simultaneously

produced at the entrance of the nozzle. Also a separated flow made by this constriction reattached the nozzle

wall. With an increase in the velocity, the reattachment point of the separated flow moved toward the nozzle exit

While velocity remained above 21.4 m/s ( Re = 8x10^ ), the reattachment point moved up-and-down at the

vicinity of the nozzle exit. With the internal flow phenomena described above, the disintegration of the issuing

jet was promoted and the break-up length and the discharge coefficient smoothly decreased by increasing the

Reynolds number as shown in Fig. 6. Photographs of the internal flows in Nozzle-d and the jets in atmospheric surroundings are shown in Fig. 9. With the velocity increase from 12.0 m/s ( Re = 45,000 ), a constricted flow and a cavitation with bubbles were

produced as in the previous photographs. However, when the velocity increased above 16.6 m/s ( Re =

61,000 ), the reattachment point moved upstream. The break-up length decreased and the spray angle increased

with a downward movement of this point. Also the 16.6 m/s velocity that produced the minimum break-up

length in the pseudo spray region consisted with the velocity where the point was nearest to the exit

M=00036 Pa=01MF& Nozzle-d L/D = A D=3.0mm M=0,56 Pa =0.1 MPa

Fig. 8 Internal flows and jets produced by Nozzle-c Fig. 9 Internal flows and jets produced by Nozzlc-d

567
The reattachment point was measured from the photographs and expressed by a distance 1 from the entrance of

the nozzle. Figure 10 shows the correlation of the break-up length and the spray angle 9 of the jets produced

by Nozzle-c, connecting with the reattachment distance 1. This result corresponded to the break-up length

smoothly changing from E to F, as shown in Fig. 5. Further it was considered to be a simulation of the

internal flow in Nozzle-a, as used in 3.0 MPa surroundings according the analogy of the internal flows in the

nozzles. In the pseudo spray region, as the injection velocity increased from 12.5 m/s, the reattachment distance 1 increased and moved toward the nozzle exit. In the complete spray region of the injection velocity above 19 m/s,

this point moved up-and-down near the nozzle exit. Accompanying these phenomena in a nozzle, the break-up

length decreased and the spray angle 0 reached a maximum at around 22 m/s.

The same trend in results concerning No/,zlc-d arc shown in Fig. 11. In this case, the behavior of the

break-up length shown by the E - G - F curve in Fig. 5 was presented by the analogy between Nozzle-b and -

d. With an increase in the velocity, the reattachment distance 1 increased and approached the exit at 18.0 m/s

in the pseudo spray region. Furthermore, the injection velocity increased, the recovery of the constricted flow in

the nozzle was promoted and the reattachment distance was decreased. However, the reattachment point moved

up-and-down at the location a little upstream from the exit. Accompanying the internal phenomena at the

injection velocity of 18.0 m/s, the spray angle 0 reached the maximum and the break-up length became the

minimum.

50

CD

i, 30 c<

o 20 a

i/i

10

450 500

L 120

10'

10 15 Injection

20 25 30

Velocity Vi m/s

5«10' 10= 15x10*

10 15 Injection20 25

Velocity Vi

10' Reynolds Number Re5x10' 10* Reynolds Number Re

15x10*

Fig. 10 Break-up length, spray angle and reattachment Fig.

distance obtamed by Nozzle-c as functions of

the injection velocity

11 Break-up length, spray angle and reattachment

distance obtained by Nozzle-d as functions of

the injection velocity

Figure 12 shows the results in which Fig. 10 was rewritten by using the non-dimensional reattachment

distance 1/L. As the 1/L increased by increasing the injection velocity, the break-up length decreased. The spray

angle reached the maximum value at around 1/L = 0.75. TheLjy^and 0 had tendencies to be reduced by the

effects of an up-and-down vibration of the reattachment point ( its range is shown by arrows in the figure ). The relation between the 1/L and the break-up behavior when the minimum break-up length appeared in the

pseudo spray region is shown in Fig. 13. The results in the pseudo spray are represented by open symbols

and solid symbols show the complete spray. As 1/L increased, Lj^/D decreased and 9 increased in the pseudo

spray region. Further increase of 1/L, the spray angle reached the maximum and LjyD became the minimum at

1/L = 0.95. In the complete spray region, 1/L was decreased by an increase of the velocity. It remained in the

range of 0.5 to 0.65 but was unstable. The increased and the spray angle decreased by an unstable effect

of an up-and-down vibration on the reattachment point. It could be considered that this form of the

disintegration might appeared in the increasing behavior of the break-up length shown by G - F curve in Fig. 5.

568
Dimenslonless Reattachment Distance l/L 0 02 0.4. 0.6 0.8 1.0 Dim en si onless Reattachment Distance I/l

Fig. 12 Break-up length and spray angle as functions Fig. 13 Break-up length and spray angle as functions

of reattachment distance ( Nozzle-c ) of reattachment distance ( Nozzle-d

)

CONCLUSIONS

In order to explain the disintegrating mechanisms of a high speed liquid jet under high pressure environment, the

correlation of the internal flow in a nozzle and the break-up length was investigated and following results were

obtained.

(1) A spray which shows the analogical behavior in break-up length of a jet in high pressure surroundings

could be reproduced by selecting the suitable sectional area ratio and length-to-hole diameter of a nozzle under an

atmospheric pressure condition.

(2) Concerning the observations of the liquid flows in the nozzle, if the separated flow and cavitation

occurred, the jet could be broken up to the spray.

(3) The closer a reattachment point of the internal flow in a nozzle approached the nozzle exit, the shorter

the break-up length and the larger the spray angle become.

(4) The reattachment point of the liquid flow oscillated in the vicinity of the nozzle exit when the break-up length

reached the minimum in the pseudo spray region.

(5) From the results described above, it is concluded that if a large disturbance contains in an issuing jet, it may

be atomized and broken-up to a spray very quickly

REFERENCES

1) Bregwere,W., "Flow pattern in Diesel Nozzle Spray Holes", Proc. Inst. Mech, Engirs., 173 (1959), 655.

2) Rupe.S.H., "On the dynamic characteristics of fuel liquid jet and a partial correlation with orifice geometry",

JPL Tech. Kept., 32 (1962), 207.

3) Lee, D.W., " the effect of nozzle design and operating conditions on the atomization and distribution of fuel

sprays", NACA Tech. Kept., No.425 (1932), 565.

4) Tanasawa, Y., High speed diesel engines I, "Fuel systems and spray combustion", Sankaido ( in Japanese ), (1956), 87.

5) Ruiz, R and Chigier, N. " Mechanism of high speed atomization", Proc. 3rd ICLASS, Vol. 1 (1985),

VlB/3/1.

6) Schweizwer.P.H., "Mechanism of disintegration of liquid jet", J. Appl. Phys. Vol.8, (1938).513.

569
6) Schweizwer,P.H., "Mechanism of disintegration of liquid jet", J. Appl. Phys. Vol.8, (1938).513.

7) Reitz, R.D. and Bracco.F.V. "Mechanism of atomization of a liquid jet", Phys Ruids Vol.25 No.lO

(1980).1730.

8) Hiroyasu.H., Shimizu, M. and Arai, M.," The break-up of high speed jet in a high pressure gaseous

atmosphere" Proc. 2nd ICLASS (1982), 69.

9) Arai.M., Tabata,M.,Hiroyasu,H. and Shimizu, M., "Disintegrating process and spray characterization of

fuel jet injected by a diesel nozzle", SAE Paper No. 840275 (1984).

10) Arai.M. Shimizu,M. and Hiroyasu.H. " Break-up length and spray angle of a high speed jet" Proc. 3rd

ICLASS, Vol.1 (1985), IB/4/1.

11) Simizu.M. Arai,M. and Hiroyasu.H. "Effects of ambient pressure on break-up length in high speed jets",

14th Conf. on Liquid Atomization and Spray systems in Japan (1986), 1.

12) Arai.M., Shimizu,M. and Hiroyasu,H. "Break-up length and spray formation mechanisms of a high speed

liquid jet" Proc. 4th ICLASS, (1988), 177.

570
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 62

EFFECT OF NOZZLE CONFIGURATION ON THE ATOMIZATION

OF STEADY SPRAY

T. Karasawa*, M. Tanaka', K. Abe^

S. Shiga' and T. Kurabayashi*

*Dept. of Mechanical Engineering, Gunma University, Kiryu, Gunma, Japan

tpuji Heavy Industry, Ohta, Gunma, Japan

^Hiroshima Institute of Technology, Hiroshima, Japan

ABSTRACT

The purpose of this study is to obtain the relationship between spray

size and the nozzle hole configuration represented by L/D and the shape of

flow inlet in steady spray under the condition of diesel injection pressure.

The spray size is apparently affected by the entrance shape. However, it is

not substantial, because the effect disappears by correcting the injection

velocity using the nominal value which can be calculated on the basis of the

discharge coefficient of the nozzle. The droplet size increases with

increasing the value of L/D. This behaviour is considered to be caused by the

recovery of the separated flow at the inlet edge, which is suggested by the

observation of photographs. The nozzle diameter also has much influence on

the spray size, where a characteristic affecting behaviour is recognized.

1. INTRODUCTION

Even the spray formation without combustion has not fully understood

because of the complexity of the phenomenon. Spray formation behaviour of an

intermittent spray has been being studied by several researchers including the

authors.^ ''"^ -^ However, less studies of steady spray can be found out,

although it is fairly important to reveal the very basic feature of spray. In

this field, studies of Tabata et al^ and Shimizu et al^ ''"^ should be

cited. Tabata measured the spray size in rather wide range of the injection

pressure, and Shimizu proposed a concept of a new parameter of break-up length

to characterize the steady spray. However, there is little study aiming at

systematically clarifying the relationship between the nozzle configuration

and the spray size. In terms of the break-up length, Shimizu et al^ showed

that the best atomization can be realized when the nozzle depth to diameter

ratio (L/D) is around 5. It also refers to the recovery of the separated flow

at the edge of nozzle inlet as the mechanism of L/D effect, although it has

not been verified clearly yet.

On the basis of this background, an experimental study was carried out to

reveal how much we have the responsibility of nozzle configuration for the

droplet size of steady spray. Several interesting aspects of the

configuration effect could be obtained, and a possible mechanism to explain

those effects could be proposed on the basis of several verification

experiments

.

571
2. EXPERIMENTAL APPARATUS AND PROCEDURE

A schematic of experimental apparatus is shown in Fig. 1. Water was used

as a tested liquid. It is contained in a flexible bag(No.4).

Most data were obtained by a phase doppler system (Aerometrics Co.,

PDPA) and were checked by the measurement of a sampling technique, immersion

sampling method having a mechanical shutter.

As the parameters to represent the nozzle configuration, two sets of

variables were chosen, which are seemed to have major effect on the

atomization; nozzle depth to diameter ratio (L/D) and the shape of flow inlet.

Three kinds of the shape of the nozzle inlet were applied form a viewpoint of

flow separation, straight type (S-type), round type (R-type) and C-type, which

are shown in Fig. 2. The naming procedure is shown in Table I. The liquid was injected downward into an atmospheric condition. The

sampling point for the immersion method is 1500 mm downstream of the nozzle

along the spray axis which was determined by a preliminary optimization test.

The measuring points for the optical technique were 1000 mm or 1500 mm

downstream of the nozzle along the spray axis which were also determined by an

optimization test.

3. RESULTS AND DISCUSSION

3-1. Effect of Nozzle Depth to Diameter Ratio (L/D) on Atomization

In Fig. 3, the droplet diameter variation of S-type nozzle with L/D is

shown at three mean injection velocities. The droplet size increases with

increasing the value of L/D in every injection velocity. In Fig. 4 of R-type

nozzle, it can be seen at least that the variation of the size with L/D is not

so large comparing with that of Fig. 3, and that there is no clear increasing

tendency of SMD. Thus, the characteristic feature of S-type nozzle is related

to the difference of the configuration of the flow inlet.

The inlet flow of S-type will be easily separated at the edge, while it

should be much harder to occur for R-type. Then, the flow separation can be

assumed to have some role to realize the characteristics of S-type nozzle,

which was suggested in the previous study^ on the break-up length

measurement. Because of the separation, the effective area for flow coming is

reduced, and the flow is free from the wall effect. Then, the velocity can

reach the nearly ideal velocity of potential flow, which leads to the

atomization improvement. When the L/D is greater than that particular value,

the separation will be recovered and the flow will be again attached to the

wall. The velocity will be decreased by the wall effect, and then the

atomization effect of the flow separation tends to disappear. In Fig. 5, the

droplet size of S-type at 1000 mm downstream of the nozzle is shown, where the

exactly similar tendency to the 1500 mm measurement can be recognized.

3-2. Verification of the Separation Effect

In order to verify the separation effect suggested above, discharge

coefficients for several conditions were measured and shown in Fig. 6. The

solid line without a point is obtained by the calculation based on the

Blasius' empirical formula of turbulent pipe flow^ The variation of the

discharge coefficient of R-type nozzle with L/D is quite similar to the

developed turbulent flow. Then, the flow is considered to be governed by the

turbulent boundary layer. In contrast to this, the coefficient of S-type

nozzle does not vary at all over the whole range of L/D. This is an

indication of the occurrence of separation, and is not the indication of the

separation recovery, because the discharge coefficient would have an increase

with increasing L/D from a certain value of L/D if the recovery could have

occurred. In any way, the recovery should occur in this case, because even

572
the separation of back step can be recovered at the downstream distance of

around 7 times as much as the step height. This may be caused by the velocity

distribution within the hole, i.e., the velocity distribution of the

separation point will be conserved to some extent for a while^ In

addition, the fully developed turbulent boundary layer of pipe flow is said to

be realized at 25 to 40 of L/D (originally called as "inlet length"), which

coincides with our results^

Further indication of the recovery occurrence is shown in the

photographic observation. In less than L/D of 4, the spray surface of S-type

looks quite smooth, while the other conditions including the spray of R-type

do not show this distinguished feature.

3-3. Factors Determining the Atomization

The droplet size dependence on the mean injection velocity is shown in

Fig. 7 for every kind of inlet shape. The size decreases exponentially with

the increase of the mean injection velocity, which may be qualitatively

consistent with the result of a previous study^"*"-^'^ \ It is also clear that

the size is smallest for S-type nozzle, largest for R-type and intermediate

for C-type. Then, it can be seen apparently that the S-type nozzle has the

best configuration in terms of the atomization. However, we need much more

injection pressure for S-type nozzle than for R-type to get the same mean

injection velocity. It is due to the difference of the discharge coefficient.

Figure 8 shows the droplet size dependence on the corrected injection

velocity which is obtained by the following relationship.

= V„/C,.

Here, : corrected injection velocity

V„ : mean iniection velocity m -' '

: discharge coefficient.

This correction implicitly assumes that the discharge coefficient is mainly

determined by the flow separation, i.e., the decrease of the effective

sectional flow area. The variation of droplet size with the nozzle

configuration almost disappears, and the dependence of every kind of nozzle

can be approximately represented by a unique line. Thus, when the corrected

injection velocity is used, that may be the nominal injection velocity at the

nozzle exit, the droplet size is not affected by the nozzle configuration. In

the other words, atomization is not determined by the nozzle configuration but

by the actual injection velocity.

Further verification of the above mechanism can be made more accurately

as shown in Figs. 9 and 10. Usually, rather precise mechanical processing is

necessary to get the reproducible value of discharge coefficient. Then, most

nozzles used in this study is made by an electric discharge method. When a

simple processing like drilling is used, pretty much variation of the

discharge coefficient is not avoidable. This variation can be seen in Fig. 9.

Four nozzles are used numbered from 1 to 4 and made by a simple processing.

Although the droplet size dependence of a particular nozzle on the mean

injection velocity is precisely expressed by an exponential curve, the

dependence varies much in every nozzle. When the corrected injection velocity

is used by the discharge coefficient, that dependence can be exactly expressed

by a unique line as shown in Fig. 10. Thus, the previous mechanism of the

importance of the nominal injection velocity is well supported by this result.

573
4. CONCLUSIONS

(1) The droplet size increases with the increase of L/D for straight type

nozzle, which is caused by the flow separation at the inlet edge and the

separation recovery downstream of the separation point.

(2) Above mechanism can be verified by a comparison study between nozzles of

S-type, C-type and R-type, measurement of discharge coefficient and

photographic observation.

(3) Atomization is determined by the flow velocity at the nozzle exit, which

can be shown by the correction of the droplet size dependence on the injection

velocity using the discharge coefficient.

REFERENCES

(1) Shiga, S. et al, Proc. ICLASS-88, (1988), 397.

(2) Takahashi, H. et al. Trans. JSME ser. B, 56-526, (1990), 1837, (in

Japanese)

.

(3) Varde, K. S. and Popa, D. M., SAE Paper, 83448(1983).

(4) Hiroyasu, H. and Aral, M., Trans. JSME, No. 21, (1980), 5, (in Japanese).

(5) Tabata, M., et al, Proc. ICLASS-85, (1985), II/l/l.

(6) Shimizu, et al. Trans. JSME ser. B, 49-448, (1983), 2886, (in Japanese).

(7) Shimizu, et al , Trans. JSME ser. B, 51-461, (1985), 257, (in Japanese).

(8) Shimizu, et al. Trans. JSME ser. B, 54-504, (1988), 2236, (in Japanese).

(9) Schlichting, H., Boundary-Layer Theory 7th ed., McGraw-Hill Co., (1979),

597.

(10) Private communication with Dr. N. Fujisawa, associate professor of fluid

dynamics laboratory in Gunma University.

(11) Schlichting, H., Boundary-Layer Theory, 7th ed., (1979), Chapter XX,

McGRAW-HILL BOOK Co.

ACKNOWLEDGEMENT

The particle sizing system of PDPA is supplied by CENTER FOR COOPERATIVE

RESEARCH GUNMA UNIVERSITY. Apparatus for experiment is partly supplied by

ZEXEL Co. Ltd. The authors acknowledge Mr. Hisao NAKAMURA, Yasuo NAKAJIMA and

Koichi HASHIMOTO for their help with the experiment.

574
Fig. 1. Schematic of experimental apparatus.

60'

(a) S-type (b) C-type (c) R-type

Fig. 2. Configuration of nozzle.

575
200

1 2 3^5 10 20 Dimensionless Length of Nozzle

50 L/D

200

150

100

J uj Water

PDPA l=lOOOmm

Vm=90

1 ^ o o

o

m/s i

E

Q

to

1 2 3 4 5 10 20 50

Dimensionless Lengtin of Nozzle L/D

Fig. 3. Variation of droplet size with Fig. 5. Variation of droplet size with

the dimensionless length of nozzle(L/D), the dimensionless length of nozzle(L/D),

S-type nozzle, measuring point; S-type nozzle, measuring point;

1500 mm downstream of the nozzle. 1000 mm downstream of the nozzle.

250

200 R03-03-

PDPA

Water

l=1500mm

Vm=90m/s

1 2 3 4 5 10 20

Dimensionless Length of Nozzle

50

L/D

Fig. 4. Variation of droplet size with

the dimensionless length of nozzle(L/D),

R-type nozzle.

576
I/) 0.5 >

: ^ 1 •51 2 3 4 5 10 20 50

Dimensionless Length of Nozzle L/D

Fig. 6. Variation of discharge coefficient with the dimensionless

length of nozzle (L/D).

Table I. Naming procedure of nozzles.

Shape of Hole L/D Lot

entrance diameter number

"Straight" "0.3mm" "^" "3"

S - 03 - 0^ - 5

577
200

150

E

=1-100

Q

50

30

R03-03-0^

PDPA

Water

1=1000 mm

S-03

\ ^

-04-3 ^

/C03-03-04

200

150

E

^100

CM •oQ

50

80 100 150 200 250 300

Mean Injection Velocity Vm m/s

30

PDPA

Water

I =1000 mm

S-03- DA-3

o

R03-03-0A'

C03-0 3-04

\a

80 100 150 200 250 300

Corrected Velocity Vc m/s

Fig. 7. Variation of droplet size with Fig. 8. Variation of droplet size with

the mean injection velocity, L/D = 4. the corrected injection velocity, L/D=4.

Fig. 9. Droplet size dependence on the

mean injection velocity, showing the

variation of individual nozzle having

the same specifications.

Fig. 10. Droplet size dependence on the

corrected injection velocity, variation

of individual nozzle disappears and

every dependence can be expressed by

a unique line.

578
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 63

A STUDY ON SPRAY FLOW ISSUED FROM PLAIN JET ATOMIZER

USING PHASE/DOPPLER PARTICLE ANALYZER

T, Inamura, N. Nagai and T. Hirai

Faculty of Engineering

Tohoku University

Sendai, Japan

ABSTRACT

Twin-fluid atomizer has been widely used in combustors, for instance spark ignition engines,

gas turbine, rocket engine and industrial furnaces, because of its high performance and simple

structure. The spray characteristics, such as the drop size, the drop velocity, the number density

and the mass flux, become increasingly important in the design of combustors and in testifying

numerical simulation of spray flow in the combustor. In spite of their importance, it was difficult to

measure these characteristics locally and simultaneously by the conventional method. On the other

hand, recently laser diagnostics technique has been developed rapidly. In particular, phase/Doppler

particle analyzer system has been popular, and the results are discussed by some researchers comparing to the conventional method. This study aims to clarify the spray characteristics of twin-fluid

atomizer and to offer the data for combustor design and the numerical simulation of a spray flow.

In this report, spatial mean drop size distribution, spatial drop velocity distribution and the relationship of the drop size and the velocity were measured by phase/Doppler particle analyzer. And

the empirical equations of spatial drop velocity distribution were deduced.

INTRODUCTION

Local and simultaneous data of spray characteristics, for example the drop size, the drop

velocity, the number density of droplets and the local mass flux of droplets are very important for

the design of combustors. However, it has been very difficult for the conventional method to obtain

these local and simultaneous data of the spray characteristics. And the designs of combustors were

generally conducted by trial and error. Recently the numerical simulation of a spray flow becomes popular in a combustion field [l],

and it has a potentiality to be applied to the design of the combustor. In that case the accurate

measurements of spray characteristics are required to validate the numerical predictions.

On the other hand, the twin-fluid atomizers have been very popular in the industrial combustors [2], because of their high performances and the flexibility of a design. Recently the airblast

atomizer has been used for the jet engine combustors instead of the pressure atomizers to decrease

soot emission [3], since the mixing between fuel spray and combustion air can be improved by the

twin-fluid atomizers.

The plain jet atomizer, which is one of the twin-fluid atomizer has simple structure and high

performances. So it has been applied to the combustor of the rocket [4], which needs high reliability.

This type of atomizer has simple structure and simple spray flow and it seems to be suitable for the

validation of numerical prediction. The present study is directed to clarify the spray characteristics

and to deduce the empirical equations of those of twin-fluid atomizers. In present report, the plain

jet atomizer was selected as the twin-fluid atomizer because of its simple spray flow. The spatial

579
drop size distribution, spatial drop velocity distribution and the relationship between drop size and

drop velocity were measured by phase/Doppler particle analyzer [5]. And the empirical equations

of drop velocity distribution were deduced.

APPARATUS AND PROCEDURES

Figure 1 shows the experimental apparatus. The atomizing air is supplied from the compressor

1 to the atomizer 11. Its flow rate is adjusted by valves 2,3,5 and metered by orifice flow meter

4. And its pressure and the temperature are measured by pressure gauge and the thermocouple.

The atomized liquid is supplied from pressure vessel 7 to the atomizer. Its flow rate is adjusted

by valves 8,10 and metered by the orifice flow meter 9. Phase/Doppler particle analyzer 12 ~ 16,

manufactured by Dantec Inc. was used to measure local drop size and drop velocity simultaneously.

The atomizer was mounted vertically in a open chamber and the spray was pulled through the

chamber by exhaust fan. And it is certified that the exhaust fan scarcely affects the drop velocity

at test section.

Fig.l Experimental apparatus Fig. 2 Structure of the plain jet atomizer

The structure of plain jet atomizer is shown in Fig. 2. The structure of twin-fluid atomizer is rich in variety and this is one of the advantages of twin-fluid atomizer. In this report the simple

structure was selected because of the purpose of this report mentioned above. The injection velocity

of the atomizing air was measured by hot wire anemometer at 1 mm downstream of the nozzle exit.

The ranges of operating conditions were, Vai=147 ~ 178 m/s(Wa=3.5 ~ 4.5 g/s), V^=0.6 ~

1.9 m/s(Wjr=0.5 ~ 1.5 g/s). Tap water was used as the atomized liquid.

RESULTS AND DISCUSSIONS

Disintegration Phenomena

Figure 3 is the magnified picture back- lighted by microflash in the vicinity of atomizer exit,

and the side-lighted streak picture of spray flow. The nozzle of 2 mm in innner diameter was used

only for taking the photograph to make the observation of disintegration phenomena easy. Figure

3(a) and (b) are in the case of low speed airflow. In the microflash photograph (a), the liquid

column is scarcely disturbed until one moves about 1.5 mm downstream of the atomizer exit. In

this region the liquid column seems not to be in contact with atomizing air stream. At 1.5 mm

580
(a) VQi=100m/s Wf=2g/s (b) Vai=100m/s Wf =2g/s

(c) Vai=16Am/s Wf = 2g/s (d) Vai=16^m/s Wf = 2g/s

Fig. 3 Disintegration phenomena and spray flow

downstream the liquid column is suddenly disturbed by air stream, and then liquid jet gets thin

gradually. This is due to the acceleration of liquid jet and the disintegration by air stream. At 15 ~

20 mm downstream the liquid column is completely disintegrated into droplets. By the observation

with the eye the tip of the liquid column vibrates right and left. And the large drops scatter to

the surrounding just after the atomization. By above observation it is deduced that under low air

speed condition the dispersion of the drops is controlled by the vibration of the liquid column, since

under low air velocity large drops are produced and a large particle maintains its initial velocity

for long time due to its large mass and low surrounding air velocity.

On the other hand in the streak photograph (b), the trajectories of the drops spread straight

from the point at about 2 mm below from atomizer exit. This point is almost coincident with the

point where the liquid column comes in contact with the air stream shown in Fig. 3(a). This means

that the dispersion of drops toward the surrounding starts at just disintegration point. The spray

angle is about 40 degree.

Figure 3(c) and (d) are in the case of high speed air flow. In the microflash photograph (c),

there exists no undisturbed liquid column. The liquid column is disturbed just below the atomizer

exit. And at 2 mm downstream the liquid column is completely disintegrated into drops. The

liquid column is disintegrated into large particles flrst, and then it is atomized again into further

fine particles. At 20 mm below the atomization seems to complete. The break-up length which

indicates the distance from the atomizer exit to the point where the liquid column is completely

disintegrated into drops is the function of liquid flow rate and the atomizing air velocity. In

particular, it decreases greatly with increase in the air velocity.

581
In the streak photograph (d), the contour of the spray spreads first. The point where the

spray contour spreads is almost coincident with the point where the atomization starts. Then the

spray contour deflates and starts to spread again at about 10 mm downstream. This point where

the spray contour spreads again is almost coincident with the point where the drops disintegrate

again into further fine drops. The spray dispersion starts after fine spray drops are produced. From

these considerations it is deduced that under high speed air condition the spray dispersion is almost

controlled by the disturbances of a air flow, since under high air velocity condition small particles are produced and a small particle follows a air flow immediately due to its small mass. Under this

injection conditions the spray angle is about 30 degree. The spray angle decreases as liquid flow

rate decreases and as atomizing air velocity increases.

Phase/ Doppler Particle Analyzer

Figure 4 compares drop size distribution measured by phase/Doppler particle analyzer (PDA)

to that measured by the liquid immersion method. The measuring point is placed at 200 mm

downstream of nozzle exit and at the center of the spray. The slide shutter was used for liquid

immersion method. The drop size distribution measured by liquid immersion method almost agrees

with the result obtained by PDA. However, by the detailed observation the result obtained by PDA

shows less small particles and more large particles. The Sauter's mean diameter measured by liquid

immersion method is 27.2 ^m and that measured by PDA is 29.8 //m. By our further investigations

xlO"'

6.0

T

£

c|"0 , . < < A.O

2.0

"0 20 40 60 80 100

d \jm

Fig.4 The comparison of drop size distributions

about the accuracy of PDA, the measuring results are influenced by three operating factors. First

one is the power of laser source. Second is the input voltage of photomultiplier and last one is the

threshold which determines whether or not the particle is spherical. High power of laser source

results in smaller drop size, high voltage of photomultiplier results in smaller drop size and severe threshold results in smaller drop size. The optimal three operating factors were determined by

the comparison of drop size distribution by PDA and by liquid immersion method and were kept

constant during the measurement. Further discussions about the accuracy of PDA are published

in other papers [6,7].

Radial Mean Diameter Distribution

Figure 5 shows the influences of the vertical distance from the nozzle exit on the radial

Sauter's mean diameter distribution. Near the nozzle exit, mean diameter decreases gradually as radial distance from the center axis increases. At a short distance from the nozzle exit, mean

Vai=229m/s Wf=1.0g/s

immersion method

PDA

582
0 10 20 30 13 10 20 30

r mm r mm

Fig. 5 The influences of vertical distance on Fig. 6 The influences of injection conditions

radial mean diameter distribution on radial mean diameter distribution

diameter decreases first and then increases gradually with increasing the radial distance. With

increasing the vertical distance from the nozzle exit the mean diameter at the center line decreases,

however, at the edge of the spray the mean diameter is almost constant. At the downstream, the

mean diameter is almost constant until r=:10 mm, then it increases gradually with increasing radial

distance.

As mentioned above, the mean diameter at the center decreases with increase in the vertical

distance. On the other hand at the edge of the spray the mean diameter scarcely changes. The

reason of this trend seems as follows. In the vicinity of nozzle exit, the mean diameter at the

center line is larger than that at the edge due to the remained liquid column at the center. Further

downstream the small particles are remained at the center because of their small momenta and

the larger particles disperse to the edge against the air stream. Consequentially the small particle

increases at the center and the mean diameter decreases. At the edge of the spray, as there are many larger particles the mean diameter is almost constant.

Figure 6 shows the dependences of radial mean diameter distribution on the injection conditions. The measuring point is placed at 40 mm downstream from the nozzle exit. Under every

injection condition the mean drop size decreases first and then increases. The mean diameter de- creases over the whole radial distance as liquid flow rate decreases and as air velocity increases.

The decrease of mean diameter is larger at the edge of the spray rather than at the center. At the

edge of the spray there are many particles produced by the shear force of the air stream on the

surface of liquid column. And the drop sizes of the particles produced by the shear force on the

surface of liquid column are greatly influenced by the air velocity and the shear force.

Radial Drop Velocity Distribution

The dependences of radial drop velocity profile of 1 ^m(it 0.5 fim) drops on the vertical

distance from the nozzle exit are shown in Fig. 7. Figure 8 shows that of 60 ^m(± 0.5 ^m) drops.

Near the nozzle exit{h < 40 mm) the velocities of large particles at the center line are smaller than

those of small particles. By further investigations, at the edge of the spray the velocities of small

particles are smaller. This trend indicates that the center part of the spray is the region in which

the drops are accelerated by the air stream and the surrounding of the spray is drop decelerating

region. Because the small particles are greatly influenced by the air stream due to those small

momenta.

On the other hand further downstream, the velocities of small particles are smaUer than those

of large particles over the whole radial distance. This trend means that further downstream the

drop velocities are larger than the air velocity over the whole radial distance.

Figure 9 shows the drop velocity at the center as a function of the vertical distance for various

drop sizes. The velocity changes of small particles are larger than those of large particles. This

trend indicates that small particles can be easily accelerated and decelerated because of their small

momenta. The drop velocities for all drop sizes have a maximum velocity at about h=50 mm. This

583
Fig. 7 The dependences of radial drop velocity Fig. 8 The dependences of radial drop velocity

profile on vertical distance (d=:l /xm) profile on vertical distance (d=60 fim.)

50 100 150 200

h mm

Fig. 9 The variation of drop velocity at the center

with the vertical distance

means that at the center the region within 50 mm downstream of the nozzle exit is drop accelerated

region, whereas the region more than 50 mm downstream is drop decelerated region.

Figure 10 shows the validity of the dimensionless empirical equation of radial drop velocity

profile of 1 //m(±0.5 fim) drops. And Fig. 11 shows that for 60 ^m(±0.5 /xm) drops. The axis of

abscissa is the radial distance made dimensionless by half the width of the radial drop velocity

profile at half the drop velocity on the center line, and the vertical axis is the drop velocity made

dimensionless by respective drop velocities on the center line. The broken lines indicate the following

empirical equation deduced by Yatsuyanagi [8] for the air velocity distribution of gas liquid twophase flow:

1.5 ^

Vr = (1-

pc 2.267 -Tp^

(1)

The experimental results of drop velocities are almost represented by the broken line for all drop

sizes. However, at the edge of the spray(r/rp^ > 1.3) there exists the trend that the experimental

results are slightly larger than those of eqn.(l) for all drop sizes.

584
Fig. 10 The comparison of radial drop velocity

profile (d=l iim)

CONCLUSIONS

0.A 06 0.8 1.0 1.2 r/fpv

Fig. 11 The comparison of radial drop velocity

profile (d=60 fim)

In order to clarify the spray characteristics issued from the plain jet atomizer, the spray

characteristics, that is spatial drop size distribution, spatial drop velocity distribution and the

relationship between drop size and drop velocity were measured by phase/Doppler particle analyzer.

The disintegration phenomena were observed by the microflash pictures and the streak pictures.

Consequently the following results were obtained:

1. Under low air speed condition the dispersion of the drops is controlled by the vibration of

the liquid column, whereas under high air speed condition the dispersion is controlled by the

disturbances of a air flow.

2. Both of the break-up length and the spray angle decrease with increasing atomizing air velocity

and with decreasing the atomized liquid flow rate.

3. In the vicinity of the nozzle the drop size decreases with increase in radial distance, whereas

further downstream the drop size increases. And at the center of the spray the drop size

slightly decreases as one moves downstream, whereas at the edge of the spray the drop size

is almost constant.

4. In the vicinity of nozzle exit there exists the drop accelerating region. In this region the

smaller the particle, the larger its velocity is. On the other hand further downstream there

exists the drop decelerating region. In this region the larger the particle, the larger its velocity

is.

5. The spatial drop velocity profiles can be represented by the empirical equation (1) for all drop

sizes.

NOMENCLATURE

d

d32

h

nN

r

v.,

drop diameter

Sauter's mean diameter

vertical distance from the nozzle exit

the number of drops

total number of drops

radial distance

half the width of the radial drop velocity profile at half the drop velocity at the center

injection velocity of the atomizing air

liquid injection velocity

mean drop velocity

585
mean drop velocity on the center line

air mass flow rate

liquid mass flow rate

ACKNOWLEDGEMENTS

The authors would like to thank Mr. T. Mori, Mr. Y. S. Kim and Mr. H. Okuyama of Tohoku

University for their help in the experiments.

REFERENCES

1. Faeth, G. M., "Evaporation and Combustion of Sprays", Prog. Energy Combust. Sci., Vol.9,

pp.1-76 (1983).

2. JSME, RC-82 Report (Chairman, Nagai, N.) (1990) (in Japanese).

3. Sato, Y., "Current Design Status and Design Trend of Fuel Injectors for Jet Engines", Journal

J.S.A.S.S., Vol.34, No.393, pp.519-530 (1986) (in Japanese).

4. Yatsuyanagi, N., "A Study on a Spray Flow Issued from Plain Jet Air Blast Atomizer (3rd

Rep.)", Trans. J.S.M.E., Ser.B, Vol.51, No.471, pp.3599-3607 (1985) (in Japanese).

5. Bachalo, W. D., Hess, C. F. and Hartwell, C. A., "An Instrument for Spray Droplet Size and

Velocity Measurements", J. Eng. Power, Vol.102, pp.798-806 (1980).

6. Dodge, L. G. and Reitz, R. D., "Comparison of Drop-Size Measurement Techniques in Fuel

Sprays: Malvern Laser-Diffraction and Aerometrics Phase/Doppler", GMR-5382, General

Motors Res. Rep., (1986).

7. Yoshizu, F., Arai, M. and Nakayama, M., "Correlation between Measured Particle Size Using

Laser Diffraction and Phase Doppler Velocimeter Measurement Systems", Proc. J.S.M.E.

Meeting, No.900-59(B), pp.509-511 (1990) (in Japanese).

8. Yatsuyanagi, N., "A Study on a Spray Flow Issued from Plain Jet Air Blast Atomizer (1st

Rep.)", Trans. J.S.M.E., Ser.B, Vol.51, No.466, pp. 1857-1866 (1985) (in Japanese).

586
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 64

SPRAY CHARACTERISTICS OF A PILOT INJECTION SYSTEM

OPERATING IN AN IDLING D.I. DIESEL ENGINE

F. Yoshizu and M. Nakayama

Department of Mechanical Engineering

Gunma University

Kiryu, Gunma, Japan

ABSTRACT

A new Diesel Injector has been developed, utilizing pilot

Injection to control the Injection rate. It offers many

advantages for use In small to medium DI dlesel engines, such

as: control of Initial spray tip penetration, good atomlzatlon,

and small droplet size. This paper presents the spray

characterization of this Injector at idle conditions. Including

time-dependent spray penetration, mean particle diameter, local

mean particle diameter on radius direction, emissions and

combustion noise.

INTRODUCTION

Many current designs of high speed direct Injection (HSDI)

Diesel Injector systems are of the initial throttle injection

type that use a smaller orifice in the vicinity of the needle

valve. The purpose of this narrower opening is to reduce both

the emissions and combustion noise. Typical of throttle injection

nozzles, are the two spring type detailed by Greeves and Wang^-*-^,

and by Kageyama'^ where the needle may be stopped in the middle

of the lift process. The results from this study showed that

emissions of HC and NOx were reduced, with CO and smoke emissions

remaining unchanged. The combustion noise was also reduced

without any change in fuel consumption. The main reason for these

improvements was a reduction in the amount fuel induced during

the onset of the injection process, allowing better precombustion

control^^. In addition to this the amount of unburnt fuel

colliding with the cylinder walls was reduced. However both

throttle and two-spring type Injectors result in Increased HC

emissions at set NOx emissions levels. This problem being most

prevalent at starting and continuous idling operating conditions.

An Injection nozzle offering a solution to this problem would

therefore have to reduce HC emissions (at constant NOx levels)

while simultaneously offering reduced idling combustion noise.

This paper therefore details the preliminary design of a novel

pilot Injector system not using the throttle nozzle principle. To

allow both the emissions and combustion noise to be reduced, it

was thought that the use of a pilot Injection system would be

beneficial. At first the atomlzatlon characteristics of this

nozzle were investigated at atmospheric conditions, with

subsequent tests carried out in a small DI engine allowing both

emissions and noise tests to be carried out.

587
INJECTOR SYSTEM DESIGN CONCEPT

It is known that the main factor causing HC emissions at

idling conditions is the complete throttling of the entire

injection period. The result of this process is a larger pressure

difference between the chamber and nozzle, leading to an increase

in the average size of spray droplet diameter. The purpose of the

pilot injection system was to control the average droplet size,

while retaining the spray penetration characteristics of throttle

type injectors, which are known to be satisfactory.

The important operating features were listed as follows:

(1) Short spray penetration at ignition delay.

(2) Small mean spray droplet size.

(3) Stratified spray droplet size distribution on radial

direction from center to outside.

(4) Stability of injection from idling to maximum speed

and maximum load.

(5) Stability of initial injection characteristics for required

operating lifetime.

(6) Easy setting and maintenance free operation.

In order to satisfy these requirements, the main design

characteristics of the new injector were as follows.

(1) Correct pilot injection characteristics.

(2) Main opening pressure higher than pilot opening pressure.

(3) Easily adjusted distance between pilot and main injection.

(4) Constant pilot injection quantity.

(5) Use of conventional nozzle types.

EXPERIMENTAL SYSTEMS

Description of Injection Systems

The types of nozzle used in this investigation are shown in

Fig. 1.1-1. 4, Figs. 1.1-1. 2 show a single spring injector in which

two nozzle types (Standard and throttle type) were used. The

nozzle shown in Fig. 1.2 has a clearance between the tip of the

needle valve and the nozzle body. Fig. 1.3 shows a two spring

injector, in which the pre-lift value of the first spring, when

used with a standard nozzle, gives the flow characteristics shown

in Fig. 2. This figure also compares the features of the throttle

nozzle. It can be seen that at the pre-lift value of the two

spring injector ( 0 . 04mm) , the volume flow rate of both nozzles is

the same

.

The new injector design is shown in Fig. 1.4. The system uses

a standard nozzle design, with a single stage lift arrangement.

However in addition to the standard fuel path leading to the

nozzle area, there is a second fuel line leading to a Dodge

Plunger* assembly

.

* A valve system allowing leakage past the plunger to force an

early valve closure thus allowing Pilot injection.

588
4 ) P i 1 ol-I n j.

Fig-. 1 Details of injection systems

The relationship between the pressurized area of the Dodg-e

plunger and the nozzle was selected according to the following

relationship

:

'DPS < (Ajsj - A^^) < A NS 'DP (1)

Where

:

ADPS

'DP

^NS

'N

Dodge Plunger seat area

Dodge Plunger cross section area

Nozzle seat area

Needle cross section area

Before the start of injection the pressure/area relationship

is stable according to the relationship - Aj)Ps*^^N~^NS • start of an injection cycle the needle valve lifts and

subsequently, when the opening pressure of the Dodge Plunger is

reached it also lifts. The movement of the Dodge Plunger creates

a drop in the line pressure due to the dead volume in the the

Plunger. At this time a new pressure/area relationship is created

- Aj^<Aqp - signifying the end of the pilot injection process,

this is also accompanied by the close of the needle valve. The

fuel line pressure then increases producing a second lift in the

needle valve, with the Dodge Plunger remaining open from the

previous line pressure increase. This process begins the main

injection

.

589
Tf =iO'

C

Pf = 1 OMRs

. STD-N

A A T/H-N

1000

: BOO o

CO\

— 600

o^ 200

1 /

/

/

/

>

//

i

t * '

PL : « I I mm

Vcaa.Bu: 0. Sm/s at 400rpa Plpe:«l. 8-400a

NPo: 18. SMPa

Np=400rpm

Q = 6 mm"/ a t

STD -InJ. T/H -InJ.

2Spr. -InJ.

PI lol-lnj.

a.5

a.a.

30

20

10 -

0

0. 3

J

O. 2

-

X 0. I -

0

Fig,

1.0 0.1 0.2 0.3 0.4. 0.5 0.6 0.7 0.8 0.9 Needle lift (mm)

2 Standard and throttle nozzle

flow characteristics

1 . J

?!

/

ilb

30

20

10

J O

a.5

z

a.

M

V

a

o

CfiM fiNGLE (DEG)

Fig, 3 Injection system

characteristics

When the line pressure begins to drop, signifying the end of

the injection process, the needle valve closes first. This is

followed by the closure of the Dodge Plunger when the line

pressure has fallen to the respective value.

The Dodge Plunger used in this investigation had a cross

sectional diameter of 6mm, with a seat diameter of 3.3mm. The

plunger lift was set to 0.65mm. The injection nozzles used were

of the 5 hole type, with a orifice diameter of 0.25mm. The

primary opening pressure was set at 18.5MPa. At idle operating

conditions the delivery pump speed was 400rpm, with aii engine

speed of SOOrpm. The injected volume of the fuel was 6mm /st. at

no load conditions. These test conditions were kept constant for

all the nozzles throughout the test procedure. Fig. 3 shows the

fuel injection pump plunger chamber pressure (PP/MPa) , path line

pressure (PN/MPa) , needle valve lift (H/mm) , and injection rate

(DQ/mm^deg"^)

.

Measurement of Spray Particles

Two different optical methods were used to measure the size

of the spray particles, a laser diffraction method (MALVERN 2600

PARTICLE SIZER) and a Phase Doppler method (AEROMETRICS PDPA100) . The MALVERN system performed a time dependent measurement

of particles within the laser beam volume, whereas the

AEROMETRICS system measured the local mean particle diameter on a

radial direction. The operating principle of both these systems

is detailed in references (4 & 5) and (6) respectively, only the

test results will be reported here.

590
DISCUSSION OF RESULTS

Spray Shape

Flash photographs of the sprays are shown in Fig. 4 at

varying time periods after the start of injection. The standard

injector shows both wide spray angles and long penetration,

however the throttle and two-spring injector arrangement (both

throttle type systems) show shorter penetration with smaller

spray angles. The new injector type (Pilot Injection) results in

both penetration and spray width values between the standard and

throttle type systems. The actual mean numerical values are shown

in comparison in Fig. 5.

The similarity between the Standard injector and the new

type at the smallest time delay (O.Sdeg.) can been seen. However

after the reduction in the fuel line pressure in the Pilot System

(due to the motion of the Dodge Plunger) the reduction in the

spray penetration can been seen. Also as the time lapse after

injection increases (1.5-2deg.) the termination of the pilot

injection process in the new injector can be clearly seen. When

compared to the standard nozzle this termination process allows

the air in the vicinity of the new nozzle to mix thoroughly with

the pilot fuel injection.

Cam ang.

from SOI

STD hole

injector

Throttle

hole

injector

Twospring

injector

Pilot

injector

0.5 1.0 1.5 2.0 3.0 4. Odeg

r ISOmni

Omm

Fig. 4 Injection spray photographs

Time Dependence of Spray Characteristics

Fig. 6 shows the Sauter Mean Diameter as produced by the

MALVERN diffraction system. The actual figure (D32) is an average

produced along the length of the laser beam's Intersection, with

the spray cone at a set period after the start of injection. The

position of the laser beam (diameter 9mm) was set at a constant

separation (80mm) away from the tip of the nozzle.

591
Np=400rpm

Q = 6 mmV s t

100

90

80

(mm) 70

60 c

m

«> 50 E

•> «

c

•> O. 30

20

to

0

©-

— STD -Inj. -A T/H -In J. 2Spr. -Inj.

-O Pi lot-Inj.

i' —

(

0 1 2 3 4. Cam angle (degl

from SOI

Fig. 5 Injection spray

penetration

Np=400rpm

Q 6 mm V s t

O-

. STD -Inj.

A T/H -Inj. —D 2Spr. -Inj.

O Pilot-Inj.

120

100

80 o

E 60

(SI COo

40

20

Lv =80

by u 9 1

} m

ig MALV ERN

\

! \

<

5 10 IS Cam angle (deg)

from SOI

20

Fig. 6 Sauter mean diameter

(MALVERN)

The time dependence of the Sauter mean diameter for all four

nozzles systems shows similar characteristics, with large initial

mean diameters, gradually decreasing with time. However if the

actual mean diameters of each injector are compared, both the

two-spring and throttle hole systems produce particles with a

mean diameter too large for efficient operation at idling

conditions. The standard nozzle does approach the mean size

distribution of the Pilot system late in the injection process,

but a comparison of the early mean diameters (5deg. after

injection) reveals a large difference between the two. The Pilot

system also produces a smaller size distribution much earlier in

the injection process (lOdeg. after injection).

Positional Dependence of Spray Characteristics

The AEROMETRICS PDPA system was used to investigate the

characteristics of the spray using single point measurements as

opposed to the beam integration of the MALVERN system. The

position of a single set of measurements, taken on a radius from

the center line of the jet, was again set at 80mra separation

from the nozzle exit. The distance between each measurement point

was 5mm.

Fig. 7 shows the results given by the PDPA for both Sauter

mean diameters. A comparison with Fig. 6 shows that the results

from this system and the MALVERN 2600 (Fig. 6) give approximately

the same particle size results at the jet center line. It is

known that the calculation routine for the evaluation of mean

particle diameter in the MALVERN system can under estimate the

592
500

Np

Q

: 4 0 0 r p m

: 6 mm*/ 8 t

. . STD -Inj.

A A. T/H -Inj.

C3 2Spr. -Inj.

O <D Pi lol-lnj.

Lv=80n

by using

AEROrCTRICS PDPA- 100^

g «0 ~ 300 o 200

100

20 30

LR (mm

)

FIG. 7 Sauter mean diameter

(AEROMETRICS PDPA)

sEL

Np = 8

T =

0 0 r pm

0 k X • m

SO 100 ISO 200 2S0 300 350 NO ( r p m

)

BO m 79

CM 78

77 m 76

75

CD 74. o 7372 oM 71 O 70 Z

-1 STD -Inj.

-A T/H -Inj.

-CD 2Spr. -I nj.

-O Pilot-Inj.

Im away from the fuel pump

side of the engine

Q

faT

c

e. v>N

FIG. 8 Engine emission and noise characteristics

true particle size. The absolute use of the ROSIN-RAMMLER curve

fit excludes, in some cases, large diameter particles. This

possible inaccuracy only occurs in lower mean diameter

distributions, where a small percentage of larger particles is

known to exist.

The size distribution of the Throttle hole injector is

different from the other systems tested, showing a large overall

mean diameter. The Two-spring injector also shows an overall

large size distribution. The Standard system gives a relatively

small particle size distribution. The Pilot injector however

shows small particle sizes at the outer edge of the jet, which

change to a larger mean size at L=25mm. The size distribution of

the Pilot jet is therefore termed as stratified. The change in

distribution of this jet is thought to result from the

interaction on the main and pilot jet fuel volumes. At the time

of the main injection the velocity and momentum of the main jet

exceed that of the pilot injection. The main jet therefore

collides with the previously injected fuel forcing it to the

outside of the jet arssea.

Emissions and Combustion Noise Tests

The engine used for this series of tests was a four cylinder

turbocharged type with a bore/stroke ratio of 93/102. The boost

pressure limit was set to 0 . 7bar , however this limit was not

encountered at idling conditions. The swept volume was 2771cc,

with a compression ratio of 17.5. The four injectors were all

tested with reference to NO and HC emissions. The combustion

noise was also measured in the 2 . 5KHz range 1 meter away from the

fuel pump side of the engine.

593
The results are shown in Fig. 8. At constant NO levels the

Throttle hole injector produced very high HC emissions when

compared to the remaining three injectors. The cause of this is

thought to be the high mean particle diameter, shown in Fig. 7 , leading to very slow combustion rates, and low noise levels (Fig.

8). Among these injectors the Pilot system produced the lowest HC

emission levels. The noise level of the standard nozzle is shown

to be the highest. Again reference to the mean particle diameter

(Fig. 7) shows this injector to have a small mean particle size.

The result of this is a high rate of combustion. The relatively

high level of noise produced by the Pilot system is thought to

result from a high rate of combustion in the outside, and center

of the jet only. These areas are shown in Fig. 7 to have the small

mean particle size necessary to produce this relatively high rate

of combustion.

CONCLUSIONS

A new Diesel injector system has been designed. The system

has been tested at atmospheric conditions with reference to spray

formation and particle size characteristics. The system has also

been tested at engine idling operating conditions to assess both

emissions and noise levels.

The main conclusions were as follows:

(1) Good controllability of spray tip penetration.

(2) Increased air/fuel mixing during pilot injection.

(3) Lower mean spray particle diameter.

(4) Lower particle diameter in spray center.

(5) Stratified spray diameter distribution.

It must be noted that the particle size and spray analysis

was performed at atmospheric conditions, so a direct comparison

of the injector system's performance within an engine, concerning

the emissions and noise levels, is not strictly valid. The next

step in this research is therefore to repeat the characterization

of the Injection system in an environment equivalent to an

operating engine.

REFERENCES

1) Greeves, G. , Wang, C.H.T., "Two stage lift injection and HSDI

combustion", SAE 905025 (1990).

2) Kageyama, K. , "Injection characteristics of two spring Diesel

injector", JSAE New Diesel Engine Symposium Tokyo (1983).

3) Gill, A. P., "Design choices for 1990's low emission Diesel

engines", SAE 880350 (1988).

4) Swithenbank, J., Beer, J.M. , Taylor, D.S., Abbot, D. and

McCreath, G.C., "A laser diagnostic technique for the measurement

of droplet and particle size distribution", AIAA No. 76-69 (1976).

5) Nakayama, M. , "Principles of laser diffraction theory and its

application to particle sizing", Japan Automotive Research

Institute. (JARI) (1989) Vol. 2.

6) Bachalo, W.D. and Houser, M.J., "Phase Doppler spray analyser

for simultaneous measurements of drop size and velocity

distribution". Optical Engineering Sept/Oct (1984) Vol. 23, No.

5/583.

594
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 65

A FEW USEFUL RELATIONS FOR CAVITATING ORIFICES

F. Ruiz

Department of Mechanical and Aerospace Engineering

Illinois Institute of Technology

Chicago, IL, U.S.A.

ABSTRACT

Elementary Fluid Mechanics is used to show the relationships between the

several cavitation parameters for plain orifice atomizers that have been

proposed over the years. Cavitation makes the orifice flow behave in a way

analogous to that of a coirpressible flow through a nozzle. With some sirrplifying

assurrptions, it is possible to estimate the order of magnitude of the energy

disappearing from the main flow and going into turbulence. A criterion for

cavitation inception in orifices is given, as well as an equation for the

discharge coefficient of the fully reattaching, non-cavitating flow.

INTRODUCTION

When the injection pressure is sufficiently high, or the back pressure

sufficiently low, the licjuid injected through an orifice will form a cavitation

bubble, surrounding a detached liquid flow or "vena contracta". This phenomenon

has been known for a long time [1], and has even been the basis of certain

devices [2]. Since the studies of Bergwerk [3] were published, orifice

cavitation has been considered a likely candidate to explain the phenomenon of

liquid atomization. Later experiments [4, 5] have also found evidence of

cavitation for the conditions that promote good atomization.

The appearance of cavitation is also linked to a drop of the discharge

coefficient [3, 4, 6, 7] . Indeed, when cavitation appears, a subsequent drop of

the back pressure will not increase the flow, that has become choked at the

cavitation region. There is an analogy between this choking and the choking of a conpressible flow.

Despite the obvious importance of long-orifice cavitation, studies have been

infrequent. A major obstacle found in searching though the literature is the

fact that different cavitation parameters are used by different authors . The

intent of this communication is to show the relationships between them, and to

formulate some useful expressions.

DEFINITIONS OF CAVITATION PARAMETER

Using the variables shown in Fig. 1, three main parameters have been

proposed:

1. Knapp, Daily and Hammitt [8], and the main body of researchers in

cavitation, prefer the parameter:

595
1 2

2 ~

where Poor are the unperturbed or characteristic values of the pressure and

velocity, respectively, and is the apparent vapor pressure (which includes

the pressure of the gases dissolved in the liquid) . For the particular case of

orifice cavitation, the parameter used is:

P - P

a =

-f (2)

1 2 — p V

2 ^

using the downstream pressure and the average flow velocity, V. When P2 =

P^, CT = 0, and the cavitation must extend all the way from the inlet to the

outlet, which constitutes the so-called "flip" condition. The flow does not

reattach to the orifice wall. Flip usually occurs below some value of O, larger

than zero. For a = 0, the pressure throughout the orifice is the vapor pressure,

and the liquid is atomized by flashing, if discharged after the orifice.

2. Bergwerk [3] preferred to use the parameter:

P - P

1 2 K = ^

(3)

p p

2 V

Using a variation of this parameter (assuming P^ = 0) , he found a relation

between the cavitation parameter and the length of the cavity, before

reattachment occurs. He also correlated the discharge coefficient against this

parameter and the Reynolds number. The same parameter was used by Spikes and

Pennington [7]

.

3. Hall [9] and, later, Nurick [4], used the parameter:

^1 - ^2

K' = (4)

P - P

1 V

and found that, by just applying this definition and that of discharge

coefficient, C^, and the continuity equation, one could deduce:

2 C^ = K- (5)

where C^, is the contraction coefficient, ratio of the narrowest cross-section of

the "vena contracta", to the orifice cross-section. This relation, although very

simplified (it does not include Reynolds number effects) , yields a good

approximate value of the discharge coefficient of a cavitating orifice, as shown

by Nurick [4]

.

These three numbers are related to each other. It is easy to see that:

596
and also:

T P - P P - P T J- 2 V 1 V -L — = — = — - 1 = 1 (6)

K P^- P^ P^ - P2 KP - P P - P

1 . = 1 - ' - = 1 - ^ ^ (7)

C 1 2 1 2 - p (v/cj - pv^

where is the average velocity at the vena contracta. Now, from Bernoulli

2

P. - P = T Pv! (8)

so:

P - P P - P

1 - c a = 1 = — = K- (9)

c p - p p - p

1 V 1 V

In other words:

2 1

K' = 1 - C O = (10)

<= 1— + 1

K

Similar equations for the discharge coefficient can be obtained by

substituting this result into equation (5)

.

ENERGY LOSS IN REATTACHMENT

It was mentioned above that there is a similarity between the flow in a cavitating orifice, and the choked compressible flow through a nozzle [1, 2] . Figure 2 illustrates the analogy, by representing the evolution of the flow, as the downstream pressure, P2 , is decreased.

In the corrpressible case, the flow rate continually increases until critical

conditions (M = 1) are reached at the throat, at which point the condition at

the throat cannot change further, since M = 1 must occur at the narrowest

section. In the orifice case, the flow rate also increases as P2 decreases,

until the pressure at the narrowest point of the vena contracts is equal to the

apparent vapor pressure, P^. A further decrease of P2 does not increase the

flow, because the pressure at the vena contracta cannot be smaller than P^^.

In the corrpressible case, if P2 drops below the critical value, a region of

supersonic flow appears downstream from the throat . This region ends in a shock

wave, which suddenly raises the static pressure. After this, the flow is

subsonic. In the orifice case, when P2 drops below the critical value, a cavitation bvibble surrounds the vena contracta. The liquid eventually reattaches

to the wall, within a short distance. This reattachment process raises the

pressure and, as we shall see later, is accorrpanied by energy degradation, like

in a shock wave. After reattachrr*ent, there are no cavitation bubbles.

If P2 is lowered sufficiently, in the corrpressible case, the shock wave will

move out of the nozzle, until all the flow downstream from the throat is

supersonic. Similarly, a lower value of P2 exists, for which the cavitation

region (which grows as P2 decreases) reaches the end of the orifice. At this

597
point the liquid totally clears the orifice walls, without reattachment: it is

the condition known as hydraulic flip.

The normal cavitating orifice flow is, thus, not unlike the compressible

flow in a convergent-divergent nozzle, with a shock wave in the divergent part.

The energy lost in the reattachment of the flow is partly transformed into

strong turbulent fluctuations. If the liquid is injected into air, these

fluctuations are responsible, to a large extent, for its atomization into fine

droplets. It is possible to correlate jet atomization characteristics with the

intensity of cavitation [5, 10] . The energy lost by the mean flow in the

reattachment can be evaluated in the following manner:

Assume that the flow up to the vena contracta is inviscid and that the

Bernoulli equation is applicable, then:

1 2 P - P = — pV (11)

1 V 9 " V ^

'

The average speed at the vena contracta, V^^ is equal, by continuity, to:

\

7 = —

V C

V

V = (12)

c

Now, the pressure head loss at reattachment, assuming that the reattachment

has been completed, and the velocity profile has been able to relax to a nearly

uniform shape, is:

Ah = P^ - (P^ + JpV^ (13)

so:

Ah = J

pv' + P^ - P^ + J

pv' = J

pv'(l/c' - 1) - (P^ - P^) (14)

and, from the definition of the parameter a:

P2-P^ = ajpv' (15)

so:

Ah = — pV^(l/C^ - 1 - C) (16)

2

or, using the other parameters:

1 2 K' 1 2 K 1

Ah = — pV ( 1) = — pV ( 1) (17) 2^ ^2 2^1 + K^2

c c

When there is no cavitation, there is still a loss associated with flow

reattachment, although the separation bubble is filled with liquid, not with a vapor. To find this value, one only needs to apply the momentum equation from

the section of the vena contracta, to a section after the reattachment:

P2 - = PV(V^ - V) (18)

Then:

598
- = pV (1/C^ - 1) (19)

Using this relationship, instead of (15) we obtain, finally:

Ah = pV^ (1/C - 1)^ (20)

nc 2

°

which is, of course, independent of any cavitation parameter.

CRITICAL CAVITATION PARAMETER

Ecjuations (15) and (18) also yield, by comparison, the condition for

cavitation inception, for which both are valid (there is no sudden juitp in head

loss when cavitation starts) . One obtains:

^ < ^crit = 2(1/C^ - 1) (21)

which, for the other cavitation parameters, becomes:

1 1

K > K 1 (22) c^it 2 C (1 - C ) c c

K' > K' , = + (1 - C )^ (23) crit c c

For a sharp-inlet orifice, for which C^;, = n/ (n+2) - 0.611, the critical

values are:

a^rit = 1-273 Kcrit = 1-103 K'crit = 0.525

This is, of course, just a simplification. Since the pressure at

reattachment is actually larger than we should expect that the actual O^;,^^^

will be smaller than the value given above, and K(,j.j^^ and K'^j-j^^ larger. The

pressure variation from reattachment to the outlet can be approximated by:

1 2 L

AP = — pV f— (24)

2 ^ D

where the friction coefficient, f, for the strongly turbulent flow after

reattachment, can be assumed, as a first approximation, to be independent of the

Reynolds number. An appropriate value is f = 0.04, which is the maximum

attaineible in a smooth-walled pipe, for transitional flow. Then:

O . = 2(1/C - 1) - 0.04 L/D (25) crit c

and the values of K(,j.£^, K'^j-^j^ would be obtained by the relations in (10), that

is:

K' , = + (1 - C )^ + 0.04 L/D (26) crit c c c

599
= 1 (27) K 2 2 2 crit C +(1-C) + 0.04 C L/D

c c c

A similar equation by Hall [9] for the prediction of Kj,^.^^ gave, instead, a prediction of the K for hydraulic flip as reported by Bergwerk [3] (observe in

reference [3] that the data that Hall erroneously took as inception, correspond

rather to the definition of flip) . That equation predicted considerably larger

values of K^rit^ because it was based on unusually small values of the discharge

coefficient, C^^.

Critical Discharge Coefficient

Equation (5) is valid at the limit of cavitation conditions, when cavitation

is barely occurring. In this condition, one can substitute the critical value of

the cavitation parameter, discussed above:

2 1 = 1 (28)

(1/C - 1) + 1 + 0.04 L/D

c

For the particular case of sharp-edged orifices, (C^, = 0.611), one obtains:

C = (1.4 + 0.04 L/D)'^^^ (29)

a

A relation of this type had been hinted by Sadek [11] . This equation can

also be used to estimate C^ in non-cavitating orifices, at high Reynolds number.

Figure 3 represents a corrparison of this estimation with experimental data, from

different sources (non-cavitating), after Lichtarowicz et al. [12].

ACKNOWLEDGMENT

The author wishes to express his thanks to the National Science Foundation

which, lander Grant No. CTS-901047 9 is supporting this research.

REFERENCES

[1] Ackeret, J., "Experimental and Theoretical Investigations of Cavitation in

Water". NACA Technical Memo. No. 1073. 1930.

[2] Randall, L.N., "Rocket ^^plications of the Cavitating Venturi" . ARS Journal,

Jan-Feb 1952, pp. 28-31.

[3] Bergwerk, W., "Flow Pattern in Diesel Nozzle Spray Holes". Proceedings of

the Institution of Mechanical Engineers, vol. 173, pp. 655-660. 1959.

[4] Nurick, W.H., "Orifice Cavitation and its Effect on Spray Mixing". Journal

of Fluids Engineering, vol. 98, pp. 681-687. 1976.

[5] Arai, M., Shimizu, M., and Hiroyasu, H., "Break-up Length and Spray Angle of

High-Speed Jet". Proceedings of ICLASS-85, p. IB/4/1. The Institute of

Energy. London, 1985.

[6] Kling, R., and Leboeuf, R., "Flow through Nozzles in Rocket Motors". The

Engineers Digest, vol. 14, pp. 461-463. 1953.

[7] Spikes, R.H., and Pennington, G.A., "Discharge Coefficients of Small

Submerged Orifices". Proceedings of the Institution of Mechanical Engineers,

vol. 173, pp. 661-664. 1959.

600
[8] Knapp, R.T., Daily, J.W., and Hammitt, F.G., "Cavitation". McGraw-Hill.

1970.

[9] Hall, G.W., "Analytical Determination of the Discharge Characteristics of

Cylindrical-Tube Orifices". Journal of Mechanical Engineering Science, vol.

5, pp. 91-97. 1963.

[10]Ruiz, F., "Experimental and Theoretical Analysis of High Speed Atomization,

and its ^^plication to Diesel Fuel Injection". Ph.D. Thesis. Carnegie-Mellon

University, Pittsburgh. 1987.

[ll]Sadek, R. comment to [3] and [4]. Proceedings of the Institution of

Mechanical Engineers, vol. 173, p. 671. 1959.

[12] Lichtarowicz, A., Duggins, R.K., and Markland, E., "Discharge Coefficients

for Incoirpressible Non-Cavitating Flow through Long Orifices". Journal of

Mechanical Engineering Science, vol. 7, pp. 210-219. 1965.

NC»1ENCLATURE

Cq contraction coefficient

Cjj discharge coefficient

D orifice diameter

f friction factor

Ah pressure head loss

K Bergwerk's cavitation parameter

K' Hall's cavitation parameter

L orifice length

M Mach number

P pressure

V velocity

p density

a standard cavitation parameter

subscripts

:

1 before the orifice

2 after coitplete reattachment and velocity profile relaxation

crit at cavitation inception conditions

V at the narrowest section of the cavitating vena contracta

Fig. 1 Schematic representation of Cavitating orifice flow

601
NOZZLE ORIFICE

subsonic

normal shock.

oblique shock

non-cavitating

77

cavitating

flipped

increase

F2

Fig. 2 Analogy between nozzle flow and cavitating orifice

WE'SB*C"

JO»:SSeN *ND NEKTON ff/r •0«94«

lAPPARC AND BE'D

lear regression

Fig. 3 Comparison between predictions of eq. (29) and experimental

values of (adapted from Lichtarowicz et al . [12])

602
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 66

HOLOGRAPHIC DIAGNOSTICS FOR DIESEL SPRAY RESEARCH

Z. Huang, L.-F. Zhang and B.-Z. Li

Power Machinery Engineering Dept.

Shanghai Jiao Tong University

Shanghai, P.R. China

ABSTE^ACT

This paper presents results of sttidies of the feasibility of applying

laser holography to Diesel spray research. A laser holographic and data

processing system for Diesel spray is described. The holographic methods

investigated include in-line, off-axis holography, laser shadowgraphy and their

capability for recording Diesel spray events is demonstrated. The concerning

problems in Diesel spray research including droplet size measurement in dense

fuel spray, fuel concentration analysis, two and three dimensional spray visualization are explored and discussed.

1. INTRODUCTION

Pulsed laser holography with it's ability to freeze three dimensions

dynamic test volume with high resolution, which has hitherto never been

possible by any other techniques, has become a powerful tool in the diagnostics

of fluid and particle flow field. [1] There have been some examples of using

holography techniques to measure fuel droplet of Diesel spray, with varying

degrees of success. [2,3,4] But the application of laser holograj^y to highly

dense Diesel spray is still uncertain. In other hand, laser holography has

great potentialities of providing new and valuable insights into physics of

Diesel spray formation and structure, its application remains to develop>e.

In this paper, an investigation on laser holographic methods for Diesel spray

research was conducted. The concerning problems in Diesel spray research

including droplet size measurement in dense fuel spray, fuel concentration

analysis, two and three dimensional spray visualization, ect. were explored

and discussed. Based on investigation, information on disintegrating process,

droplet distribution and fuel-air mixture formation in Diesel spray from holographic visiialization is useful to clarify the mechanism of atomization and

will greatly help to promote combustion reseairch.

2. EXPERIMEKTAL TECHNIQUE AND APPARATUS

Holography is a two-step image-forming technique for recording and reconstructing the light waves. [5] Application of the pulsed laser holograpJiy

enables a record of a high velocity transient fuel spray to be frozen in time.

603
This recording can subsequently be reconstructed into the original threedimensional form and analyzed at leisure without the deterioration of the

sample. In this work, a laser holographic and data processing system for

Diesel spray has been set up. The Diesel spray is injected into the various

ambient gas density from hole type nozzle. To maintain clean window, an electronically controlled injection device enables single injection of fuel to be

achieved with good repeatability. The system controller receives signals frcan

the needle lift transducer and triggers the pulse laser at the fixed time to

synchronize emission of the laser and fuel injection.

The layout of the recording system is shown schematically in Fig . 1 . The

ruby laser is Q-switched with a Pockels cell and emits a 30 nanosecond light

pulse. the optical wedge is used to produce two beams, which are expanded

and collimated to the size of the quartz glass window. One is led into the

vessel to be transmitted through the spray as object beam. The other one is

led to the holographic film as reference beam combined with the object beam.

All the optical components and the vessel are arrajiged on a massive isolated

table. This off-axis layout could be easily changed into in-line layout by

removing the reference beam.

The reconstruction and data processing system is shown in Fig. 2. A coherent plane wave provided by a helium-neon gas laser illuminates the hologram

1. Fuel puip rig

2. System controller

3. Pulse ruby laser

4. Air bottle

5. Quartz window

6. Transducer

7. Fuel injector

8. Pressure vessel

9. Quartz window

10. Auxiliary injector

11. Angle Barkers

12. Injection pump

1. He-Ne laser

2. Spatial filter/EC

3. Holograa

4. Reconstructed iiiage

5. Imaging lens

6. TV camera

7. Pcvision plate

8. Monitor

9. Computer

10. D/A circuit

11 .Step-motor power

12. 3-D carriage

Fig. 2 Schematic of reconstruction and data processing system.

Fig.l Schematic of holographic system.

3 If 5

- /

10

5

604
which is transported on a three-dimensional motor-driven carriage controlled

by computer to reconstruct the spray field. The reconstructed real images are

magnified and relayed into the TV camera by image lens. The system transforms

the video signal into digital signal and provides a image of matrix of 512*512

elements v^ich is quantized linearly with 256 brightness levels from 0 to 255.

The digital image is processed by computer with the threasholding and infocus

discrimination techniques to open data source for studying the spray field. [6]

3. DROPLET SIZE MEASUREMENT

The droplet size and distribution of Diesel spray have great effect on

the spray evaporation, fuel-air mixture formation and combustion. Holography

is the most attractive and only known method by which high resolution images

of individual particles in dynamic Diesel spray can be produced.

In-line far-field holography has the advantage of simplicity and should

be explored first when holographic techniques is applied to Diesel spray. The

theory of far-field holography has been well developed. [7,8] Here we will

discuss some sp>ecial consideration in Diesel spray including droplet density,

depth of the field and their limit on far-field holography.

The in-line holography is the recorded interference pattern between the

light diffracted by the object and a collinear background wave. The necessary

condition is that, first, the location of particles to be recorded must

satisfy far-field condition, second, the particle density to be measured must

be such that a significant amount of light could pass through the field without

modulation so as to serve as an effective reference beam. In order to make good

quality holograms of Diesel spray, an experiment was designed and conducted to

test the effect of particle density on the quality of reconstructed image.

Samples of glass spheres of size 50-150 Am and with SMD = 110 nm were used.

Holograms of various particle density were made. The quality of the image

including shape feature, focusing and particle intensity distribution were

examined in the reconstruction and data processing system. Fig. 3 is a example

of reconstructed image with 1 mm depth of particle field. The hologram of

particle density equal to 8 particles/mm' is seen to be good, v«^ile those of

48 particles/mm' would be considered as marginal. Whereas those of 96

paxticles/mm' are bad. In that case, light passes through the field was badly

deteriorated, contributing to noisy hologram and degradation of the reconFig.3 Effect of particle field density on quality of the image

(a) 8 (b) 48 (c) 96 particles/mm'

605
structed image. Particles can't be resolved individually.

Based on experimental results, according to the 80% fraction of transmitted light criterion, [9] the effects of various monodisperse particle diameters,

particle density and depth of the field on in-line far-field holography were

calculated. In combination with far-field condition, the limit on the holography is then found to be:

n^ 254-10V(d'l) (1)

d*4^Z$100dVA (2)

1 ^ (100dJ,;,-di„)/A (3)

Where 1 is depth of particle field, d the diameter of particle, n the

particle density, Z the far-field distance and A the wavelength of light used.

In practice, size distribution has to be considered in Diesel spray. In that

case, d in Eq. (1) could be replaced by SMD of the spray field. Fig. 4 illustrates the relationship between particle density, particle size and depth of

the field. It is seen that the smaller the jjarticle size, the shorter the

depth of the field to be measured. On the other hand, the large the particle

size, the more stringent the requirements for particle density. Therefore,

particle density and depth of the field are of vital importance to feasibility

of holographic technique and quality of reconstructed image.

The off-axis holography, in which the reference beam is added separately,

thus could make a good reference wave and easily control the 0/R ratio. It

will handle the case with higher particle density and produce higher quality

image than in-line holography. But off-axis holography is also restricted by

particle density because of multiple scattering and shadowing effects in the

object field.

The Diesel spray is very dense and its droplet density goes far beyond

the density limit established above. In other hand, Diesel droplet size is

very small and depth of the field is relatively large. Light passes through

Fig. 4 The limit on in-line far- Fig. 5 Laser shadowgraphy of Diesel spray

field holography. (a) 0.35 (b) 0.95 ms after injection.

606
the spray field was badly deteriorated due to multiple scattering and shadowing effects. The recording can then be classed as a good coherent shadowgraphy

,

but with very poor hologram characteristics. Fig. 5 illustrates the character

of such recording. Therefore in-line holography can't be used to measixre

droplet size and distribution of such dense spray field directly. It is found

that the technique may be recoverable with the use of sampling method, which

means that some special measures are to

be taken to purposely decrease density

and depth of the field to be measured.

In this work, Diesel spray was

measured by sampling a slice of the

spray, as shown in Fig. 6. It is worthwhile to note that the sampling slot

and position must be elaborately

designed to make sure of disturbing

the atomization process and flow of

spray field as less as possible and

producing representative spray sample.

[10] The image lens provide a enlarged

image of the spray close to the holographic film to relax far-field condition and depth limit. Fig. 7 is the fuel

droplet images taken of 2-mm slice of

Diesel spray, in vdiich both in-focus

and out of focus droplets can be

observed. The data of holograms were

analysed by the computer data processing system and droplet size distribution is presented in Fig. 8. Fig. 6 Schematic of sampling method.

Diesel spra;

Saipllng slot

Laser

llluiinatioD

Holographic

fill

0 6 /ff i5 20 2S 3t 35 U ^f5 5S S5 60

Droplet diaieter (ai)

Fig. 7 Reconstructed image of fuel

droplet in Diesel spray.

4. SPRAY FIELD VISUALIZATION

Fig. 8 Droplet size distribution.

The understanding of the physics of Diesel spray formation and structure

is very important indeed in producing more effective and clean engine. The

technique of holography enables a way of direct visualization of the process.

As mentioned above, although the in-line holography can not yet be used

607
for droplet size measurement in dense Diesel spray directly, it can nevertheless produce high quality shadowgraphy of Diesel spray by virtue of "freezing"

effect of pulsed laser and high resolution of holographic film, as shown in

Fig. 5. Even without reconstruction, laser shadowgraphy could well define the

spray shape and structure in two dimensions, which has never been possible by

conventional imagery.

In particular, by analysing the image of such recordings with the help

of digital image processing technique, the distribution of light intensity

transmissivity t(y) in Diesel spray shadowgraphy could be obtained, as shown

in Fig. 9. According to the light extinction theory, assuming that the spray is

of axial symmetry, the optical thickness of the particle field is thin. The

following fuel concentration C(r) of nonevaporating Diesel spray is found to

be:[ll]

f

C(r).r-dr _ o.767-P4.D»-lg{t(y)l- (4)

where R is the radius of spray cross section to be measured, r the radial

distance, the density of Diesel fuel, ^«jrt the extinction coefficient, in

general which lies very near to value of 2.0 [3] and k the film constant. Dja

lilkt inteiiiitr triniiiiiiTit; (relitive ictle)

1 I —^ 1 1 I 1 I 0 It H 31 » iO 61 Axial distance x (nun)

Fig. 9 Distribution of light intensity transmissivity of Diesel spray.

0iz5*f(>fl 0I234-S679

Radial distancse (mn) Radial distance (om)

Fig. 10 Fuel concentration of non-evaporating Diesel spray.

X is the axial distance.

608
is the sauter mean diameter of the droplets which is measured by in-line holography in this work.

By means of symmetrizing treatment and numerical method, the temporal and

spatial developnent of fuel concentration is given in Fig. 10. It is seen that

at different time and section, the the distribution of fuel concentration is

different. It is worth noticing that (1) although this method is only suitable

for non-evaporating Diesel spray, on these grounds one could well predict fuelair mixture formation and distribution of Diesel spray in the working cylinder.

(2) this method can not be uesd to calculate the fuel concentration in core

region of the spray because of incomplete break up and large optical thickness.

Off-axis holography, in which the far-field condition (in in-line holography) does not set any limit to it since the reconstructed images are separated angularly, is particularly useful in spray visualization. In this work,

the off-axis holography with direct illumination of object field has been

successfully used to record and study Diesel spray formation and structure.

New atomization phenomenon, such as wavelike, perforated dendroid, fragmental and ligamentous fuel inner structure were discovered. Fig. 11 presents

example of reconstructed images. By scanning through the images of such

recordings in data processing system, one can not only observe clearly the

process of fuel injection and atomization, but also determine the size and

shape of fuel droplets.

Fig. 12 The shape and strucFig.ll Reconstructed image of spray inner structure. ture of Diesel spray.

Based on observation, it is found that the atomization of Diesel spray

usually exhibits three modes, that is, wave, dendritic, and chaotic disintegration. [12] ITie shape and structure of Diesel spray could be classified into

core region, postatomization region and spray mantle, as shown in Fig. 12.

To enhaince the understanding of the evaporating spray field, holography

interferometry has also be applied in this work. Theoretically, this technique

will allow one to visualize not only the movement of spray but also the

development of density gradient existing in the spray field. However, it is

found that from the practical standpoint the value of such recording is uncer609
tain at this time because the spray field are so complex that interferometic

fringes are extremely difficult to interpret.

5. CONCLUSION

This investigation has demonstrated that the laser holography has great

capability for providing new and valuable insights into the physics of Diesel

spray formation aad structure. Bsused on investigation, the following conclusions can be made.

1. In-line holography is usefiiL for the measurement of the particle field,

but suffers particle density limit and tedious data reduction. But the

technique may quite be of use with the aid of the sampling method. In this

work, a density and depth limit that can be applied to in-line holographic

recording is established and a computer-aided data processing system has been

developed which could make the technique more effective and attractive.

2. Laser shadowgraphy, without reconstruction, could well define spray

shape and structure in two dimensions. Particularly with the help of digital

image processing technique, it can provide information on fuel concentration

of non-evaporating Diesel spray.

3. Off-axis holography is best suited for spray field visualization. In

this work, the inner structure of Diesel spray was firstly visualized and new

atomization phenomenon were discovered by means of off-axis holography with

direct illumination. Information on disintegrating process, droplet formation,

shape and structure of Diesel spray from the holographic visualization will

greatly help to advance these studies.

REFERENCE

1. Lee, Y.J. and Kim, J.H. : "A Review of Holography Application in MultipJiase

Flow Visualization Study" Journal of Fluids Engineering. Vol.108, Sep.

pp. 279 (1986)

2. Trolinger, J.D. and Bentley,H.T. et,al: "Application of Electro-optical

Techniques in Diesel Engine Research" SAE 740125. pp.633

3. Isao Shimizu et,al: "A Light Scattering and Holographic Technique for

Determine Droplet Size and Volume Density Distribution in Diesel Fuel

Sprays" SAE 820355. pp.155

4. Keiya Nishida and Hiroyuki Hiroyasu et,al: "Holographic Measurement of

Evaporating Diesel Spray at High Pressure and Temperature" JSME International Joural. Vol.107, pp. 30 (1987)

5. Collier,R. J. : Optical Holography. Academic Press. 1971

6. Xiao Bo, Xia Xiao-mei, Huang Zhen et,al: "Data Processing on Particle

Field Hologram." Paper presented at the 4th China ICE Diagnostics and

Measuerment Meeting. April, 1988

7. Siliverman,B.A. , Thompson, B.J. et,al: "A Laser Fog Disdrometer" Journal

of Applied Meteorology. 3, 792(1964)

8. Cartwright , S . L . , et, al: "F^rticle Sizing Using Far-field Holography"

Optical Engineering. Vol.19, NO. 5, pp.727 (1980)

9. Trolinger, J.D. : "Particle Field Diagnostics by Holography." AIAA-80-0018

10. Huang Zhen: "In-line Holography for Analysis of Dynamic Particle Field."

Thesis of Shanghai Jiao Tong University. 1990.

11. Huang Zhen, Zhang Lian-fsmg: "Measurement of Fuel Concentration in Nonevaporating Diesel Spray" Science & Technology of SJTU. No. 57, pp. 57 (1990)

12. Huang Zhen, Zhang Lian-fang: "The Laser Holographic Study on Fuel Atomization of Diesel Spray" International Symposium on CCMODIA *90 pp. 303-

308 (1990)

610
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 67

RESULTS OF THE ASTM INTERLABORATORY STUDY ON

CALIBRATION VERIFICATION OF LASER DIFFRACTION PARTICLE

SIZING INSTRUMENTS USING PHOTOMASK RETICLES

E.D. Hirleman', P.G. Felton^ and J. Kennedy*

* Mechanical and Aerospace Engineering Dept., Arizona State University, Tempe, AZ, U.S.A.

^ Dept. of Mechanical and Aerospace Engineering, Princeton University, Princeton, NJ, U.S.A.

* United Technologies Research Center, East Hartford, CT, U.S.A.

ABSTRACT

The results of the Interlaboratory Study (ILS) supporting the proposed ASTM Standard Test Method for

Calibration Verification of Laser Diffraction Particle Sizing Instruments using Photomask Reticles are presented.

The study was limited to the class of instruments known as laser diffraction particle sizing instruments which

analyze the near-forward light scattering signature of ensembles of particles and measure properties of the particle

size distribution. The standard materials used in the ILS, termed photomask reticles, consist of a quartz substrate

with a 2-D photo-etched array of randomly-positioned circular disks of chrome thin film. The reticles represent an

orthogonal projection of the cross-sections of the spherical particles in a specified 3-D reference population. The

near-forward optical scattering properties of the ensemble of circular chrome disks approximate that of the

corresponding reference particle population. Values for statistical parameters of the size distribution of the

reference population can be directly related to measurements obtained using optical microscopy and are therefore

traceable to the standard meter. The participants in the ILS represented thirteen laboratories using twenty-one

different instruments of nine different models produced by four different instrument manufacturing companies.

Each lab was asked to obtain four test results on two nominally identical photomask reticles for three different

receiving lenses. ILS results on precision for the volume median diameter Dyo.5 for three lenses indicated a withinlab repeatabiUty of about 1% (2.8a) and a between-lab reproducibiUty of about 10% (2.8o). Results for D^nj and

Dyo.9 showed greater variabilities by factors ranging approximately from 2 to 5. Bias between the mean of me ILS

test results for all lenses and the corresponding image-analysis reference values was approximately -8% for DyQ.i.

-6% for DyQ 5, and +6% for D^q o- The bias values were not statistically significant at 95% confidence in Ught of

the observed reproducibility and the uncertainty limits on the reference values.

INTRODUCTION

Since the mid-1970's ASTM Subcommittee E29.04 has been addressing standards related to the

measurement of liquid particle sizes. Since there exists a large variety of techniques and instruments for the sizing

of particles and droplets in fluid suspension, work on the comparison of the results provided by these various

instruments has been ongoing. The earliest interlaboratory comparisons of data on, for example, reference liquid

sprays showed significant variability (greater than 100% differences in measured mean drop size diameters in some

cases). More recent results fi:om an interlaboratory study have also shown significant (albeit not as severe)

discrepancies within and between the various instrument types as reported by Dodge [1].

In that context, it is important to isolate the factors causing the observed variability in order to better

understand the results and the implications of the results. Two primary probable causes for the variability are: 1)

difficulties associated with creating a reproducible Uquid spray with the reference spray nozzle; and 2) fundamental

differences in the sampling and performance characteristics of the various types of instruments used in the study.

This ELS and the associated Standard minimize the possible influence of these two factors by: 1) using a stable

artifact to simulate the liquid spray; and 2) considering only one type of liquid particle sizing instrument, i.e. those

commonly known as laser diffraction instruments.

The reference material used in this ILS, a photomask reticle, represents a 2-D projection of a 3-D particle

population. The reticle is comprised of about 10^ randomly positioned circular discs of chrome thin film deposited on a transparent optical substrate. Further details on the reference material are discussed below. The potential

usefulness of the photomask reticle in calibration verification was confirmed by an informal study of the

611
performance of laser diffraction instruments manufactured by Malvern Instruments Ltd. as reported by Hirleman

and Dodge [2]. That study generated sufficient interest to pursue development of an ASTM Standard Test Method

specifying the procedure and a quantitative assessment of the expected within-lab repeatability, between-lab

reproducibility, and bias errors which would be associated with the use of photomask reticles as a reference standard

for laser diffraction instruments. The ILS (round-robin) discussed in this paper provided the data on precision and

bias errors included in the proposed Standard Test Method. A discussion of the ILS including reference material,

participants, procedure, and results are presented below.

REFERENCE MATERIAL

The ideal calibration test samples for laser diffraction particle sizing instruments would be comprised of the

actual particle or droplet material of interest in the actual environment of interest with a size distribution closely

approximating that encountered in practice. Unfortunately the use of such calibration test samples is not currently

feasible because multi-phase mixtures may undergo changes during a test and because actual samples (e.g. a spray)

are not easily collected and stabilized for long periods of time. A photomask reticle is an artifact designed to simulate the optical-scattering properties of a 3-dimensional constellation of particles or droplets. A reticle depicts

the orthogonal projection, onto a plane, of the geometric cross-sections of all particles in the simulated 3-

dimensional particle constellation, other words, a reticle is equivalent to the 2-D image of the 3-D particle field

which would be obtained with an infinite-depth-of-field imaging system in the absence of diffraction. In addition to representing a 2-D rendering of a 3-D particle cloud, a photomask reticle also simulates the near-forward scattering

properties of the same 3-D particle cloud. It is this latter property of a reticle which makes it useful as a standard for

laser diffraction particle sizing instruments.

One problem with using a reticle to simulate a 3-D constellation of particles is that of nonuniqueness, i.e. since all information concerning the position of any particle along an axis parallel to the orthogonal projection is lost there are an infinite number of ways in which a given population of particles can be distributed in 3-D space and

stUl produce the same 2-D projection. Unfortunately the near-forward scattering properties of the constellation of a

given population of particles will also, in general, depend on the exact positions of the particles. As an example,

consider two particles positioned such that they are intersected by the same line-of-sight (i.e. if the z-axis is pardlel

to the line of sight and normal to the reticle plane, then both particles encompass some common x and y

coordinates). If these two particles were very close, say within a few diameters, then the particles would not scatter

light independently as each would influence the scattering process of the other. In that case single-scattering models

for isolated particles would not adequately predict the interactive or multiple scattering by these adjacent particles.

From a geometric optics point-of-view, one particle would fall within the shadow of the other and the light

scattering properties of the doublet would not simply be a superposition of the scattering signatures which would be

produced by the two particles if they were isolated.

Now the shadow produced by a spherical particle illuminated by a planar light beam does not actually extend

to infinity, but rather disappears some distance behind the particle due to diffraction. The length of the shadow is approximately D^/A, where D is the particle diameter and X is wavelength. After this distance the local disturbance

to the planar incident light beam is dissipated, and in some sense diffraction effects work to reconstitute a wave

which is again approximately planar. If a particle was encountered after this distance, even if it was located in the

line-of-sight, the light scattering process would occur basically independently of the presence of the first particle.

However, on a reticle the images of these two particles would be overlapped regardless of the interparticle spacing,

and it is therefore clear that a photomask reticle will not be able to perfectly match the light scattering

characteristics of all possible 3-D particle constellations which it may be representing.

The photomask reticles used in this ILS consist of a two-dimensional array of thin, opaque circular discs

(termed particle artifacts) deposited on a transparent substrate. The substrates of the two samples used in this DLS,

RR-50-3.0-0.08-102 serial numbers 246 and 247, were Schlieren-grade quartz 6.25 x 6.25 x 0.635 cm thick with a narrowband anti-reflection coating for operation at 0.6328 nm. The chrome particle artifacts were randomly

positioned within an 8 mm diameter sample area centered in the 6.25 cm square substrate. The 23 primary particle

sizes used in the discrete size distribution were approximately uniformly spaced in a geometric progression of

diameters from 4 to 91 |im. The reference population represented by the reticle contained a total of 10,441 particles

distributed over the 23 primary sizes and approximated a Rosin-Rammler particle size distribution of nominal

parameters X = 50 |im and N = 3.0. This specific particle population, if distributed within a cylinder of 8 mm

diameter and arbitrary length, would result in a projected area of covering nominally 8% of the cylinder crosssection. Hence the reticle, which represents a projection of the population, also has nominally 8% of the 8 mm

diameter sample area covered with particle artifacts. Further details of the size distribution of the particle artifacts on the reticle are available elsewhere [5].

Since the photomask reticle is used as a reference material for particle sizing instruments based on near- forward light scattering, the degree to which the reticle can approximate the near-forward scattering properties of

the reference particle population is crucial. In regimes where the diffraction approximation to light scattering is very accurate, i.e. for near-forward angles and particles large compared to the wavelength, the scattering signature

produced by die projection of a particle (an opaque circular disc on the reticle) is very close to that produced by the

612
corresponding spherical particle independent of the particle chemical composition. Further, the properties of the

reference particle population and of the near-forward scattering properties thereof can be estimated from an

independent characterization (using optical microscopy) of the number, positions, and sizes of the discs on the

reticle.

REFERENCE VALUES

A complete description of a particle population requires a large amount of information including the size,

positions, and chemical composition of each of the constituents. In order to reduce this information to a manageable

level, statistical parameters (e.g mean diameters and moments) of the particle size distribution are often used. Two

important examples of statistical parameters, the volume median diameter D^q 5 and the relative span (D^q 9

- DyQ j)/DyQ c as defined in E799-87, were used in this study. These representative parameters are quite important

in many appucations and are measured by most if not all laser diffraction instruments.

A standard test method can be used to characterize the performance of an instrument in either a relative

(precision error) or absolute (bias error plus precision error) sense. For a relative measurement, it is the

repeatability of the instrument which is under study and the stability of the reference material which must be

confirmed. For absolute measurements, the true values (and associated uncertainties) of the representative

diameters to be measured must also be known. The true values of volume parameters such as D^q ^ for a photomask reticle are those that apply to the reference particle population for which the reticle is a smiulation.

However, since the reference particle population is an imaginary entity, we must estimate these true values using

measurements on the photomask reticle combined with information on the design of the reticle. Estimates of these

true values can be established using optical or scanning electron microscope measurements of the diameters of the

particle artifacts on the reticle combined with design information and standard mathematical procedures (E799-87).

The values so established are termed image-analysis reference values and are used as the accepted reference values

in this study. The image-analysis reference values are fimdamental in that they are defined independent of

scattering or diffraction and can be traceable to a NIST length standard. The stability of DyO.S' ^® relative span,

and all other statistical parameters representative of the particle artifact size distribution for a reticle and the ability to produce nearly identical replicate copies contribute to the usefulness of a reticle as a reference material.

Several factors complicate the determination of image-analysis reference values for photomask reticles.

First, there is a problem associated with the rather large number of particles (10,441) in the reference population.

The most straightforward approach to determining image-analysis reference values involves measuring the size of

each of the numerous particle artifacts on the reticle. However, this is a very large number of measurements if accuracy to some fraction of a micrometer is desired, and acceptable levels of uncertainty can be obtained with

fewer measurements. A second complication involves the fact that some of the particles faJl within the same lineof-sight (i.e. the projections of these particles on the reticle are overlapped). The image-analysis reference values of

interest are parameters of the volume distribution of the reference particle population, and there is obviously some

ambiguity associated with estimating the volume of a particle from its 2-D projection. If the particles are assiuned

to be spherical and not in the same line-of-sight the uncertainty is small, but overlapping images complicates the

determination.

While a simulation of a "real" particle population would require a continuous distribution of particle sizes,

the photomask reticles used in this ILS have multiple rephcations of only a limited number (23) of primary particle

sizes. This facilitates basing calculated image-analysis reference values on measurements of the sizes of only a

representative subset of the particle artifacts. The image-analysis reference values reported in Table II are based on

measurements of the maximum chord (in the x-direction) of several hundred of the 10,441 possible particle

projections. The sizes of particle artifact projections that were not measured were determined statistically assuming

that the within-primary-size distribution (due, for example, to manufacturing variability) of the measured particle

artifacts is representative of all artifacts of that primary size.

Overlapped rNoncircular) Particle Artifact Projections

The particle artifact images on the photomask reticles used in this study contain overlapped images. For this reason, some additional uncertainty is introduced in determining the image-analysis reference values. Four methods

to account for these overlapped particles in determining reference values for the DyQ ^ are considered. The first is effectively the theoretical value, and the others attempt to account for the change in the apparent population caused

by loss of information due to projecting the particles onto the reticle plane.

Reference Population. In the first approach the overlapped images are neglected, and properties of the

originally-specified particle population used. In this approach, small particles which fall completely within the

boundary of the projection of a larger particle and are still counted even though they would be invisible on the

reticle. Overlapped particles are therefore treated as if the 3-D constellation of the population had very large

interparticle spacings and all particles contribute independentiy to the Hght scattering signature with equal

weighting. If the 3-D constellation were such that no portions of any particles fell within a common Une-of-sight,

the reference population values would be used.

613
Parsed analysis. In this approach partially-overlapped projections of spherical particles are counted (number

basis) as a fraction of a particle. Only the smaller of an overlapping pair of circular discs is reduced to a fractional

count, and the fraction is equal to the ratio of the nonoverlapped area (area not occluded) to the total area of a

isolated circular projection of the same primary size. For the reticles used in the ILS, this results in a number

distribution consisting of non-integer effective numbers of the 23 discrete primary sizes.

A parsed analysis would provide an exact representation of what a laser diffraction instrument would "see" if nonoverlapped portions of the smaller (daughter) particle artifacts scattered with the same relative angular scattering

signature as an isolated (nonoverlapped) circular disc but with the total scattering reduced by an amount

proportional to the fraction of the disc area which is contained in the overlapped region. This would also imply

independent scattering by the mother (larger) particle and the non-occluded portion of die smaller particle of an

overlapped pair.

Equivalent Volume of Planar Spherical-Particle Multiplet. This method assumes tiiat: 1) the partial circles

comprising a multiplet (overlapped) image are 2-D projections of intersecting spheres of the same diameter, and 2)

the centers of the spheres of the multiplet are coplanar in a plane normal to the line-of-sight (e.g. in the plane of the

reticle). A resulting volume-equivalent diameter, necessarily larger than tiiat of any of the primary singlets, is thereby assigned to each overlapping multiplet. This volume-equivalent diameter D^^ is assigned such that die total apparent volume of the intersecting particles equals

Area-equivalent weighting. This method also assumes that overlapping particle projections are due to a

multiplet particle made up of intersecting singlets. Regions where projections of two or more particle artifacts are overlapped are assumed to represent one spherical particle of the area-equivalent diameter such that the total projected area of the overlapped particles equals 7tDae^/4. The equivalent volume of the multiplet then equals

In other words, two or more interconnected, overlapped particles are assumed to scatter as one particle of

the area-eqiuvalent diameter.

Values for the various estimates of the reference values for the reticles used in the ILS are given in Table II. Size values based on the reference population are the smallest, and the area-equivalent analysis gives the largest

values for Dvo.x- Th^ differences between weightiung approaches is greatest at large volume fractions where the

effect has the "greatest influence. The area-equivalent weighting adds considerable apparent volume to each

overlapped multiplet due to die volume-of-rotation effect (the differences are not so great when considering D^o.x)-

The planar spherical-particle multiplet approach to generate an apparent volume gives smaller volumes and thereoy

lower values for Dyg x- The four values in Table II are based on reasonable models and should certainly bound the "best" value. For

that reason the values were averaged to arrive at die image-analysis reference values. The uncertainty limits in

Table II contain two bias components: 1) bias in the measuring equipment used to size the particle artifacts, and 2) one standard deviation of the values obtained from the four methods. An estimate of the resulting overall bias limits

was obtained by summing in quadrature die two components according to Abemathy and Thompson [7]. The bias

limits were added [7] to 95% confidence precision error limits due to sizing a limited sample of the particle artifacts to arrive at die total uncertainty hmits in Table II.

PARTICIPANTS

Participation in the ILS was open to any laboratory or individual with a laser diffraction particle sizing

instrament available. The ILS was announced publicly in a monthly ASTM publication and in various technical

forums where interested parties were involved. Eleven laboratories representing twenty-two instruments

participated in the ILS which ran for a littie over two years beginning in March 1988. A list of die instruments used

by the participants along widi die laboratory identifier codes used in this report are shown in Table I. A list of the

individual participants, not identified with the particular instruments, can be found in the acknowledgements.

PROCEDURE

The overall procedure followed in the ELS was based on die ASTM Standard for Interlaboratory Studies,

E691-87. A document specifying die protocol to be followed was sent to each participant along widi the most

recent version of die draft Standard Test Method for Calibration Verification of Laser Diffraction Particle Sizing

Instruments using Photomask Reticles (Draft STM). The ILS Protocol specified: the preferred options in cases where die Draft STM allowed for user choices; die conditions of die tests to be run including die laser diffraction

receivmg lens focal lengdis (63, 100, and 300 mm); die number of test results required for each sample/instrument

configuration (4); the number of samples to be analyzed (2); and die data reporting forms. The Draft STM specified

the procedure to be used in obtaining a single test result including: methods for presenting the photomask reticle to the instrument; methods for minimizing reflections; acquisition of background and signal measurements signatures;

and finally die test results to be reported which included D^q i, DyQ^, D^n 9 and volume span as defined in ASTM

E799-87. For further details the reader is referred to die Standard Test Method document which is available from

ASTM [3]. An ASTM Research Report which provides further details on die procedure is available [4].

614
ILS RESULTS

The objective of an ILS is to provide data to support precision and bias statements for a Standard Test

Method. The results of the ILS in terms of precision are summarized in Table IE. The statistical parameters

tabulated correspond to those specified in E691-87. Each row of results corresponds to the data for all instiiiments

for all lens focal lengths and the indicated test results. The repeatability (within-lab) limits r indicate the range

(95% confidence) of measured results which could be expected from measurements of a single reticle repeated over short time intervals by a single operator with a single instrument. The reproducibility (between-lab) limits R

indicate the range (95% confidence) of measured results which could be expected from measurements on a single

reticle in different labs using different instruments and possibly different focal lengths. The results indicate very

good within-lab repeatability for Dyp 5 (<1%) as would be expected since this parameter is least sensitive to noise

sources. The repeatability is somewhat worse for Dyp 1 arid Dyg 9 which fall on the tails of the size distribution.

TTie reproducibility limits are considerably larger as woiild be expected due to differences in instrument design.

The apparent bias of the test method is obtained by subtracting the mean of the series of test results obtained

in the ELS from the reference value. The apparent bias terms for each of the 3 complete cases are shown in Table

IV. The apparent bias errors are about -6% for Dyn 1 and Dyp 5 and +6% for Dyo o. However, it should also be

noted that these bias estimates are not significant (95% confidence) given the very large reproducibility limits in

Table III.

Another method for visualizing results of interlaboratory studies on which two nominally identical samples

are used is the Youden plot [6]. In tiiis ILS two test samples were used, photomask reticles RR-50-3.0-0.08- 102

serial number 246 and 247. Since the data obtained for the two nominally identical reticles were independent,

plotting the measured results against each other as shown in Figs. 1 and 2 is of interest. Data spread along the 45

deg. line (lower left to upper right) in the plots indicates between-lab variability, and data spread normal to that line

indicates within-lab variability. In standard ASTM terms, the spread along the 45-degree line is related to the

reproducibility of the Test Method, and spread normal to the Une is indicative of repeatability. Ideally, the data

would fall in a circular pattern, indicating a good balance between instrument-to-instrument variability and the

repeatabiUty limits of die test mediod. The ellipticities of die data shown in Figs. 1 and 2 range from about 3.0 to 5.0 indicating relatively poor between-instrument reproducibility.

CONCLUSIONS

An ILS was successfully carried out to quantify the precision and bias errors associated with calibration

verification of laser diffraction particle sizing instruments using photomask reticles. The within-lab repeatability

was observed to be quite good, on the order of 1% (2.8a) for Dyg 5, confirming Uie applicabiUty of die test method

for laser diffraction instrument performance verification on a relative basis. The between-lab, between-lens

reproducibilities were about 10% for tiie best possible parameter, Dyp 5, and were about 20% and 35% for Dyp j

and Dyo 9respectively. These values are rather high given the levels of accuracy desired for some particle size

measurerrient applications. However, the bewteen-lab reproducibility was somewhat better for certain instrument

configurations, and ultimately the metiiod might operate at several percent reproducibility.

ACKNOWLEDGEMENTS

The ELS was directed by a Task Group consisting of E29.04 members, who assisted in preparation of the

Standard and monitoring results. The Task Group was comprised of Dr. Phil Felton, Princeton University, Dr. E.

Dan Hirleman, Arizona State University, and Dr. Jan Kennedy, United Technology Research Center. Dr. Charles

Martin, originally affiliated with Parker Hannifin Co. and presentiy with Goodyear Aerospace, was a member of the

Task Group during the first year of the ILS. Mr. Lee Dodge, Southwest Research Institute, provided muchappreciated assistance during the study. The capable leadership of the Chairman of E29.04, Dr. Julian Tishkoff, has

been instrumental to the success of the group and is acknowledged.

Finally, the author wishes to acknowledge the participants and sponsoring laboratories who participated in

the ILS: Dr. T. Allen, E. I. Dupont, Mr. F. Biancaniello, National Institute of Standards and Technology, Mr. C.

Fagan, Parker Hannifin Corp., Dr. P. Felton, Princeton University, Mr. R. Gourley, The Munhall Co., Dr. D.

Holve, Insitec Inc., Mr. E. Hovenac, NASA Lewis Research Center, Mr. C. Hughes, Malvern Instruments USA,

Dr. R. Ingebo, NASA Lewis Research Center, Dr. J. Kennedy, United Technologies Research Center, Mr. S. Kenney, Arizona State University, Mr. J. Melvin, Parker Hannifin Corp., Mr. S. Miller, University of Texas, Dr.

S. Ridder, National Institute of Standards and Technology, Mr. J. Sager, TextiTon, Prof. P. Sojka, Purdue

University, and Mr. R. Williams, Garret Auxihary Power Division, Allied-Signal.

615
REFERENCES

1. Dodge, L., "Comparison of Performance of Drop-sizing Instruments", Applied Optics, V. 26, pp. 1328-1341,

1987.

2. Hirleman, E. D. and Dodge, L., "performance of Laser Diffraction Drop Size Analyzers", paper 85-64,

ICLASS 85 Conference Proceedings, The Institute of Energy, London, 1985

3. Standard Test Method for Calibration Verification of Laser Diffraction Particle Sizing Instruments using

Photomask Reticles, available from the American Society of Testing Materials, 1916 Race Street,

Philadelphia, PA, 19103.

4. Report on the Interlaboratory Study supporting the Standard Test Method for Calibration Verification of

Laser Diffraction Particle Sizing Instruments using Photomask Reticles, Research Report, available from the

American Society of Testing Materials, 1916 Race Street, Philadelphia, PA, 19103.

5. Hirleman, E. D., Oechsle, V. and Chigier, N. A., "Response Characteristics of Laser Diffraction Particle

Sizing Systems: Optical Sample Volume and Lens Effects," Optical Engineering, V. 23, pp. 610-619

(1984).

6. Youden, W. J., "Graphical Diagrams of Interlaboratory Test Results", Journal ofIndustrial Quality Control,

V. 15, pp. 133-137, 1959.

7. Abemathy, R. B. and Thompson, J. W. "Measurement Uncertainty Handbook", Instrument Society of

America, Research Triangle Park, NC, 1980.

Table I. Instruments used by Participants in the ILS

Diffraction

Instrument Instrument Software

Lab ID Supplier Model Version

1 Sympatec^ Helos

2 Malvern 2600HSD VF.5

3 Malvem 2600HSD M5.4

4 Malvem 2600C M3.1

5 Malvem 2600C

6 Malvem 3600EC M3.0

7 Malvem 2600C 6.10

8 Malvem 3600E M5.4

9 Malvem 2600 VF.3

10 Malvem^ 2600HSD In-house^

11 Malvem Mastersizer

12 Malvem 3600 M3.0

13 Insitec EPCS-P In-house^

14 Sympatec^ Helos

15 Malvem 2600HSD M5.4

16 Malvem 3600E M5.4

17 Malvem 2600 M3.1

18 Malvem 2600HSD VF.6

19 Insitec^ EPCS-P 1.1 20 Malvem 3600E M5.4

21 Munhall Co. PSA-32 1.90

The Sympatec Helos instrument referenced here used focal lengths of 100 and 200mm. The 200 mm data

obtained here were grouped with the 300mm data from other instruments (see note 2 below).

The Sympatec Helos instrument referenced here used a focal length of 500 mm, but the data were grouped for

analysis along with the 300 mm data (see note 1 above).

Two test results taken with an in-house diffraction instrument are included with data from this lab. For the other

test results the scattering signature was obtained with a Malvem 2600HSD and in-house inversion software was

used to obtain the reported size distribution.

In-house means the instrument or software was developed for internal use by the laboratory and is not

commercially available.

The data from this lab were obtained using the Insitec EPCS-P detector but with in-house receiving optics and

inversion software.

The Insitec EPCS-P instrument used a receiving lens of 58 mm focal length, but the data were grouped with the

63 mm results for analysis!

616
DV0.1 (urn) Reticle #246

39 40 41 42 43 44 45 46 47 48 49 50 51

DV0.5 (urn) Reticle #246

Figure 1. Plot of the final ELS test results for DyQ.i and Dyo.s- The symbols indicate the mean of the data for the

four repeated measurements for each reticle/lens combination in each lab. The error bars indicate one standard

deviation of these four test results. The dashed lines indicate the median of the data for the three different lenses,

and the ellipticity of the data for each lens are shown in the legend.

617
50 55 60 65 70 75 80 85 90 95 100 105 110

DV0.9 (urn) Reticle #246

Span (V) Reticle #246

Figure 2. Plot of the final ILS test results for D^q 9 ^nd relative span (volume basis). The symbols indicate the

mean of the data for the four repeated measuremente for each reticle/lens combination in each lab. The error bars

indicate one standard deviation of these four test results. The dashed lines indicate the median of the data for the

three different lenses, and the ellipticity of the data for each lens are shown in the legend.

618
Table II. Image-analysis Reference Value for Reticles used in the ILS.

Test

Result

Reticle

Serial #

Reference

Population

Value (|jjn)

Parsed

Number

Value dm)

Volume Equiv.

Sphere Multiplet

Value dim)

Area

Equivalent

VaJue (jiin)

Image-analysis

Reference

Value ^ dm)

Dvo.i

Dvo.l

246

247

25.74

25.82

26.16

26.24

26.39

26.46

29.63

29.60

27.0 ±2.3

27.0 ± 2.4

Dvo.5

%0.5

246

247

47.54

47.26

47.95

47.66

49.80

49.41

51.38

51.26

49.2 ±2.3

48.9 ±2.7

Dvo.9

DvO.9

246

247

69.15

69.33

69.40

69.57

72.78

72.77

78.52

78.75

72.5 ± 5.1

72.6 ±6.3

^The image-analysis reference value is taken as the mean of the four estimates. The uncertainty limits for the

image-analysis reference values represent 95% confidence intervals based on estimates of precision and bias error

in assigning the values.

Table in. Precision Statistics for the Test Method (all focal lengths grouped).

Test

Result

Reticle

Serial #

Mean of

ILS Results

X(|im)

Within-lab

Repeatability

Limit r^ dm)

Between-lab

Reproducibility

Limit R^ (|im)

Rel. Within-lab

Repeatability

Limit r/X(%)

Rel. Between-lab

Reproducibility

Limit R/X(%)

Dvo.i 246

247

25.09

25.04

±1.25

±0.91

±5.18

±5.37

±5.0

±3.7

±20.7

±21.5

Dvo.5

Dvo.5

246

247

46.05

46.05

±0.56

±0.59

±4.26

±4.57

±1.2

±1.3

±9.3

±10.0

DvO.9

Dvo.9

246

247

77.47

77.33

±3.95

±6.80

±26.50

±29.18

±5.1

±8.8

±34.2

±37.7

^The precision indices r and R, defined in E691-87, are 2.83 times the repeatability (within-lab) and

reproducibility (between-lab) standard deviations, respectively.

Table IV. Bias of the Test Method (all focal lengths grouped).

ILS Results Image-analysis Bias Estimates

Test

Result

Reticle

Serial #

Mean

Value dm)

Reproducibility

Limit R (|im)

Reference Uncertainty Measured Statistically

Value dim) Limits ^ (|jjn) Bias(|jjn) Significant?

Dvo.l

%0.1

246

247

25.09

25.04

±5.18

±5.37

27.0 ±2.3

27.0 ±2.4

-1.9

-2.0

No

No

Dvo.5

246

247

46.05

46.04

±4.26

±4.57

49.2 ±2.3

48.9 ±2.7

-3.1

-2.8

No

No

Dvo.9

%0.9

246

247

llAl

n.33

±26.50

±29.18

72.5 ± 5.1

72.6 ±6.3

5.0

4.7

No

No

iThe uncertainty limits for the image-analysis reference values represent 95% confidence intervals based on

estimates of precision and bias error in assigning the values. The statistical significance indication also is for 95%

confidence.

619
1

I

I

I

i

i

I

I

I

I
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 68

CORRECTING MALVERN PARTICLE SIZE MEASUREMENTS

FOR PHASE DISTORTION

R.A. Pietsch, P.E. Sojka and G.B. King

Thermal Sciences and Propulsion Center

School of Mechanical Engineering

Purdue University

West Lafayette, IN, U.S.A.

1 ABSTRACT

The two effects of phase distortion on particle size measurements made with a Malvem

Spray Analyzer are identified as beam spread and a shift of the Airy pattern extrema to higher

spatial frequencies. Beam spread is shown to be present regardless of the laser beam coherence

length (4), while the Airy pattern shift occurs only when the size of the particle being sampled

exceeds 4. A correction scheme that removes the effects of phase distortion when possible is

oudined, and then evaluated by comparing phase distorted particle size information with its

undistorted counterpart. The enhanced agreement indicates the proposed technique successfully

removes the effects of phase distortion when possible. Finally, guidelines are provided for determining when phase distortion is present.

2 INTRODUCTION

The quality of atomization cannot be ignored when considering the performance of any liquid fueled combustion system. Information about both the mean particle size and distribution of

particle sizes is important to the combustion engineer in his search for more efficient combustor

designs and alternative fuels. A popular optical method of collecting this size information utilizes Fraunhofer diffraction of a laser beam as it passes through a spray. Currently, measurement

systems based on Fraunhofer diffraction are commercially available; the Malvern Spray

Analyzer is one such system. Unfortunately, recent studies of particle sizing in high temperature

turbulent environments have shown that this type of instrument, in its current form, is incapable

of accurately measuring the scattered light profile and yielding the desired particle size data

[1,2]. The inaccuracies noted in References 1 and 2 result from opdcal perturbations present in

the system being studied.

The primary perturbing optical effects associated with laser beam propagation in a high

temperature turbulent environment are known as beam wander and phase distortion. Beam wander occurs when the laser beam encounters an index of refraction gradient and is displaced, or

steered, from its normal path of propagation. Phase distortion occurs when different points on

the same laser beam wavefront vary in phase after passing through an index of refraction gradient.

The effects of beam wander were studied in detail by Miles et al [2] and a technique was

developed to successfully compensate for this phenomenon. Miles et al also suggested that any

errors remaining in the particle distribution measurements after correcting for beam wander were

due to phase distortion. Support for their hypothesis was provided by performing a Young's

double slit experiment at various temperatures to measure Michelson's visibility, arguing that

Michelson's visibility was equivalent to the complex degree of coherence between two points on

the laser beam, and then noting that the observed decrease in coherence length was indicative of

an increase in the level of phase distortion.

Leader [3] studied phase distortion, concentrating on how it affects laser beam propagation

through weakly and strongly turbulent atmospheres. He discovered that as the turbulence

621
became stronger, phase distortion's effects were dominant over those of beam wander. Leader

also noted that as phase distortion increased, the ability to focus the beam to a given spot size

decreased. Additionally, the unfocused beam's diameter, known as the beam spread, was seen to

increase log-normally as phase distortion increased.

This increase in beam diameter is an important issue since some of the incident beam's

energy will emerge from a spray as unscattered light. In many diffraction based instruments, this

unscattered energy is removed from the particle diffraction profile by measuring the background

light intensity incident on the photodiode. Leader's work therefor suggests that a background

light intensity reading for a phase distorted beam will have a greater intensity at lower spatial

frequencies than for an undistorted beam.

Two conclusions can be drawn from the work of Miles et al and Leader. First, the effects

of phase distortion become more pronounced as the turbulence within a system increases. Second, since the amount of beam spread, and thus additional background intensity, is dependent on

the flow conditions, a separate background intensity reading should be taken at each temperature

or flow velocity being investigated in order to eliminate this effect of phase distortion.

Two issues, however, remain to be addressed. First, how is a researcher to know when the

effects of phase distortion are important, and second, are there any further effects of phase distortion that must be eliminated or compensated for.

The goal of the this work then is twofold: to develop a technique that eliminates the errors resulting from phase distortion whenever possible and to outline a procedure that will aid users

in determining when the effects of phase distortion are present. The next section provides a summary of the theory needed to accomplish these objectives.

3 THEORY

In order to develop a correction scheme to remove the effects of phase distortion on optical

particle sizing with the Malvern instrument, it is first necessary to determine if phase distortion is

present and, if so, to describe its impact. Phase distortion is present if the medium of interest

reduces the spatial coherence length of a laser beam passing through it. Its effect on Malvern

particle size measurements depends on the relative size of the perturbed laser beam's spatial

coherence length and the diameter of the largest particle to be sampled.

A level of phase distortion such that the spatial coherence length of a beam passing

through the medium is less than the diameter of the largest particle in a spray indicates that all

particles are not illuminated coherently. Fraunhofer diffraction by apertures illuminated using

partially incoherent light was studied by Shore et al [4] who showed that a gradual loss in the

intensity contrast of the diffraction signature occurs as the coherence length decreases. Furthermore, when the coherence length approaches zero, the diffraction maxima and minima shift to

higher spatial frequencies. This indicates that more light would be incident on the Malvern

detector array at larger radii, thus skewing the particle size distribution to smaller diameters.

Phase distortion on this level can be detected by employing a calibrated aperture. Specifically, a pinhole may be introduced into the phase distorting media and its intensity profile

obtained using the Malvern receiver. If the diffraction maxima and minima of the resultant Airy

pattern are shifted to higher spatial frequencies, the particle size measurements cannot be corrected. Consequently, a series of calibrated apertures is inserted into the flow with the smallest

pinhole exhibiting no shift in its diffraction extrema, for a given flow condition, indicating the

largest particle that can be accurately sized. A correction scheme was developed for this case. When no shift in the position of the diffraction extrema occurs for any particle, the complex degree of coherence, y, is used to describe the optical perturbations. By examining the scattering geometry illustrated in Figure 1, and following the notation from Hecht and Zajac [5], the

complex degree of coherence may be derived in the following fashion.

Consider two identical apertures or particles as the source of the diffracted radiation. If

they are located at a distant plane along the path of propagation, the electromagnetic fields can

be combined as: Epit) = K,E,{t - 1,) + K^E^t - (1)

622
where ti = r^lc and = Tj/c. The K factors depend on the size and location of the scatterers relative to the point P and also account for absorption, diffraction and other physical effects. Defining 1 = ^2"^! allows the time origin to be shifted and the K factors to be removed with the result

that the mutual coherence function can be defined as r,,(x)=<E,(x + t)E;(t)> (2)

Normalizing the mutual coherence function by the magnitude of the two sources gives the complex degree of coherence

<E,(x + t)E;it)> = ; z—Tji (3)

(< I E,\ ' X I E,\ ' >f'

The magnitude of y^CO is between 0 and 1. If it is equal to unity the field at P is the result of two

coherent waves out of phase by some quantifiable amount while if the magnitude is zero then the

two waves are said to be incoherent. Any value of yi2(?) between these two extremes defines

partial coherence.

Thompson [6] studied diffraction and interference from multiple apertures and arrays by

looking at the generalized interference law:

I,= i i (IJJ" I Y^(0)| cos(|3_(0) + (4)

n = 1 TO = 1 where /„ and I„ are the intensities produced by each beam individually, | is the modulus of

the complex degree of coherence between apertures n and m, and

p_(A) = 27C^A + «r^(Y^(A)) (5)

6_ = (27C/X)(r,-rJ (6)

Here and r„ are the path lengths from the two source points to the observation point, A is the

path difference divided by the speed of light, v is the mean frequency, and X is the mean wavelength of the laser beam. It is important to note that

lY^(0)l=l, n=m

lYnroCO)! =lYTOn(0)l

For a circular source illuminating a regular array of circular apertures, only nearest neighbor apertures along the array directions can be arranged to be incoherent with respect to each

other. That is, for a regular array, Yn.n + i(0) = 0 is a possible occurrence. However, because the

spacing of the Airy ring minima is not constant, some degree of coherence between an aperture

and its non-nearest neighbors will always exist. This implies that unless the coherence length is

much less than the spacing of the array, no difference will be seen in the resulting diffraction

pattern.

This observation is important to the operation of particle sizing instruments based on forward light scattering since it suggests that if phase distortion produces a beam whose coherence

length is greater than the diameter of the largest particle to be sampled then the shape of the

diffraction profile will be the same as for the unperturbed case. Thompson also investigated diffraction from irregular arrays using lycopodium powder.

Of specific interest were the diffraction profiles that resulted from using illumination whose

coherence length ranged from several thousand to only a few particle diameters. The results

showed no loss in fringe contrast with a drop in coherence length, as perhaps might have been

predicted from the work of Shore et al. The only difference was a "smoothing" of the intensity

profiles as the coherence length decreased since the values of contribute less to the sum in

Eqn (4).

In summary, the level of phase distortion can be determined by sampling the Airy patterns

produced by a series of calibrated apertures. The minimum size aperture yielding a pattern not

623
shifted to higher spatial frequencies must be greater than the largest particle to be sized or the

measurements cannot be corrected. If this condition holds, the work of Thompson shows that the

phase distorted diffraction profile will not differ from the undistorted profile. Furthermore,

under these conditions the only effect of phase distortion that need concern the user of a Fraunhofer diffraction based particle sizing instrument is beam spread. Fortunately, this spread, which

will appear as undiffracted light in the energy profile, can be removed by acquiring the

instrument's background signal under the conditions that lead to phase distortion and then processing the particle scattering data using the standard software. A demonstration of this technique is provided in Section 5.

4 EXPERIMENTAL APPARATUS

A Malvern 2600HSD diffraction based particle sizing instrument was employed in this

investigation since the Malvern serves as a sort of default standard in diffraction based systems

and Malvems have been used extensively in industry and research applications by many individuals and groups. A high temperature flow tunnel was used to generate the desired optical perturbations. It has been described in detail by Miles [7].

In order to isolate the effects of phase distortion on the particle diffraction profile, it was

first necessary to remove the effects of beam wander. This was accomplished by utilizing the

conditional sampling hardware developed by Miles et al [2]. This hardware was added to the

optical train developed to study phase distortion and is denoted by the extension ".BW" in Figure

2. It includes a 4% reflection beam splitter, a focusing lens (250 mm focal length), a fourquadrant photodetector, and electronic circuitry used to compare the laser beam's wandered position to that of the ambient situation. If the wandered beam's position is within an acceptable

window, the conditional circuitry triggers the Malvem receiver to collect the diffraction profile

information.

In order to ensure that a repeatable particle size distribution could be obtained throughout

the course of the experiments, a standard reticle was employed [8]. It had best-fit RosinRammler parameters of N = 3.20, X = 52.8 |im, and SMD (Sauter mean diameter) = 40.27 jim.

Several issues had to be considered when observing the effects of phase distortion on the

unscattered laser beam, on the diffraction pattern from the reticle, and on the diffraction pattern

from the calibrated apertures. The ability to reproduce the same diffraction pattern from either a

calibrated aperture or the reticle was limited by the ability to reproduce their exact physical location within the optical train. The repeatability of such a task was determined through trial and

error to be low, regardless of the positioning method employed. In order to observe each of the

above profiles individually, without having to reposition the reticle or calibrated aperture, the

experimental apparatus shown in Figure 2 was utilized.

The optical hardware illustrated in Figure 2 consists of a 10 mW HeNe laser whose beam

is spatially filtered and collimated to a diameter of 4.5 mm, plus two Mach-Zehnder interferometers in series that sample and redirect the laser beam and the diffraction profiles resulting from

the reticle and calibrated aperture. As shown in Figure 2, the calibration aperture and reticle are permanently positioned in separate arms of the interferometers, thus guaranteeing that their diffraction profiles are reproducible. The interferometer arrangement also allows for observation of

the effects of phase distortion on the undiffracted laser beam by adjusting the iris diaphragms to

block off both the redcle and aperture. As a result, the user can observe each of the relevant

intensity profiles individually without having to reposition the reticle or calibrated aperture

between runs. Experimental data obtained using the apparatus presented in Figure 2 are introduced in the

next section and indicate the effects of phase distortion on the standard instrument's performance, plus the enhanced agreement that results from correcting particle size data for phase distortion.

624
SLIT/APERTURE

SOURCES

DETECTION PLANE

Figure 1. Scattering geometry for two identical apertures.

eHeNe LASER

QD.BW

^ FL.BW

BS BS

IZI S 1 10

BS.BW PR IR ^ 1

IR

IR

? I s

CA BS

CSC BW

C MR

PDM

PDM - PHASE DISTORTION MEDIA BS - BEAMSPLITTER

FL - FOCUSING LENS QD - QUADRANT DETECTOR

M- MIRROR

IR - IRIS DIAPHRAGM

CA - CALIBRATED APERTURE

PR - PARTICLE RETICLE

IR DAQ

MR - MALVERN RECEIVER DAQ - DATA ACQUISITION CSC - CONDITIONAL SAMPLING

CIRCUITRY .BW - BEAM WANDER HARDWARE

Figure 2. Experimental apparatus.

RING NUMBER

Figure 3. Fraunhofer diffraction profiles for a 600 \xm diameter pinhole.

625
5 RESULTS

Data acquisition commenced by placing the calibrated aperture and reticle in their respective interferometer arms, centering the Malvern photodiode array on the laser beam, and aligning

the beam wander conditional sampling hardware. Three sets of ambient background, calibrated

aperture diffraction profile, and particle diffraction profile data were then collected by shutting

the appropriate iris diaphragms. At this point, the blower supplying air to the test section was

started and an air velocity of 25 m/s was established in the test section. The air temperature was

adjusted to 200, 400 and then 600 C with one background profile, three calibrated aperture diffraction profiles, and three particle diffraction profiles obtained at each temperature. The entire

process was performed once a day for three days for a total of nine data points at ambient and

each of the elevated temperatures.

As mentioned previously, the calibration aperture serves as a users' guide to recognize

when phase distortion is present and indicates the largest particle diameter that can be sized without phase distortion affecting the light intensity profile. Sample intensity profiles for a phase distorted and non-phase distorted aperture (circular pinhole, d = 600 \im) are shown in Figure 3.

The agreement between the two profiles indicates no shift in the maxima or minima and therefor a coherence length greater than 600 [im. Consequently, all particles up to the 564 |im limit measurable using the Malvern's 300 mm focal length lens will be coherently illuminated and the

only effect that need be considered is beam spread.

A sample phase distorted diffraction profile from the reticle appears in Figure 4. Compared to a non-phase distorted profile, the perturbed profile reveals large intensity variations at

lower spatial frequencies, i.e. in the innermost photodiode rings, due to beam spread. However,

at larger spatial frequencies no intensity variation is observable supporting the conclusion that

the spatial coherence length was larger than the largest particle in the distribution. Figure 5

shows the background readings corresponding to each of the profiles in Figure 4. By subtracting

the background taken under phase distorting conditions from the phase distorted particle profile, a corrected diffraction profile is generated and used to determine the particle size distribution.

An example profile appears in Figure 6; comparing it to the light intensity profile of Figure 4

clearly demonstrates the improvement achieved when using the correction procedure outlined

above. This improvement is also apparent from the Malvern calculated Rosin-Rammler size distribution parameters for the uncorrected and corrected profiles, as shown in Tables 1 and 2.

Uncertainties in these tables are reported as one standard deviation. Note that the discrepancies

between the ambient and 200, 400 and 600 C corrected values forX, N and SMD are within the

sum of their standard deviations, with the exception of the 600 C value for N. In contrast, the

discrepancies between the ambient and uncorrected 400 and 600 C values forX, N and SMD are

at best one standard deviation and most often greater than 2a.

6 SUMMARY

The two effects of phase distortion on the diffraction pattern produced by a particle, beam

spread and a shift of the Airy pattern extrema to higher spatial frequencies, have been noted.

Beam spread is present regardless of the laser beam coherence length (4), while the Airy pattern

shift occurs only when the size of the particle being sampled exceeds 4. A procedure to determine the coherence length under phase distorting conditions was described and a correction

scheme was developed to remove the effects of phase distortion when all particles have

diameters less than 4.

The correction scheme is straightforward, requiring only that the Malvern's background

signal be acquired under the same phase distorting conditions as the particle diffraction profile.

The data may then be processed using the standard Malvem software.

The correction technique was evaluated by comparing phase distorted diffraction profiles

and background intensities with their undistorted counterparts that were collected under identical

flow conditions. Results showed that phase distorted intensity profiles yielded particle distribution parameters that were as much as 9% in error. Parameters calculated from corrected intensity

626
/

/

n

0 S 10 15 » 2S 30

RING NUMBER

Figure 4. Fraunhofer diffraction profiles for the reticle.

1 3 S 7 8 11 13 15 17 10 21 23 26 27 2a 31 2 4 8 I 10 1 2 14 18 11 20 22 24 » 21 30

RINGNUhfiER

Figure 5. Ambient and high temperature backgrounds.

T-400C

qI

0 5 10 15 20 25 30

RING NUMBER

Figure 6. Fraunhofer diffraction profile corrected for phase distortion.

627
profiles, however, were within 2 % of those obtained under ambient conditions. This indicates

the proposed technique successfully removes the effects of phase distortion when possible.

We thank E.D. Hirleman (Arizona State University) for several helpful discussions and

B.H. Miles (US Army Waterways Experiment Station) for providing comments on an early ver- sion of this manuscript.

REFERENCES

1. Roesler, T.C., Lefebvre, A.H. and Sojka, P.E., "Fuel Atomization at High Ambient Temperatures," Central States Section/Combustion Institute paper 86-6D1, NASA-LeRC, Cleveland, OH

(1986).

2. Miles, B.H., Sojka, P.E. and King, G.B., "Malvern Particle Size Measurements in Media with

Time Varying Index of Refraction Gradients," Applied Optics, Vol. 29, No. 31, pp. 4563-4573

(1990).

3. Leader, J.C., "Beam Properties of Partially Coherent Curved Beam Waves in the Turbulent

Atmosphere," Journal Optical Society America Vol. 70, pp. 682-688 (1980).

4. Shore, A., Thompson, B.J. and Whitney, R.E., "Diffraction by Apertures Illuminated with

Partially Coherent Light," Journal Optical Society America, Vol. 56, 733 (1966).

5. Hecht, E. and Zajac, A., Optics, Addison-Wesley, Reading MA, (1979).

6. Thompson, B.J., "Multiple-Beam Interference with Partially Coherent Light," Journal Optical

Society America, Vol. 56, pp. 1157-1161 (1966).

7. Miles, B.H., "Optical Particle Sizing in a High Temperature Environment," M.S. Thesis, Purdue University (1988).

8. Hirleman, E.D., "On Line Calibration Techniques for Laser Diffraction Particle Sizing

Instruments," ASME Paper No. 83-GT-232.

Table I. Particle Size Distribution Parameters for Uncorrected Data

Temperature, C X,[Un N SMD, \im

20 52.34+0.18 3.4010.03 41.9710.04

200 52.8610.07 3.3910.00 42.3310.57

400 54.9910.73 3.2410.06 43.4710.33

600 56.2510.32 3.1710.17 44.1510.48

Table n. Particle Size Distribution Parameters for Corrected Data

Temperature, C X,[im N SMD, \un

20 52.3410.18 3.4010.03 41.9710.04

200 52.4510.00 3.3910.00 42.0010.00

400 52.0310.70 3.3510.07 41.5310.49

600 51.8511.05 3.5510.08 42.0310.61

628
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 69

A COAXIAL-BEAM DROP SIZING INSTRUMENT

FOR DENSE SPRAYS

W.M. Grissom

Physics Department

Morehouse College

Atlanta, GA, U.S.A.

ABSTRACT

A low-cost instrument has been developed to measure drop size and speed in dense sprays. The

instrument uses two coaxial laser beams of different colors. The smaller locater beam defines the center of the

larger main beemi jmd triggers a measurement only when a drop passes through the center of the main beam.

Light scattered perpendicular to the beam is focused through a double sht imaging system which avoids partial

obscuration of the image. A vibrating glass tube produces a uniform drop stream for calibration. A simple

"scattering tomography" analysis accounts for beam attenuation. Fluctuations in beam intensity are also

considered. Data in a spray nozzle are reported. Both the diameter and speed histograms are best fit by a log-normal function.

INTRODUCTION

A review of most drop sizing techniques up to 1978 is given by Azzopardi [1]. Since then the

Polarization Ratio [2,3], Phase Doppler [4], and Pulse Intensity [5-7] techniques have been developed. The

present instrument is of the latter type.

The Pulse Intensity techniques determine the drop size by measuring the scattered pulse as a drop

crosses a laser beam. This gives a robust signal with simple optics. The main problem is that the scattered

intensity depends upon the path the drop follows across the beam. Since the laser beam normally has a Gaussian intensity distribution, a drop passing far from the beam center gives a smaller pulse height than if it passed through the center. This problem is termed "Gaussian ambiguity".

An obvious solution to the Gaussian ambiguity problem is to generate a beam with a uniform ("tophat")

profile, but this has proven difficult [8]. Instead, several researchers [5,6] developed deconvolution algorithms

to analytically correct for the ambiguity. Another obvious solution is to accept a Gaussian beam, but define the

center of the beam such that measurements are triggered only when a drop passes through the center. Hess [7] used a laser velocimeter to define the beam center. Since the velocimeter signal is weak,

forward scattered light was collected by viewing the beam at an oblique angle, greatly increasing the probe

volume size. Since the velocimeter probe volume must be about 4 times wider than the drop to give a good

visibility signal and the main beam must be about 5 times wider than the velocimeter probe volume, the

minimum average drop spacing is about 20 drop diameters, which is fairly sparse.

629
EXPERIMENTAL APPARATUS

The optical arrangement is shown in Figure 1. A small yellow beam (He-Ne, 594 nm) defines the center

of a coaxial red beam (He-Ne, 633 nm). The red main beam is about 5 times larger than the yellow locater

beam at the probe volume. A similar coaxial arrangement was reported by Wang and Hencken [9]. Light

scattered at 90° is imaged onto the main slit. The slit width, together with the beam diameter defines the size

of the probe volume. The two colors are separated by an edge filter. The red light, from the main beam, is

reflected into a photomultiplier. The yellow light, from the locater beam, is re-imaged onto the "locater sht". The locater sht views only the center of the main sht, ignoring drops which are partially obscured by the

edges of the main sht. The locater photomultipUer triggers a measurement only when a drop passes through

the center of the main beam and is centered in the view of the main sht. Umhauer [10] reported a similar two

sht imaging system, with the shts in parallel.

A locater pulse triggers storage of a main photomultipher pulse with a Rapid Systems R2000 digital

O-scope. The digitized pulse is fit to a Gaussian curve by a logarithmic transform to a parabola and linear

regression. About 50 samples per pulse are needed to average over ripples on the pulse. The pulse peak

determines the drop diameter and the width determines the speed transverse to the beam. Checks on the

fitting error help ehminate drop coincidences in the probe volume. Since both size and speed are obtained,

either a flux or spatial drop size distribution can be generated and speed-size correlations can be obtained.

Since the pulse amphtude is approximately proportional to the square of the drop diameter, care must

be taken to provide a wide dynamic range. Previous researchers used logarithmic amphfiers. Instead, the

programmable gain feature of the digitizer is used. Measurements are made at three gain ranges. The

processing rate is usually slower than the data rate. To avoid bias against the busier ranges, a set of pulses is

digitized and the elapsed time recorded. The pulses are then processed and the histogram bins incremented

inverse to the digitization time.

The transmitted beam is focused onto a pinhole in front of a phototube. A 4 axis positioning system

controlled by a Precision Micro Control DCX-8 board allows translations in the scattering plane and rotations

about the spray and collection axes. A vacuum collects the spray drops, allowing operation with the spray

pointing upward.

CALIBRATION

A glass pipet is drawn to a small diameter and cleaved. A laminar jet of liquid is forced from the pipet

while vibrated by a speaker. At certain frequencies the jet breaks into a stream of uniform drops. Sakai [12]

gave a detaiiled report on the phenomenon. A time trace is shown in Figure 2. The drops almost overlap in the

main beam. The ripples in the main pulses are due to noise in the beam profile. Reducing the beam diameter

B

Fig. 1 Optical amangement

time (msec)

Fig. 2 Photomultiplier signals witli uniform 230 umD drops,

700 umD main beam, 200 umD locater beam

630
yields smoother pulses. The spatial filter partially obscures the main beam, resulting in steeper edges than a

true Gaussian pulse, which is advantageous in decreasing overlap of adjacent pulses.

Knowing the frequency and flow rate, the drop diameter is easily calculated. The caJculated size has

been verified using a microscope, with stroboscopic illumination. Photographing the drop stream and

measuring the drop spacing allows calculation of the speed. Comparing this with the e"^ pulse width

(temporal) allows calculation of the e'^ beam width, which is the calibration constant for speed.

Since the drops follow a very defined path through the beam, the drop stream must be scaimed back

and forth across the beam to test the instrument's resolution. The resulting histograms are shown in Figure 3. Equally narrow histograms are obtained for scans along the laser heam. Disabling the locater beam gives the

histograms shown in Figure 4, displaying the "Gaussian ambiguity". It appears difficult to deconvolve the

diameter histogram to appear as that in Figure 3, as suggested by Holve and Self [5] and Mizutani [6]. The

speed histogram is less affected. Theoretically, with a Gaussian beam, it should be unaffected [5]. However, in

practice the degradation is severe. Scaiming the drop stream across the laser beam also allows determination of the size of the locater

probe volume, by noting the hmits over which the locater beam is triggered. Wang and Hencken [9] found that

large drops are more likely to trigger a measurement. However, testing with drops from 163 to 279 /imD,

using a locater beam about 200 /imD and locater slit 500 ijm wide, showed no difference in probe volume size.

Therefore no correction is made for size bias. The likely explanation is that most of the light scattered at 90°

is due to total internal reflection from a small region on the back of the drop, so that the effective scattering area of a drop is much smaller than the actual cross section.

Calibration results are shown in Figure 5. Ideally, the drop diameter should be proportional to the

square root of the normalized scattered intensity, however a slight curvature is apparent as the drops increase

in size. An explanation results by averaging the beam intensity over the drop cross section, giving the fit shown

in Figure 5. The best fit occurs by assuming a main beam diameter (D) of 700 /imD, although the curvature is so slight that a much larger value would make Uttle difference. The beam diameter estimated by eye was

550 ^mD, while the speed calibrations determined 578±63/imD. The difference is again attributed to the

effective scattering diameter being smaller than the actual diameter.

For drops smaller than about 10 nmD, the light collection cone is not large enough to average over the

Mie interference pattern and the calibration curve becomes multivalued. However, few drops this small exist

in the sprays tested. Calibration streams smaller than 150 /zmD can be generated, but the drops overlap in the

probe volume, so that the lower end of the calibration curve has not been verified. One solution is to decrease

the number density by electrostatically deflecting away every other drop [13].

Calibrations using 8 different Hquids suggest a linear relation between the apparent diameter increase

and the refractive index relative to water, as shown in Figure 6. The incident light was polarized perpendicular

(vertical) to the scattering plane. For most low molecular weight liquids the correction is fairly small.

50-

40-

I

30i

20

10

230 umD drops 700 umD main beam

200 umD beater beam

200 counts

4/2/91

^^^^^^^fwnwfl^iffTTTnrniTnnfniTTnnnTHnHHrTTTTrTTfnTnTHTnHfH' 0 10 20 30 40 50 60 70 80 90 100

diameter (%range)

Fig. 3 a) Diameter histogram with drops scanned across beam

C

3

O

U

20

15

10

230 umD drops 700 umD main beam

200 umO iocater beam

200 counts

4/2/91

0 10 20 30 40 50 60 70 80 90 1 00

speed (% range)

Fig. 3 b) Speed histogram wkh drops scanned across beam

631
120

diameter (%range)

Fig. 4 a) Diameter histogram without locater beam

gospeed (% range)

Fig. 4 b) Speed histogram without iocater beam

By averaging over the beam area, a large drop gives a slightly wider pulse as it crosses the beam than a

small drop. The resulting pulse is still essentially Gaussian. Numerical integration predicts a fractional

increase in pulse width of 0.539(d/D)^ .3 4 6

^ where d is the effective scattering diameter. The present scatter

in the speed cahbrations (± 11%) preclude verifying this prediction.

BEAM ATTENUATION

In a dense spray the laser beam is attenuated in traveling to the probe volume and the scattered light

suffers attenuation in exiting the spray. These attenuations cause the drop pulses to appear smaller than

normal and must be corrected.

The transmittance to the probe volxmie is found by passing the beam from the left and right, in turn:

tji = [(tji /tf)T]'^ , where t^ and t^ are the transmittances for the beam from the left and right, respectively, and

T is the total transmittance. A linear "absorption" law (actually scattering) is assumed, i.e. T = tj^ tj.. The pulse height histograms appear similar in each case, with the abscissas simply scaled by the ratio of

the beam intensities. The ratio t^ /t^ is thus found by ratioing the peak values of the two histograms. The

collected light transmittance is the same fraction in each case, so that the ratio is unaffected. It is not actually

necesseu-y to pass a beam from the right. The same geometry results by rotating the spray 180° about the

collection axis, i.e. so that the spray points upward.

This analysis has general application to any problem where the attenuation is due partially to scattering,

and the "absorptivity" at every point in a cross section is desired, for example in stud3dng soot generation. The

632
"absorptivity" at every point along a line through the region can be found by differentiating the transmittances

along the line. This analysis is best termed "scattering tomography".

The transmittance along the collection path can be found by rotating the spray 90° about the spray axis,

as shown in Figure 7, making the former collection path (tg ) aliga with the beam, allowing its transmittance to be determined. Denote the peak of the measured pulse height histogram by Pj^^, where k-Jl or r for the

beam from the left or right, respectively, and 6 = 0° or 90° for the spray at the original orientation or rotated

90° about the vertical, respectively, tj denotes the transmissivity along the four paths shown and T^ and

are the transmittances along the two perpendicular paths across the spray. The corrected pulse height peak is found as: P = ^^0° ''(^i ^3 )• Expressing t^ and t^ with the scattering tomography analysis above eventually

results in an identity (note T J ^"^^2 "^^4)*

P =

L Tl T2

pp

Referring to Figure 7, the possibilities for the numerator are:

(tit3P)(t2t4P) = Pji 0° Pjl 180° ' Pi 0° Pr270° ' Pr90° 180° ' o^" Pr90° Pr270°

(tit4P)(t2t3P) = P^ 90° PrO° > ^Jl 90° ^Ji 270° ' ^rlSO" PrO° ' ^rlSO" ^Jl 270°

These are summarized as two procedures: (1) after the first measurement, rotate the spray 180° about

the spray axis, or (2) after the first measurement, rotate the spray 90° about the spray axis and 180° about the

collection axis (to pass beam from right). The first procedure is more convenient since the spray remains

pointing downward. It is interesting that including the collection attenuation leads to a simpler procedure than

considering the input attenuation alone. Beretta [2] states the same result for the case of an axially symmetric

spray. As seen here, the result is not restricted to axial symmetry.

A related problem is that the beam attenuation is not constant, due to statistical fluctuations in the

number of drops along the path to the probe volume. The collected light also fluctuates, but it is not significant

due to the large collection cone. A drop scatters most of the light incident upon its cross-sectional area out of the beam. An equal

amount of light is diffracted, but most of it remains within the confines of the beam. Scattering causes

localized shadows in the beam profile, giving ripples on the pulses. However, the Gaussian fitting largely

negates the ripples, so that only temporal fluctuations in the spatially averaged intensity need be considered.

Assuming a uniform beam intensity and a mono-disperse spray, the fraction of beam energy scattered

by a single drop is (d/D) ^ . The total fraction attenuated is: A = l-exp[-N(d/D) ^

], where N is the number of

drops in the beam. According to Poisson statistics, the attenuated fraction fluctuates by an rms fraction A/J N,

so the rms beam intensity fluctuation is:

= (d/D) A[-ln(l-A)]-^

For both very low and very high attenuations there is little fractional intensity fluctuation. The

maximum fluctuation occurs at an attenuation of 71.5%, in which case Ij.jj^5/I = 0.638(d/D). Given the square

root relation, the corresponding diameter fluctuations are half this. It is desirable to converge the main beam

as it approaches the probe volume, so that its average diameter through the spray is increased, decreasing the

fluctuations.

Hess [11] measured intensity fluctuations of +/-14%, at a beam attenuation of 10.7%. The beam

diameter was about 600 ^ mD at the probe volume, and somewhat larger in the spray. Since the attenuating

drops were far upstream, a factor of 2 should be added to account for diffraction. The analysis would then

predict an rms fluctuation of J 2(d/D)(0.318). The average drop size in the spray was not stated. To agree

with the above analysis, the drops would have to have been an average of 187 ^m in diameter.

633
1 sproy

e«0 e-90

TIq. 7, Input ond collecilon ottonuoLlon

1.2

0.8

0.4-

main beam

locater beam

DelavanWDB20

40 psi, z^l 52mm,

on axis, 4/4/91

Time (ms)

Fig. 8 Photomultiplier signals in spray

D.5

--0.5

140-

0 9 18 26 34 42 52 62 72 82 (umD)

Fig. 9 (a) Drop size distribution in spray

4/1 2^91 Delavan WDB20-45 deg

40 psi, z=1 SZmm, on aotis 3041 counts

log-normal fit; V=C/x exp{-0.5[ln(x/x0)/^ 2}

x0=980

s - 0.305 C= 1030

peak - 8.93 m/s

3.5 6.9 10.4 13.9 17.3 20.8 24.3

Fig. 9 (b) Speed distribution in spray

(m/s)

EXPERIMENTAL RESULTS

Measurements in a Delavcin WDB20-45° nozzle spraying water at 276 kPa, 685 cc/min, 152 mm from

the nozzle on the centerline, are shown in Figures 8 and 9. Similar shaped drop size distributions have been

reported [6,14,15].

To insure that the rapid cutoff on the low side of the drop size histogram was not due to instrument

sensitivity, the time traces were monitored. The smallest drops detected had locater pulses well above the

trigger threshold, so that drops smaller than the 5 /jniD limit in Figure 9a could have been detected. The

limiting drop size is probably due to the Umited kinetic energy available to atomize the liquid.

The log-normal function gives the best fit to both the diameter and speed histograms, as shown in

Figure 9:

f(x) = (C/x)exp{-%[ln(x/xO)/s]2} ; normalize: C = l/[s(27t)'^]

The parameters xO and s control the peak and width of the distribution. An upper limit modification to the

log-normal function has been suggested [16], however, the addition of a third pzirameter confuses the data

analysis. Such a limit is better added at the modeling stage.

634
The Nukiyama-Tanasawa (N-T) function has also been used:

f(x) = C x"^ exp[-(x/xO)P] ; normalize: C = p/{xO ^ (m + 1) T [(m + l)/p]

}

where F () is the gamma function. When exponent m is restricted to (p-1) it is termed the Weibull function

[17]. One difficulty is that the N-T function does not maintain a similar shape. For wide distributions it is skewed left, appearing similar to the log-normal function, whereas for narrow distributions it becomes skewed

right. Also, if m < 0 it predicts an infinite number of drops as the diameter approaches zero. Another popular function is the Rosin-Rammler, usually expressed as a cummulative volxmie fraction:

Fy(x) = l-exp[-(x/xO)P]

however, ^ ^

so that it is seen to be simply the Nukiyama-Tanasawa function with m = p-4, a fact noted by Bevans [18].

For application to combustion models, cm efficient procedure should exist to generate random variates

from the selected distribution function. When the cumulative distribution function (cdf) can be analytically

inverted, random variates are generated simply as X = F'^ (R), where F'^ () is the inverse cdf and R is a random number from 0 to 1 [19]. Examples are the exponential function, used in Los Alamos' KIVA code

[20], and the Weibull function. The log-normal cdf cannot be inverted, however efficient methods exist for

generating random variates from it [19, p 132].

The Delavan nozzle used was the subject of a "round-robin" test, using 6 different drop sizing

instruments [21]. At the test conditions of Figure 9a they reported Sauter Mean Diameters (SMD) from 65 to 180 jjmD and "relative spans" from 0.95 to 1.7. The present data gives an SMD of 46.3 ^mD and 1.19 relative

span. However, the SMD is a poor basis for comparing distributions from single particle counters since it is overly sensitive to the tail, where the counts are very low.

The beam attenuation through the center of the spray was 19.4%, with an rms fluctuation of 2.42%.

The previous attenuation analysis calculates an equivalent monodisperse drop size of 16.4 ^mD, for the

400 /jToD (minimum) main beam, which is close to the peak value of 14.8 /imD in Figure 9a. Doubling the

pressure to 552 kPa hardly changed the peak diameter (15.8 /jtoD) but narrowed the distribution (SMD =

36.3 ^mD), while the peak velocity increased from 8.93 m/s to 11.0 m/s.

CONCLUSIONS

A low-cost instrument has been developed to allow drop size and speed measurements in dense sprays.

The design overcomes previous difficulties with Pulse Intensity methods. A simple "scattering tomography"

analysis accounts for beam attenuation through the spray. Drop size and speed distributions in the Delavan

spray are best fit by a log-normal distribution.

ACKNOWLEDGEMENTS

The research followed a 1986 UES/AFOSR Summer Faculty Research Program (SFRP) tenure at the

Air Force Phillips Lab and a 1988 SFRP tenure at Arnold Engineering Development Center. Michael Powell

at the Phillips Lab assisted with the initial design and Chad Limbaugh at Arnold Engineering supported tests

in a model rocket injector. Research sponsored by the Air Force Office of Scientific Research, under

Contract F49620-88-C-0053.

635
REFERENCES

1. Azzopardi, B.J., "Measurement of Drop Sizes", Int. Journal of Heat and Mass Transfer, Vol 22,

pp. 1245-1279 (1979).

2. Beretta, F., Cavzdiere, A., and D'Alessio, A., "Drop Size and Concentration in a Spray by Sideward Laser

Light Scattering Measurements", Combustion Science and Technology, Vol 36, pp 19-37 (1984).

3. Presser, C, Santoro, R.J., and Semerjian, H.G., "Velocity and Droplet Size Measurements in a Fuel Spray",

ALAA 24th Aerospace Sciences Meeting, AL\A-86-0297, (Jan 1986).

4. Bachalo, W.D. and Houser, M.D., "Phase/Doppler Spray Anjdyzer for Simultaneous Measurements of

Drop Size and Velocity Distributions", Optical Engineering, Vol 23, no 5, pp 583-590, (Sept/Oct 1984).

5. Holve, D. and S.A. Self, "Optical Particle Sizing for in situ Measurements, Part 11", Applied Optics 18,

p. 1646 (1979)

6. Mizutjmi, Y., Kodama, H., and Miyasaka, K., "Doppler-Mie Combination Technique for Determination

of Size-Velocity Correlation of Spray Droplets", Combustion and Flame, Vol 44, pp 85-95 (1982).

7. Hess, C.F. "Nonintrusive Optical Single-Particle Counter for measuring the Size and Velocity of Droplets in

a Spray", Applied Optics, Vol. 23, no 23, pp 4375-4382 (Dec 1984).

8. Eauret, P.R. et dl, "Measurement of Drop Size and Velocity in Vaporizing Sprays, in Proceedings, Second

Int'l Conference on Liquid Atomization and Spray Systems. Madison, Wisconsin, p 261 (June 1982).

9. Wang, J.C.F. and Hencken, K.R., "In situ Particle Size Measurements using a Two-Color Laser Scattering

Technique", Applied Optics, Vol 25, no 5, pp 653-657 (March 1986).

10. Umhauer, H., "Particle Size Dist. Analysis by Scattering Light Meas. using an Optically Defined

Measurement Volume", Journal of Aerosol Science, Vol 14, no 6, p 765 (1983).

11. Hess, C.F., "An Instrimient to Measure the Size and Velocity of Particles in Particle Laden Flows", 21st

Joint Propulsion Conference, Moterey, California, AIAA-85-1443 (July 1985).

12. Sakai, T., Sadakata, M., Saito, M., et al, "Uniform Size Droplets by Longitudinal Vibration of Newtonian

and non-Newtonian Fluids", Second Int'l Conference on Liquid Atomization and Spray Systems. Madison,

Wisconsin, pp 37-45 (June 1982).

13. Keng, Weaver, and Orr. "A Study of Vibrating Capillary Atomizers", Georgia Institute of Technology

Engineering Experiment Station Report, Grant B-279 (April 1967).

14. Bolle, L. and Moureau, J.C., "Spray Cooling of Hot Surfaces" in Multiphase Science and Technology,

Vol 1. (G.F. Hewitt, J.M. Delhaye, and N. Zuber, Eds.), pp. 1-10, Hemisphere Publ., Washington (1982).

15. Hiroyasu, H., Kadota, T., and M. Arai, "Supplementary Comments: Fuel Spray Characterization in Diesel

Engines" in Combustion Modeling in Reciprocating Engines (Mattavi, J.N. and Ammna, C.A., Eds.),

pp. 369-405, Plenum Press, N.Y. (1980),

16. Dobbins, R.A. "Light Scattering and Transmission Prop, of Sprays", Phd. Thesis, Princeton Univ. (1960).

17. Derman, C, Gleser, L., and Olkin, I., A Guide to Probability Theory and Applications. Holt, Rinehart, and

Winston (1973).

18. Bevans, R.S., "Mathematical Expressions for Drop Size Distributions in Sprays", Conference on Fuel

Sprays. Univ. of Michigam (March 1949).

19. Dagpunar, J. Principles of Random Variate Generation. Clarendon Press (1988).

20. Amsden, A.A., J.D. Ramshaw, et al. "KIVA: A Computer Program for 2 and 3 - Dimensional Fluid Flows

with Chemical Reactions and Fuel Sprays", Los Alamos National Laboratory Report LA-10245-MS

(Feb 1985).

21. Dodge, L,G,, "Comparison of Drop-Size Measurements for Similar Atomizers", Special Report No.

SwRI-8858/2, Office of Naval Reseach, (December 1986).

636
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 70

MEASUREMENT OF DROPLET SIZES BY LIQUID NITROGEN

FREEZING METHOD

T. Karasawa*, S. Shiga* and T. Kurabayashi^

*Dept. of Mechanical Engineering, Gunma University, Kiryu, Gunma, Japan

^Hiroshima Institute of Technology, Hiroshima, Japan

ABSTRACT

A liquid nitrogen freezing method for accurate measurement of droplet sizes

has been developed since 1971 and recently the final setup has been completed

through the efforts for overcoming some problems and difficulties. Droplet

sizes of water spray are measured by both the freezing method and the

immersion sampling method and the results are compared. It is revealed that

the number of droplets caught in the freezing method is several times as much

as that of the immersion sampling method. The reason of the decrease of

droplets number in the immersion sampling method is considered to be caused by

the disappearing of smaller droplets due to evaporation. Therefore the Sauter

mean diameter obtained by the immersion sampling method tends to be larger

than that by the freezing method. Through the comparison experiment with

immersion sampling method, it has been revealed the present apparatus can

afford highly accurate value of droplet sizes.

INTRODUCTION

It is strongly desired to obtain accurate droplet sizes in the fields of

liquid atomization technique. While various measuring methods have been used,

few of them seems to be able to afford the reliable value of droplet sizes due

to individual shortcoming essential to each method. The freezing method, in

which liquid droplets are solidified by low temperature atmosphere possesses

several advantages to obtain their accurate sizes. Since the first study on

this method by Longwell(l) in 1943, several works in this field have been made

by different investigators (2)-(7). In recent ten years, however, no

successful results has reported.

The authors started the research on this freezing method in 1971. A

cylindrical freezing chamber is air-proved to avoid generation of cloud in the

chamber. Liquid nitrogen fed into the double wall space of the chamber

vaporizes and forms cooling atmosphere. Liquid to be tested is injected

downward through a nozzle and its frozen droplets fall onto a photographic

film set on a bottom plate of the chamber. In order to avoid the severe

problems of frosting and optical difficulty, a simple photographic technique

called photogram is employed to get shadows of the droplets. In 1978, at the

637
first ICLASS, history of development of the research from the first trial

setup to the sixth setup was reported and availability of this method for

accurate measurement was suggested on the basis of the experimental results.

In 1979 and 1980, remarkable improvements had been made in the method

and the seventh trial setup was designed and developed. Prior to the

construction of this device, distance neccessary for droplets to freeze was

theoretically estimated and evaporation of droplets were also calculated. The

most important improvement in the setup was employment of 100 feet long film

in a film suply system, which enables to obtain a significantly increased

number of data in one series of experiment using liquid nitrogen filled in a

150 liter Dewer tank. The film suply system was also newly designed to protect

the film against sudden exposure to the atmosphere with large temperature

difference from room to freezing chamber, to avoid the temperature shock which

deteriorates film emulsion. Using the setup, adequate exposure time to

compensate the decrease of light sensitibity of the film emulsion due to

lowered temperature, and adequate axial temperature gradient in the test

chamber to form spherical frozen particles were searched experimentally. All

these results were already reported (8).

Through further improvement made in recent years, the final setup has

been completed for accurate measurement of droplet sizes. In this paper

procedure of the improvements is described at first, and the accuracy of the

freezing method is examined by comparing the droplet sizes obtained by this

method with those by the another method. This time, the immersion sampling

method is chosen as a representative of the other methods, which is widely

used at present and is often employed as a standard method when comparison is

required.

EXPERIMENTAL SETUP AND PROCEDURE

The arrangement of experimental devices is shown in Fig.l. The freezing

chamber (32) is composed of a double wall type cylinder and a top and bottom

end plates to keep airtight for avoiding generation of fog in the chamber. The

inside of the chamber is kept dark. A length of 2 meter of the chamber was

determined considering the analytical result mentioned above, and the inner

diameter is 600 mm. The cylinder walls are made of 0.5 mm stainless steel

plate. The thickness of the previous body wall was 1.0 or 1.2 mm, and this

reducton of the wall thickness produced decreased heat capacity of the body,

nitrogen consumption economy and easy control of ambient gas temperature.

Liquid nitrogen in the Dewer tank (14) is poured into the gap between the

inner and outer walls of the chamber cylinder and the evaporated nitrogen

forms freezing atmosphere in the chamber. The amount of nitrogen supplied is

measured by using a spring balance (13). The chamber is thermally insulated

by foamed polystyrene with a thickness of 150 mm. Distribution of temperature in the freezing chamber is measured by thermocouples (29) set at seven

heights, which can be moved radially by rotating a support bar (30). Thermocouples are also inserted in other parts, injection nozzle (5), film boxes

(17) and (28), etc.. All of these sixteen thermocouples are connected to a

data-logger(l ) and their measurement time is controled by a micro computer.

When the temperature distribution in the chamber achieves required

profile, a new part of a 35 mm roll type long film is pulled out from the film

box (28) through the temperature buffer (25) at adequate speed for avoiding

heat shock, and is set on a film guide (23). On the guide the spread film

forms a collective area of frozen particles with a width of 24 mm and a length

of 600 mm. These operations are made automatically by a rolling system

638
Fig.l The eighth apparatus

consisted of a photo-detect counter (27), an inverter motor and an electronic

control circuit. Liquid to be tested is then injected downward into the

freezing chamber. When the smallest frozen particles are considered to reach

the film completely, shadows of the frozen particles are recorded on the film

by the technique called photogram, at which direct exposure is made to the

light from the bulb (2) for adequate duration determined depending on the

coldness of the film. As soon as the exposure is finished, the exposed part

of the film is rolled up into the film box (17), and the frozen particles on

the film are replaced by a suction unit (15), The photogramed particles are

printed with magnification of 20 and then measured.

Since the experiments for comparison of the two methods have to be

carried out under the same conditions as far as possible, the chamber of the

freezing method is also used to carry out the experiment of immersion sampling

method and the sampling cells are placed at the same position as that of the

film supply system in the freezing method. All the experiments are carried

out on water spray injected through a hole nozzle with 0.3 mm diameter.

Injection pressure is changed in four steps, i.e. 12, 10, 8 and 6 MPa, and the

injection period is changed also in four steps, i.e. 150, 90, 40 and 30

millisecond. The injection period of 90 ms is set as a standard injection

period. The quantities of water injected are chosen as small enough not to

cause the coalescence of droplets at the immersion liquid surface and large

enough to offer reasonable number of particles.

639
RESULTS AND DISCUSSIONS

Influence of Exposure Time to Low Temperature Atmosphere on its Negative

It is known that the application of sudden change of temperature to the film

sometimes causes unevenness in density and spots or patches with irregular

shape on its negative. However the details of these phenomena have not been

clear. Therefore the following tests were carried out. The elapsed time from

the setting of a film in alow temperature atmosphere of -100 C to the light

exposure was changed as 0, 30 and 60 minutes, and then the film was developed.

The result showed that the negative without elapsed time is normal, but a lot

of fine spots appeared on the negative with 30 minutes elapsed time and the

number of the spots increase in the case of 60 minutes. However since the

contour of the spot is not clear, while that of the shadow of the frozen

particle is quite clear regardless of its size, both the spots and the

particles can be easily distinguished. Therefore, it is considered that the

elapsed time of 30 minutes which generates relatively small number of the

spots does not affect on the result of measurement. To avoid these

undesiarable phenomena, in the present experiments the film is introduced into

the chamber immediately before the liquid injection and rolled up into the

film box as soon as the exposure is finished.

Comparison of the Number of Droplets Captured

In Fig. 2 radial profiles of number density of droplet caught in the freezing

r-o r-o

X

10

Eu\to

o.o

i

6

Freezing method

PI=10MPa(<25;]ni)

Freezing method PI=12MPa

O

Freezing method PblOMPo

Freezing method

PI=12MPa«25^m)

Freezing method

Pl=8MPo

^ PI=8MPQK25/im)

D—cr Freezing method

"PNSMPa

n K- V

^^--PI=6MPa(<25pm)

Immersion method/PI=12MPo

yPI=6MP0

0 100 200

Distance from nozzle axis r mm

300

Fig. 2 Variation of the number of

droplets caught by both

the freezing method and

immersion method

• '.. .* : ".'^ .v,» «: •

.,• JL : I

, . ..... "A . .

r» too _)

1

B

O

O

Q

_)

r-200

' .1 .

>.

:

Freezing method Immersion method

Fig. 3 Photographs of droplets

640
method is compared with those in the immersion sampling method for four

injection pressure levels. Some representative photographs of droplets in the

two methods are shown in Fig. 3. The figure shows that the number of droplets

caught in the freezing method is nine times as much as that in the immersion

sampling method, and more than half of the number is occupied by the fine

droplets smaller than 25 micrometer. These facts are caused by evaporation

of the droplets during their flights in the ambient gas, and the increased

evaporation rate as the diameter becomes small (8). In the freezing method

droplets are prevented from evaporation due to the extremely low temperature

atmosphere and almost all droplets can remain and are captured. Fig. 2 also

shows the same tendency as above.

Comparison of Sauter Mean Diameter in Two Methods

Fig. 4 shows the radial distributions of local Sauter mean diameter obtained by

freezing method and immersion sampling method. In Fig. 4(a) it is shown that

the Sauter mean diameters by the immersion sampling method are much larger

than those by the freezing method over whole range of radial derection. A

similar tendency was observed in the case of injection pressure of 10 MPa.

The reason of these facts is that in the immersion sampling method the number

of finer droplets is decreased by evaporation as mentioned above. Radial

profiles of the mean diameter in (a) are more gentle than those in (b) and

(c). This tendency may be caused by the fact that in the high pressure

injection the size of droplets becomes smaller even at the center of the

spray, and the ralatively larger droplets in the vicinity of spray axis are

apt to be shifted toward periphery due to the high injection velocity.

In figure (b) in the case of injection pressure of 8 MPa, Sauter mean

diameters by the two methods come to nearly the same at the vicinity of spray

center. This is caused by that although the number of droplets obtained by

the freezing method is larger than that of the immersion sampling method, the

total number of droplets generated decreases and the number of large droplets

over a diameter of 200 micrometer becomes to increase as the injection

pressure decreases.

In figure (c) Sauter mean diameters by the immersion sampling method

becomes larger again than those by the freezing method. As the injection

pressure decreases the dispersion of spray becomes weak and the sizes of

droplets increase. These tendencies bring larger possibility of coalescence

of droplets when they are captured by immersion liquid. This fact shifts the

value of mean diameter near the spray center. In the periphery of spray the

mean diameters are shifted down nearly to the values of freezing method. In

this region the increase of mean diameter due to the evaporation of fine

droplets becomes small, the number of fine droplets decreases under the lower

injection pressure.

Comparison of Size Distribution

In Fig. 5 number distributions of droplets obtained by the freezing method are

compared with those by the immersion sampling method. Difference is seen in

the region of fine droplet under the condition of high injection pressure.

Influence of Injection Period on Sauter Mean Diameter

Injected liquid quantity was changed by changing the injection period to

reveal their influence on Sauter mean diameter, and the results are shown in

Fig. 6. As seen in Fig. 6(a) the mean diameter by the freezing method with

injection periods of 30 and 40 millisecond is larger than those with 90 and

150 millisecond. This may be caused by the relatively longer transient period

641
200

N 150 8

01E

o

glOO

0)E

L.V

Z3O

W

D

50

PI=)2MPo

f=90msec

Y=2000mm

Woler

o Freezing method

A Immersion

sompling method

1 1

1

200

0 100 200

DIsfonce (rom nozzle axis r mm

(a) PI = 12 MI'a

300

|.50

100

u

dj +-

«>E

a

'ac

o

«>E

a> +-a

ina _ 8 50

_J

Pi=8MPo

f =90msec

Y=2000mm

Woter

o Freezing method

A Immersion

sompling method

JL

0 100 200

Distance from nozzle axis r mm

(b) PI = 8 MPa

300

200

Fig. A Comparison of local SMD byfreezing method with that

by immersion method E

Nl50 8

100

o> +-

0)E

D

TJC

o

4)E

L.

«J +-3

O

V)

^ 50

Pi=6MPo

t=90msec

Y=2000mm

Woter

o Freezing method

A Immersion

sompling method

o

0 too 200 300

DIstonce from nozzle oxis r mm

(c) PI - 6 MPa

642
of nozzle needle motion in the case of shorter injection period where

insufficient atomization occurs. In the immersion sampling method these

differences are not recognized. Based on this result, it can be said that

the freezing method is more sensible than the immersion sampling method.

Pl=12MPo

t=90msec

Y=2000mm

Woter

—O— Freezing method

--£ic— Immersion

sampling method

ot the center of spray

300 Diameter d400

80

60

c

u

c

£ 20

rl-blvlra t=90msec

Y=2000mm

Woter

—o— Freezing method

-A- Immersion

sampling method

-

at the center ot spray

0 100 200 ^ ^00 400 Diameter d /im

(a) Pi = 6 MPa (b) Pi = 12 MPa

Fig. 5 Distribution of droplet size

200r

E

~ 1501-

0}E

_o

-oiiOOh

03E

O

3

O

W

o

o

O S 5

A

50-

Freezing method

Pi=12MPo

Y=2000mm

Water

t=30msec

A 40

o 90

V 150

V O

<37

_i_

200

E

5-150

E

§100

03E

o

^ 50

V

A

D

a>o

Immersion

sampling method

Pi=12MPo

Y=2000mm

Water

o t=30msec

A 40

O 90

V 150

2 e

0 100 200

Distance from nozzle axis r mm

300 0 100 200

Distance from nozzle axis r mm

300

(a) Freezing method (b) Immersion method

Fig. 6 Influence of injection period on local SMD

643
CONCLUSIONS

A new apparatus for accurate measurement of spray droplet sizes based on the

liquid nitrogen freezing method has been constructed after the constructions

of seven trial setups. Using this apparatus comparison tests with the

immersion sampling method were carried out and the results are summerized as:

In the present apparatus improvements were made on some points for more

reliable measurement, easy operation and nitrogen economy.

The immersion sampling method presents larger mean droplet diameter

mainly due to disappearing of fine droplets by evaporation and coalescence of

droplets occurs when they are captured by immersion liquid.

In the freezing method the evaporation and coalescence of droplets are

almost avoided and, therefore, the method can afford highly accurate spray

diameter.

Since the freezing method is based on a quite simple principle and has

little possibility of error, the present setup is considered to be a standard

measuring apparatus.

ACKNOWLEDGMENTS

The authors wish to thank Mr . Hisao Nakamura for his enthusiastic helps

for constructions and tests of the apparatus.

This work was partly supported by the Tanikawa Foundation, Osaka, Japan.

REFERENCES

1 Longwell J P, Doctorial dissertation, Massachusetts Institute of Technology, 1943.

2 Taylor E H and Harman D B, 'Measuring Drop Sizes in Sprays', Industrial and

Engineering Chemistry, Vol.46, No. 7, pp. 1455-1457 , July 1954.

3 Tanasawa Y, 'Automobile Engines, Lecture', Toyota Gijutu, Vol.5, No. 7,

pp. 24-34, 1952.

4 Joyce J R, 'The Atomization of Liquid Fuels for Combustion', Inst. Fuels

J., Vol.22, No. 124, pp. 150-156, 1949.

5 Nelson P A and Stevens W F, 'Size Distribution of Droplets from Centrifugal

Spray Nozzles', J. Am. Inst. Chem. Eng., Vol.7, No.l, pp. 80-86, Mar. 1961.

6 Street P J and Danaford V E J, 'A Technique for Determining Droplet Size

Distributions Using Liquid Nitrogen', J. Inst. Petroleum, Vol.54, No. 536,

pp. 241-242, Aug. 1968.

7 Rao K V L, 'Liquid Nitrogen Cooled Sampling Probe for the Measurement of

Spray Drop Size Distribution in Moving Liquid-Air Sprays', Proc. 1st

ICLASS, Tokyo, pp. 293-300, Aug. 1978.

8 Karasawa T, Kaeriyama H and Kurabayashi T, 'Freezing Method by Liquid

Nitrogen for Measuring Spray Droplet Sizes', J. of the Fuel Society of

Japan, Vol.62, No. 679, pp. 901-912, Nov. 1983.

644
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 71

LDV AND PDA ANALYSIS OF SPRAY-WALL INTERACTION

G. Bninello*, S. Callera', A. Coghe^ G.E. Cossali^ and F. Gamma*

*CNPM-CNR, Peschiera Borromeo (MI), Italy

^Politecnico di Milano, Milano, Italy

*Universitk "La Sapienza", Roma, Italy

ABSTRACT

A diesel spray impacting on a wall in a close chamber, under

ambient air conditions, has been analysed by means of Laser

Doppler Velocimetry (LDV) and Phase Doppler Anemometry (PDA) in

order to measure droplet size and velocity in regions close to

the wall. Spray behaviour has been characterized in terms of

different injection periods and different spatial regions, and

proper data analysis allowed to identify the main features of

the droplet-wall interaction phenomenon.

INTRODUCTION

Recent theoretical predictions, experimental observations and

engine tests [1] evidenced that the interaction of a diesel

sprays with the piston wall might be beneficial to engine

combustion because of the influence on fuel vaporization and

heterogeneous mixture formation processes. Many experiments have

been performed in the past by different groups [2,3,4,5] to

investigate the spray-wall interaction, generally using highspeed photography, and the first submodel developed to describe

the spray-wall interaction process has been presented very

recently [6]. Spray-wall interaction characteristics seem to be

controlled mainly by the drop Weber number although other

parameters as wall temperature, fuel boiling temperature, drop

impinging angle, swirl intensity, air density and temperature,

Leidenfrost temperature, wall surface roughness have also been

proven to be important. Analysis of spray-wall interaction is

furthermore complicated, due to the interaction of spray

droplets with entrained air flow [7], which makes unrealistic to

extrapolate data obtained in experiments with single droplets.

Development and validation of a spray impingement model reguires

a large amount of experimental data on the droplets of a real

spray, like size and velocity in the region close to the wall.

The present research work was aimed to a detailed analysis of

the spray wall impingement process, by using the Phase Doppler

Anemometer (PDA) to obtain simultaneous measurements of spray

drop size and velocity in the close proximity of the wall.

645
EXPERIMENTAL SET-UP

The experiment was performed in a closed cylindrical chamber (89

mm i.d.,300 mm height) equipped with two glass windows

positioned to allow the use of the PDA in the 70 deg forward

scattering configuration to reduce the effective probe volume

dimensions ( about Imra in length and o.l5 mm diameter) and the

sensitivity of the phase shift to refractive index of the drops

[9]. A standard diesel injector (Bosch single hole nozzle, 0.25

mm diameter) injecting at 45 deg with respect to the nozzle

axis, was mounted off-axis on the top wall of the chamber. The

injector was fed by a Bosch high-pressure pump driven at 500 rpm

by an electrical motor and diesel oil was used as fuel. The

impact wall, an aluminium disc (50 mm diameter) with a polished

surface, was centred on the chamber axis and positioned at a

distance of 20 mm from the nozzle tip. (see fig.l). The

impingement distance (L) and the impingement angle (45 deg) were

kept constant. The nozzle opening pressure was fixed at 20 MPa

and the injection pressure history is shown in Fig. 2. Operating

experience indicated that the injection system provides very

reproducible pressure time histories. The ambient gas was air at

room temperature and pressure. The operating conditions of the

experiment are summarized in Table 1.

Table 1 - Operating conditions

Nozzle diameter, d (mm) 0.25

Max. fuel in j . pressure (MPa) 70

Pump speed (RPM) 500

Volume of fuel injected (mm'/inj) 36

Duration of injection (ms) 3.4

Fig.l Schematic of the impingement Fig. 2 The liquid inexperiment (all dimensions in mm). jection pressure.

646
Drop size and velocity were measured by the Phase Doppler

Anemometer (Dantec PDA); the light source was an Argon-ion laser

(5 W nominal power) and the transmitting optics comprised a 40

MHz Bragg cell for frequency shifting and directional

sensitivity and a front lens of 300 mm focal length. The

receiving optics was the 57X10 Dantec unit equipped with three

preamplified photomultipliers . The covariance processor of the

PDA instrument was capable of 1 [is time resolution in data

acquisition and the data transfer rate, to a PC, was higher than

100 kHz. The frequency bandwidth (36 MHz) and the optical

configuration of the transmitting optics defined a measurable

velocity range from -25 to 125 m/s and a measurable size range

from 2 to 80|im which resulted the most appropriate in the

present case. Normal and tangential velocity components were

measured at the same point at different times and drop size was

measured together with the velocity component normal to the flat

plate.

Measurements were performed at several locations on three lines

at distances of 0.5, 1 and 2 mm from the surface (see Fig.l),

laying on the symmetry plane normal to the wall. The

investigated region extended from the impingement area to the

boundary of the impact disc, in the downstream direction. Data

acquisition, triggered by the needle lift opening at 20 MPa,

covered a period of 5 ms for each injection and 5,000 to 10,000

data were collected over many consecutive injections for each

measurement point. An encoder pulse with 1 \is resolution allowed

a subsequent reconstruction of the size and velocity time

history of each individual injection and also the application of

time-ensemble averaging procedures.

RESULTS AND DISCUSSION

A typical result for the drop size and two velocity components,

measured downstream the nominal spray impinging location, is

shown in Fig. 3, where all the individual measurements collected

over a large number of injections are reported versus time, the

origin, t = 0, being defined by the 20 MPa triggering level on

the injection pressure trace. As in free sprays [8] three main

periods are distinguishable: the first period with strong

unsteadiness due to the passage of the spray head, a main period

where most of the fuel is injected, and a late period

characterized by the trailing edge, where the smaller droplets,

left behind by the passage of the previous structures, and the

bigger ones, generated during the late injection at low

pressure, are travelling together at low velocity. Close to the

surface (h=0.5 mm in Fig. 3 ), after the tip arrival, tangential

velocity stabilizes, with high rms, around a value larger than

the tip velocity and than decays after completion of injection.

The corresponding normal (i.e. perpendicular to the wall)

velocity component shows that only droplets traversing the

control volume with negative velocities (i.e. back from the

wall) were detected (Fig. 3 ). The size distribution shows an

upper bound of approximately 30 [xm and few data at higher

values. At the same location (X = 6 mm) and larger distance from

647
the wall (h=2 mm) the tangential component (Fig. 4 ) shows a

transient behaviour in correspondence of the early and late

injection periods, while the velocity of a large fraction of

droplets is almost zero during the main injection period,

suggesting that the wall spray does not extend beyond 2 mm from

the wall. At the same point, it appears clearly the coexistence

of droplets traversing the control volume with positive and

negative normal velocity components (fig. 4 ). Comparison with

Fig. 3 shows that the droplet sizes appear lower at h = 2 mm,

d^uring the main injection period.

The relative steadyness of normal and tangential velocity

components during the main injection period, although the

variability is high, suggested the following data analysis:

tangential velocities were ensemble averaged over a period from

1.5 to 2.5 ms after injection start (AIS) and the results

plotted versus distance from the wall, at X = 14 mm, and the

same was done with the size data. The results (Figs. 5 and 6)

show that the maximum mean velocity and mean diameter were

reached both at the closest distance from the wall with a steep

gradient, suggesting that the most of the tangential droplet

momentum remains concentrated close to wall during the main

injection period.

80.ee

6e.ee

4e.ee

ze.ea

e.ee

2e.ee

le.ee

e.ee

-ie.ee

-2e.ee

H 0 . 5 mm

e leee 2eee aeee norm. voXoc:±ty Cm/s>

4080 seee

r~. .u

'' :v:WMl

-J,' ;.-- :. '.-arr^.—

-^I>iaH«t«p Cum] H — 2 mm

20.00

10.00

0.00

-10.00

-20.00

0 1000 2000 3eee

1^ noirm. \7-oloc;±ty Cm/s)

4ee0 seee

2000 3oee 4eee seee

leee 2eee aeee 4eee seee tang

120.00 .

100. eo

80. oe

60.00

40.08

20.ee

0.08

-2o.ee

voloc±-ty Cm/s)

2088 3088 4088 5008

120.00 -I tang, velocity Cm/s)

100.00 4

30.00

60.00 .

40.00 J

20.00

0.00

-20.00

I

Ilaee 2000 30O8 4008

t±m9 C/^s>

5000

Fig. 3 Time evolution of drop

size and velocity at X = 6 ram.

Fig. 4 Time evolution of drop

size and velocity at X = 6 mm.

648
DV0.1 (um) Reticle #246

39 40 41 42 43 44 45 46 47 48 49 50 51

DV0.5 (um) Reticle #246

Figure 1. Plot of the final ILS test results for Dyo.i and Dyo.s- The symbols indicate the mean of the data for the

four repeated measurements for each reticle/lens combination in each lab. The error bars indicate one standard

deviation of these four test results. The dashed lines indicate the median of the data for the three different lenses,

and the ellipticity of the data for each lens are shown in the legend.

617
50 55 60 65 70 75 80 85 90 95 100 105 110

DV0.9 (um) Reticle #246

0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 1.5 1.6

Span (V) Reticle #246

Figure 2. Plot of the final ILS test results for D^q 9 and relative span (volume basis). The symbols indicate the

mean of the data for the four repeated measurements for each reticle/lens combination in each lab. The error bars

indicate one standard deviation of these four test results. The dashed lines indicate the median of the data for the

three different lenses, and the ellipticity of the data for each lens are shown in the legend.

618
Table II. Image-analysis Reference Value for Reticles used in the ILS.

Test

Result

Reticle

Serial*

Reference

Population

Value (pm)

Parsed

Number

Value (|jjn)

Volume Equiv.

Sphere Multiplet

Value (|im)

Area

Equivalent

Value {\m)

Image-analysis

Reference

Value^ (jxm)

Dvo.l

Dyo.l

246

247

25.74

25.82

26.16

26.24

26.39

26.46

29.63

29.60

27.0 ± 2.3

27.0 ± 2.4

Dvo.5

246

247

47.54

47.26

47.95

47.66

49.80

49.41

51.38

51.26

49.2 ± 2.3

48.9 ±2.7

Dvo.9

Dvo.9

246

247

69.15

69.33

69.40

69.57

72.78

72.77

78.52

78.75

72.5 ± 5.1

72.6 ±6.3

^The image-analysis reference value is taken as the mean of the four estimates. The uncertainty limits for the

image-analysis reference values represent 95% confidence intervals based on estimates of precision and bias error

in assigning the values.

Table in. Precision Statistics for the Test Method (all focal lengths grouped).

Test

Result

Reticle

Serial*

Mean of

ILS Results

X(nm)

Within-lab

Repeatability

Limit r^ (|im)

Between-lab

Reproducibility

Limit r1 dm)

Rel. Within-lab

RepeatabiUty

Limit r/X(%)

Rel. Between-lab

Reproducibility

Limit R/X(%)

Dvo.l

IVo.l

246

247

25.09

25.04

±1.25

±0.91

±5.18

±5.37

±5.0

±3.7

±20.7

±21.5

Dvo.5

246

247

46.05

46.05

±0.56

±0.59

±4.26

±4.57

±1.2

±1.3

±9.3

±10.0

DvO.9

Dvo.9

246

247

77.47

77.33

±3.95

±6.80

±26.50

± 29.18

±5.1

±8.8

±34.2

±37.7

^The precision indices r and R, defined in E691-87, are 2.83 times the repeatability (within-lab) and

reproducibility (between-lab) standard deviations, respectively.

Table IV. Bias of the Test Method (all focal lengths grouped).

ILS Results Image-analysis Bias Estimates

Test

Result

Reticle

Serial*

Mean Reproducibility

Value (|am) Limit R(|im)

Reference Uncertainty Measured Statistically

Value dim) Limits ^(nm) Bias(jLim) Significant?

%0.1

%0.1

246

247

25.09

25.04

±5.18

±5.37

27.0 ± 2.3

27.0 ±2.4

-1.9

-2,0

No

No

246

247

46.05

46.04

±4.26

±4.57

49.2 ±2.3

48.9 ± 2.7

-3.1

-2.8

No

No

DVO.9

%0.9

246

247

77.47

77.33

±26.50

±29.18

72.5 ± 5.1

72.6 ±6.3

5.0

4.7

No

No

iThe uncertainty limits for the image-analysis reference values represent 95% confidence intervals based on

estimates of precision and bias error in assigning the values. The statistical significance indication also is for 95%

confidence.

619

ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 68

CORRECTING MALVERN PARTICLE SIZE MEASUREMENTS

FOR PHASE DISTORTION

R.A. Pietsch, P.E. Sojka and G.B. King

Thermal Sciences and Propulsion Center

School of Mechanical Engineering

Purdue University

West Lafayette, IN, U.S.A.

1 ABSTRACT

The two effects of phase distortion on particle size measurements made with a Malvem

Spray Analyzer are identified as beam spread and a shift of the Airy pattern extrema to higher

spatial frequencies. Beam spread is shown to be present regardless of the laser beam coherence

length (4), while the Airy pattern shift occurs only when the size of the particle being sampled

exceeds 4. A correction scheme that removes the effects of phase distortion when possible is

outlined, and then evaluated by comparing phase distorted particle size information with its

undistorted counterpart. The enhanced agreement indicates the proposed technique successfully

removes the effects of phase distortion when possible. Finally, guidelines are provided for determining when phase distortion is present.

2 INTRODUCTION

The quality of atomization cannot be ignored when considering the performance of any liquid fueled combustion system. Information about both the mean particle size and distribution of

particle sizes is important to the combustion engineer in his search for more efficient combustor

designs and alternative fuels. A popular optical method of collecting this size information utilizes Fraunhofer diffraction of a laser beam as it passes through a spray. Currently, measurement

systems based on Fraunhofer diffraction are commercially available; the Malvem Spray

Analyzer is one such system. Unfortunately, recent studies of particle sizing in high temperature

turbulent environments have shown that this type of instrument, in its current form, is incapable

of accurately measuring the scattered light profile and yielding the desired particle size data

[1,2]. The inaccuracies noted in References 1 and 2 result from optical perturbations present in

the system being studied.

The primary perturbing optical effects associated with laser beam propagation in a high

temperature turbulent environment are known as beam wander and phase distortion. Beam wander occurs when the laser beam encounters an index of refraction gradient and is displaced, or

steered, from its normal path of propagation. Phase distortion occurs when different points on

the same laser beam wavefront vary in phase after passing through an index of refraction gradient.

The effects of beam wander were studied in detail by Miles et al [2] and a technique was

developed to successfully compensate for this phenomenon. Miles et al also suggested that any

errors remaining in the particle distribution measurements after correcting for beam wander were

due to phase distortion. Support for their hypothesis was provided by performing a Young's

double slit experiment at various temperatures to measure Michelson's visibility, arguing that

Michelson's visibility was equivalent to the complex degree of coherence between two points on

the laser beam, and then noting that the observed decrease in coherence length was indicative of

an increase in the level of phase distortion.

Leader [3] studied phase distortion, concentrating on how it affects laser beam propagation

through weakly and strongly turbulent atmospheres. He discovered that as the turbulence

621
became stronger, phase distortion's effects were dominant over those of beam wander. Leader

also noted that as phase distortion increased, the ability to focus the beam to a given spot size

decreased. Additionally, the unfocused beam's diameter, known as the beam spread, was seen to

increase log-normally as phase distortion increased.

This increase in beam diameter is an important issue since some of the incident beam's

energy will emerge from a spray as unscattered light. In many diffraction based instruments, this

unscattered energy is removed from the particle diffraction profile by measuring the background

light intensity incident on the photodiode. Leader's work therefor suggests that a background

light intensity reading for a phase distorted beam will have a greater intensity at lower spatial

frequencies than for an undistorted beam.

Two conclusions can be drawn from the work of Miles et al and Leader. First, the effects

of phase distortion become more pronounced as the turbulence within a system increases. Second, since the amount of beam spread, and thus additional background intensity, is dependent on

the flow conditions, a separate background intensity reading should be taken at each temperature

or flow velocity being investigated in order to eliminate this effect of phase distortion.

Two issues, however, remain to be addressed. First, how is a researcher to know when the

effects of phase distortion are important, and second, are tiiere any further effects of phase distortion that must be eliminated or compensated for.

The goal of the this work then is twofold: to develop a technique that eliminates the errors resulting from phase distortion whenever possible and to outline a procedure that will aid users

in determining when the effects of phase distortion are present. The next section provides a summary of the theory needed to accomplish these objectives.

3 THEORY

In order to develop a correction scheme to remove the effects of phase distortion on optical

particle sizing with the Malvern instrument, it is first necessary to determine if phase distortion is

present and, if so, to describe its impact. Phase distortion is present if the medium of interest

reduces the spatial coherence length of a laser beam passing through it. Its effect on Malvern

particle size measurements depends on the relative size of the perturbed laser beam's spatial

coherence length and the diameter of the largest particle to be sampled.

A level of phase distortion such that the spatial coherence length of a beam passing

through the medium is less than the diameter of the largest particle in a spray indicates that all

particles are not illuminated coherently. Fraunhofer diffraction by apertures illuminated using

partially incoherent light was studied by Shore et al [4] who showed that a gradual loss in the

intensity contrast of the diffraction signature occurs as the coherence length decreases. Furthermore, when the coherence length approaches zero, the diffraction maxima and minima shift to

higher spatial frequencies. This indicates that more light would be incident on the Malvern

detector array at larger radii, thus skewing the particle size distribution to smaller diameters.

Phase distortion on this level can be detected by employing a calibrated aperture. Specifically, a pinhole may be introduced into the phase distorting media and its intensity profile

obtained using the Malvern receiver. If the diffraction maxima and minima of the resultant Airy

pattern are shifted to higher spatial frequencies, the particle size measurements cannot be corrected. Consequently, a series of calibrated apertures is inserted into the flow with the smallest

pinhole exhibiting no shift in its diffraction extrema, for a given flow condition, indicating the

largest particle that can be accurately sized. A correction scheme was developed for this case. When no shift in the position of the diffraction extrema occurs for any particle, the complex degree of coherence, y, is used to describe the optical perturbations. By examining the scattering geometry illustrated in Figure 1, and following the notation from Hecht and Zajac [5], the

complex degree of coherence may be derived in the following fashion.

Consider two identical apertures or particles as the source of the diffracted radiation. If

they are located at a distant plane along the path of propagation, the electromagnetic fields can

be combined as: E,{t) = K,E,{t-t,) + K^E^{t-t^) (1)

622
where ti = r^lc and ?2 = ^-Jc . The K factors depend on the size and location of the scatterers relative to the point P and also account for absorption, diffraction and other physical effects. Defining X = ^2 - ?i allows the time origin to be shifted and the K factors to be removed with the result

that the mutual coherence function can be defined as r,2(t)=<^,(T-Hr)£;(0> (2)

Normalizing the mutual coherence function by the magnitude of the two sources gives the complex degree of coherence

<£i(T-Kr)E;(f)>

The magnitude of YiaCO is between 0 and 1. If it is equal to unity the field at P is the result of two

coherent waves out of phase by some quantifiable amount while if the magnitude is zero then the

two waves are said to be incoherent. Any value of Yi2(?) between these two extremes defines

partial coherence.

Thompson [6] studied diffraction and interference from multiple apertures and arrays by

looking at the generalized interference law: = i i {IJS^ 1 7^(0)1 cos({3^(0) 5^) (4)

/I = 1 OT = 1 where /„ and are the intensities produced by each beam individually, | y^l is the modulus of

the complex degree of coherence between apertures n and m, and

(3^(A) = 27i^A + flr^(Y^(A)) (5)

5^ = (27C/X)(r^-rJ (6)

Here and r„ are the path lengths from the two source points to the observation point, A is the

path difference divided by the speed of light, v is the mean frequency, and X, is the mean wavelength of the laser beam. It is important to note that

17^(0)1=1, n=m

lY„.(0)l =17^(0)1

For a circular source illuminating a regular array of circular apertures, only nearest neighbor apertures along the array directions can be arranged to be incoherent with respect to each

other. That is, for a regular array, 7„.„ + i(0) = 0 is a possible occurrence. However, because the

spacing of the Airy ring minima is not constant, some degree of coherence between an aperture

and its non-nearest neighbors will always exist. This implies that unless the coherence length is

much less than the spacing of the array, no difference will be seen in the resulting diffraction

pattern.

This observation is important to the operation of particle sizing instruments based on forward light scattering since it suggests that if phase distortion produces a beam whose coherence

length is greater than the diameter of the largest particle to be sampled then the shape of the

diffraction profile will be the same as for the unperturbed case. Thompson also investigated diffraction from irregular arrays using lycopodium powder.

Of specific interest were the diffraction profiles that resulted from using illumination whose

coherence length ranged from several thousand to only a few particle diameters. The results

showed no loss in fringe contrast with a drop in coherence length, as perhaps might have been

predicted from the work of Shore et al. The only difference was a "smoothing" of the intensity

profiles as the coherence length decreased since the values of 7^ contribute less to the sum in

Eqn (4).

In summary, the level of phase distortion can be determined by sampling the Airy patterns

produced by a series of calibrated apertures. The minimum size aperture yielding a pattern not

623
shifted to higher spatial frequencies must be greater than the largest particle to be sized or the

measurements cannot be corrected. If this condition holds, the work of Thompson shows that the

phase distorted diffraction profile will not differ from the undistorted profile. Furthermore,

under these conditions the only effect of phase distortion that need concern the user of a Fraunhofer diffraction based particle sizing instrument is beam spread. Fortunately, this spread, which

will appear as undiffracted light in the energy profile, can be removed by acquiring the

instrument's background signal under the conditions that lead to phase distortion and then processing the particle scattering data using the standard software. A demonstration of this technique is provided in Section 5.

4 EXPERIMENTAL APPARATUS

A Malvern 2600HSD diffraction based particle sizing instrument was employed in this

investigation since the Malvern serves as a sort of default standard in diffraction based systems

and Malvems have been used extensively in industry and research applications by many individuals and groups. A high temperature flow tunnel was used to generate the desired optical perturbations. It has been described in detail by Miles [7].

In order to isolate the effects of phase distortion on the particle diffraction profile, it was

first necessary to remove the effects of beam wander. This was accomplished by utilizing the

conditional sampling hardware developed by Miles et al [2]. This hardware was added to the

optical train developed to study phase distortion and is denoted by the extension ".BW" in Figure

2. It includes a 4% reflection beam splitter, a focusing lens (250 mm focal length), a fourquadrant photodetector, and electronic circuitry used to compare the laser beam's wandered position to that of the ambient situation. If the wandered beam's position is within an acceptable

window, the conditional circuitry triggers the Malvem receiver to collect the diffraction profile

information.

In order to ensure that a repeatable particle size distribution could be obtained throughout

the course of the experiments, a standard reticle was employed [8]. It had best-fit RosinRammler parameters of N = 3.20, X = 52.8 |im, and SMD (Sauter mean diameter) = 40.27 )im.

Several issues had to be considered when observing the effects of phase distortion on the

unscattered laser beam, on the diffraction pattern from the reticle, and on the diffraction pattern

from the calibrated apertures. The ability to reproduce the same diffraction pattem from either a

calibrated aperture or the reticle was limited by the ability to reproduce their exact physical location within the optical train. The repeatability of such a task was determined through trial and

error to be low, regardless of the positioning method employed. In order to observe each of the

above profiles individually, without having to reposition the reticle or calibrated aperture, the

experimental apparatus shown in Figure 2 was utilized.

The optical hardware illustrated in Figure 2 consists of a 10 mW HeNe laser whose beam

is spatially filtered and collimated to a diameter of 4.5 mm, plus two Mach-Zehnder interferometers in series that sample and redirect the laser beam and the diffraction profiles resulting from

the reticle and calibrated aperture. As shown in Figure 2, the calibration aperture and reticle are permanently positioned in separate arms of the interferometers, thus guaranteeing that their diffraction profiles are reproducible. The interferometer arrangement also allows for observation of

the effects of phase distortion on the undiffracted laser beam by adjusting the iris diaphragms to

block off both the reticle and aperture. As a result, the user can observe each of the relevant

intensity profiles individually without having to reposition the reticle or calibrated aperture

between runs. Experimental data obtained using the apparatus presented in Figure 2 are introduced in the

next section and indicate the effects of phase distortion on the standard instrument's performance, plus the enhanced agreement that results fi^om correcting particle size data for phase distortion.

624
I

SLIT/APERTURE

SOURCES

DETECTION PLANE

Figure 1. Scattering geometry for two identical apertures.

QD.BW IR

HeNe LASER

<=» FL.BW

BS BS IR 0 SI I 0 ? I s

BS.BW PR IR CA BS

CSC.BW

C MR

PDM

IR DAQ

PDM - PHASE DISTORTION MEDIA

BS - BEAMSPLITTER

FL - FOCUSING LENS QD QUADRANT DETECTOR M - MIRROR

IR - IRIS DIAPHRAGM CA - CALIBRATED APERTURE

MR MALVERN RECEIVER DAQ DATA ACQUISITION CSC CONDITIONAL SAMPLING

CIRCUITRY .BW - BEAM WANDER HARDWARE

PR PARTICLE RETICLE Figure 2. Experimental apparatus.

RING NUUBEn

Figure 3. Fraunhofer diffraction profiles for a 600 \}jn diameter pinhole.

625
5 RESULTS

Data acquisition commenced by placing the calibrated aperture and reticle in their respective interferometer arms, centering the Malvern photodiode array on the laser beam, and aligning

the beam wander conditional sampling hardware. Three sets of ambient background, calibrated

aperture diffraction profile, and particle diffraction profile data were then collected by shutting

the appropriate iris diaphragms. At this point, the blower supplying air to the test section was

started and an air velocity of 25 m/s was established in the test section. The air temperature was

adjusted to 200, 400 and then 600 C with one background profile, three calibrated aperture diffraction profiles, and three particle diffraction profiles obtained at each temperature. The entire

process was performed once a day for three days for a total of nine data points at ambient and

each of the elevated temperatures.

As mentioned previously, the calibration aperture serves as a users' guide to recognize

when phase distortion is present and indicates the largest particle diameter that can be sized without phase distortion affecting the light intensity profile. Sample intensity profiles for a phase distorted and non-phase distorted aperture (circular pinhole, d = 600 [im) are shown in Figure 3.

The agreement between the two profiles indicates no shift in the maxima or minima and therefor a coherence length greater than 600 |im. Consequently, all particles up to the 564 [im limit measurable using the Malvern's 300 mm focal length lens will be coherently illuminated and the

only effect that need be considered is beam spread.

A sample phase distorted diffraction profile from the reticle appears in Figure 4. Compared to a non-phase distorted profile, the perturbed profile reveals large intensity variations at lower spatial frequencies, i.e. in the innermost photodiode rings, due to beam spread. However,

at larger spatial frequencies no intensity variation is observable supporting the conclusion that

the spatial coherence length was larger than the largest particle in the distribution. Figure 5

shows the background readings corresponding to each of the profiles in Figure 4. By subtracting

the background taken under phase distorting conditions from the phase distorted particle profile, a corrected diffraction profile is generated and used to determine the particle size distribution.

An example profile appears in Figure 6; comparing it to the light intensity profile of Figure 4

clearly demonstrates the improvement achieved when using the correction procedure outlined

above. This improvement is also apparent from the Malvern calculated Rosin-Rammler size distribution parameters for the uncortected and corrected profiles, as shown in Tables 1 and 2.

Uncertainties in these tables are reported as one standard deviation. Note that the discrepancies

between the ambient and 200, 400 and 600 C cortected values forX, N and SMD are within the

sum of their standard deviations, with the exception of the 600 C value for N. In contrast, the

discrepancies between the ambient and uncorrected 400 and 600 C values forX, N and SMD are

at best one standard deviation and most often greater than 2a.

6 SUMMARY

The two effects of phase distortion on the diffraction pattern produced by a particle, beam

spread and a shift of the Airy pattern extrema to higher spatial frequencies, have been noted.

Beam spread is present regardless of the laser beam coherence length (4), while the Airy pattern

shift occurs only when the size of the particle being sampled exceeds 4. A procedure to determine the coherence length under phase distorting conditions was described and a correction

scheme was developed to remove the effects of phase distortion when all particles have

diameters less than 4.

The correction scheme is straightforward, requiring only that the Malvern's background

signal be acquired under the same phase distorting conditions as the particle diffraction profile.

The data may then be processed using the standard Malvem software.

The correction technique was evaluated by comparing phase distorted diffraction profiles

and background intensities with their undistorted counterparts that were collected under identical

flow conditions. Results showed that phase distorted intensity profiles yielded particle distribution parameters that were as much as 9% in error. Parameters calculated from corrected intensity

626
0 S 10 15 20 2S 30

RING NUMBER

Figure 4. Fraunhofer diffraction profiles for the reticle.

1 a 5 7 8 11 13 IS 17 II 21 23 25 27 2B 31 2 4 S I 10 12 14 IB II 20 22 24 28 21 30

RINGNUWBER

Figure 5. Ambient and high temperature backgrounds.

T-400C

gl

0 5 10 IS 20 25 30

RING NUMBER

Figure 6. Fraunhofer diffraction profile corrected for phase distortion.

627
profiles, however, were within 2 % of those obtained under ambient conditions. This indicates

the proposed technique successfully removes the effects of phase distortion when possible.

We thank E.D. Hirleman (Arizona State University) for several helpful discussions and

B.H. Miles (US Army Waterways Experiment Station) for providing comments on an early ver- sion of this manuscript.

REFERENCES

1. Roesler, T.C., Lefebvre, A.H. and Sojka, P.E,, "Fuel Atomization at High Ambient Temperatures," Central States Section/Combustion Institute paper 86-6D1, NASA-LeRC, Cleveland, OH

(1986).

2. Miles, B.H., Sojka, P.E. and King, G.B., "Malvern Particle Size Measurements in Media with

Time Varying Index of Refraction Gradients," Applied Optics, Vol. 29, No. 31, pp. 4563-4573

(1990).

3. Leader, J.C., "Beam Properties of Partially Coherent Curved Beam Waves in the Turbulent

Atmosphere," Journal Optical Society America Vol. 70, pp. 682-688 (1980).

4. Shore, A., Thompson, B.J. and Wliitney, R.E., "Diffi-action by Apertures Illuminated with

Partially Coherent Light," Journal Optical Society America, Vol. 56, 733 (1966).

5. Hecht, E. and Zajac, A., Optics, Addison-Wesley, Reading MA, (1979).

6. Thompson, B.J., "Multiple-Beam Interference with Partially Coherent Light," Journal Optical

Society America, Vol. 56, pp. 1157-1161 (1966).

7. Miles, B.H., "Optical Particle Sizing in a High Temperature Environment," M.S. Thesis, Purdue University (1988).

8. Hirleman, E.D., "On Line Calibration Techniques for Laser Diffraction Particle Sizing

Instruments," ASME Paper No. 83-GT-232.

Table I. Particle Size Distribution Parameters for Uncorrected Data

Temperature, C X, jim N SMD, \im

20 52.34±0.18 3.40±0.03 41.97±0.04

200 52.86±0.07 3.39+0.00 42.3310.57

400 54.99±0.73 3.24±0.06 43.4710.33

600 56.25±0.32 3. 17m17 44.15±0.48

Table n. Particle Size Distribution Parameters for Corrected Data

Temperature, C X,\im N SMD, [im

20 52.34±0.18 3.40±0.03 41.97±0.04

200 52.45±0.00 3.39±0.00 42.00+0.00

400 52.0310.70 3.3510.07 41.5310.49

600 51.8511.05 3.5510.08 42.0310.61

628
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 69

A COAXIAL-BEAM DROP SIZING INSTRUMENT

FOR DENSE SPRAYS

W.M. Grissom

Physics Department

Morehouse College

Atlanta, GA, U.S.A.

ABSTRACT

A low-cost instrument has been developed to measure drop size and speed in dense sprays. The

instrument uses two coaxial laser beams of different colors. The smaller locater beam defines the center of the

larger metin beeun and triggers a measurement only when a drop passes through the center of the main beam.

Light scattered perpendicular to the beam is focused through a double sUt imaging system which avoids partial

obscuration of the image. A vibrating glass tube produces a uniform drop stream for calibration. A simple

"scattering tomography" analysis accounts for beam attenuation. Fluctuations in beam intensity are also

considered. Data in a spray nozzle are reported. Both the diameter and speed histograms are best fit by a log-normal function.

INTRODUCTION

A review of most drop sizing techniques up to 1978 is given by Azzopardi [1]. Since then the

Polarization Ratio [2,3], Phase Doppler [4], and Pulse Intensity [5-7] techniques have been developed. The

present instrument is of the latter type.

The Pulse Intensity techniques determine the drop size by measuring the scattered pulse as a drop

crosses a laser bejun. This gives a robust signal with simple optics. The main problem is that the scattered

intensity depends upon the path the drop follows across the beam. Since the laser beam normally has a Gaussian intensity distribution, a drop passing far from the beam center gives a smaller pulse height than if it passed through the center. This problem is termed "Gaussian ambiguity".

An obvious solution to the Gaussian ambiguity problem is to generate a beam with a uniform ("tophat")

profile, but this has proven difficult [8]. Instead, several researchers [5,6] developed deconvolution algorithms

to analytically correct for the ambiguity. Another obvious solution is to accept a Gaussian beam, but define the

center of the beam such that measurements are triggered only when a drop passes through the center. Hess [7] used a laser velocimeter to define the beam center. Since the velocimeter signal is weak,

forward scattered light was collected by viewing the beam at an oblique angle, greatly increasing the probe

volume size. Since the velocimeter probe volume must be about 4 times wider than the drop to give a good

visibility signal and the main beam must be about 5 times wider than the velocimeter probe volume, the

minimum average drop spacing is about 20 drop diameters, which is fairly sparse.

629
EXPERIMENTAL APPARATUS

The optical cirrangement Is shown in Figure 1. A small yellow beam (He-Ne, 594 nm) defines the center

of a coaxial red beam (He-Ne, 633 nm). The red main beam is about 5 times larger than the yellow locater

beam at the probe volume. A similar coaxial arrangement was reported by Wcmg and Hencken [9]. Light

scattered at 90° is imaged onto the main slit. The slit width, together with the beam diameter defines the size

of the probe volume. The two colors are separated by an edge filter. The red light, from the main beam, is reflected into a photomultipher. The yellow light, from the locater beam, is re-imaged onto the "locater sUt". The locater slit views only the center of the main slit, ignoring drops which are partially obscured by the

edges of the main slit. The locater photomultipher triggers a measurement only when a drop passes through

the center of the main beam and is centered in the view of the medn sht. Umhauer [10] reported a similar two

slit imaging system, with the slits in parallel.

A locater pulse triggers storage of a main photomultipher pulse with a Rapid Systems R2000 digital

O-scope. The digitized pulse is fit to a Gaussian curve by a logarithmic transform to a parabola and linear

regression. About 50 samples per pulse are needed to average over ripples on the pulse. The pulse peak

determines the drop diameter and the width determines the speed transverse to the beam. Checks on the

fitting error help eliminate drop coincidences in the probe volume. Since both size and speed are obtained,

either a flux or spaticd drop size distribution can be generated and speed-size correlations can be obtained.

Since the pulse amphtude is approximately proportional to the square of the drop diameter, care must

be taken to provide a wide dynamic range. Previous researchers used logarithmic ampHfiers. Instead, the

programmable gain feature of the digitizer is used. Measurements are made at three gain ranges. The

processing rate is usually slower them the data rate. To avoid bias against the busier ranges, a set of pulses is

digitized and the elapsed time recorded. The pulses are then processed and the histogram bins incremented

inverse to the digitization time.

The transmitted beam is focused onto a pinhole in front of a phototube. A 4 axis positioning system

controlled by a Precision Micro Control DCX-8 board allows translations in the scattering plane and rotations

about the spray and collection axes. A vacuum collects the spray drops, allowing operation with the spray

pointing upward.

CALIBRATION

A glass pipet is drawn to a small diameter and cleaved. A laminar jet of liquid is forced from the pipet

while vibrated by a speaker. At certain frequencies the jet breaks into a stream of uniform drops. Sakai [12]

gave a detciiled report on the phenomenon. A time trace is shown in Figure 2. The drops almost overlap in the

main beam. The ripples in the main pulses are due to noise in the beam profile. Reducing the beam diameter

0

1 Dcot-ei- beom

( Lja 1 I av ]

Fig. 1 Optical ornangemenL

c

Ol

4/2/91 mam

locater

J. A J

0 0.5 1 1.5 time (msec)

Fig. 2 Photomultiplier signals with uniform 230 umD drops,

700 umD main beam, 200 umD locater beam

630
yields smoother pulses. The spatial filter partially obscures the main beam, resulting in steeper edges than a

true Gaussian pulse, which is advantageous in decreasing overlap of adjacent pulses.

Knowing the frequency and flow rate, the drop diameter is easily calculated. The calculated size has

been verified using a microscope, with stroboscopic illumination. Photographing the drop stream and

measuring the drop spacing allows calculation of the speed. Comparing this with the e"^ pulse width

(temporal) allows Ccilculation of the e'^ beam width, which is the caUbration constant for speed.

Since the drops follow a very defined path through the beam, the drop stream must be scanned back

and forth across the beam to test the instrument's resolution. The resulting histograms are shown in Figure 3. Equally narrow histogrcuns are obtained for scans along the laser beam. Disabling the locater beam gives the

histograms shown in Figure 4, displaying the "Gaussian ambiguity". It appears difficult to deconvolve the

diameter histogram to appear as that in Figure 3, as suggested by Holve and Self [5] and Mizutani [6]. The

speed histogram is less affected. Theoretically, with a Gaussian beam, it should be unaffected [5]. However, in

practice the degradation is severe. Scanning the drop stream across the laser beam also allows determination of the size of the locater

probe volume, by noting the Umits over which the locater beam is triggered. Wang and Hencken [9] found that

large drops are more likely to trigger a measurement. However, testing with drops from 163 to 279 ^mD,

using a locater beam about 200 /imD and locater slit 500 fim wide, showed no difference in probe volume size.

Therefore no correction is made for size bias. The Ukely explanation is that most of the light scattered at 90°

is due to total internal reflection from a small region on the back of the drop, so that the effective scattering area of a drop is much smaller than the actucil cross section.

Calibration results are shown in Figure 5. Ideally, the drop diameter should be proportional to the

square root of the normalized scattered intensity, however a slight curvature is apparent as the drops increase

in size. An explanation results by averaging the beam intensity over the drop cross section, giving the fit shown

in Figure 5. The best fit occurs by assuming a main beam diameter (D) of 700 /imD, although the curvature is so slight that a much larger value would make Uttle difference. The beam diameter estimated by eye was

550 /imD, while the speed calibrations determined 578±63/imD. The difference is again attributed to the

effective scattering diameter being smaller than the actual diameter.

For drops smaller than about 10 /imD, the light collection cone is not large enough to average over the

Mie interference pattern and the cahbration curve becomes multivalued. However, few drops this small exist

in the sprays tested. Calibration streams smaller than 150 /imD can be generated, but the drops overlap in the

probe volume, so that the lower end of the calibration curve has not been verified. One solution is to decrease

the number density by electrostatically deflecting away every other drop [13].

Calibrations using 8 different Uquids suggest a linear relation between the apparent diameter increase

and the refractive index relative to water, as shown in Figure 6. The incident light was polarized perpendicular

(vertical) to the scattering plane. For most low molecular weight liquids the correction is fairly small.

60- 25-

50- 230 umO drops 700 umD main beam

200 umD locater beam

200 counts

4/2/91 15-

230 umO drops 700 umD main beam

200 umD locater beam

200 counts

4/2/91 C

C

3

20-

5-

10-

0 10 20 30 40 50 60 70 80 90 100

diameter (%range)

, li l iiiInU

0 10 20 30 40 50 60 70 80 90 100 speed (% range)

Fig. 3 a) Diameter histogram wilii drops scanned across beam Fig. 3 b) Speed histogram with drops scanned across beam

631
12a

diameter (%range)

Fig. 4 a) Diameter histogram without locater beam

90-

speed (% range)

Fig. 4 b) Speed histogram without locater beam

Drop Diameter (d) (umD) refractive index

Fig. 5 Diameter calibration Fig. 6 Effect of refractive index on drop size , rel. to H20

By averaging over the beam area, a large drop gives a slightly wider pulse as it crosses the beam than a

small drop. The resulting pulse is still essentially Gaussian. Numerical integration predicts a fractional

increase in pulse width of 0.539(d/D)2 - ^ * ^ , where d is the effective scattering diameter. The present scatter

in the speed cahbrations (± 11%) preclude verifying this prediction.

BEAM ATTENUATION

In a dense spray the laser beam is attenuated in traveling to the probe volume and the scattered hght

suffers attenuation in exiting the spray. These attenuations cause the drop pulses to appear smaller than

normal and must be corrected.

The transmittance to the probe volume is found by passing the beam from the left and right, in turn:

^Jl ~ /^r)^]'^ > where tj^ and tj. are the transmittances for the beam from the left and right, respectively, and

T is the total transmittance. A linear "absorption" law (actually scattering) is assumed, i.e. T = t^ tj.

The pulse height histograms appear similar in each case, with the abscissas simply scaled by the ratio of

the beam intensities. The ratio t^/t^ is thus found by ratioing the peak values of the two histograms. The

collected Ught transmittance is the same fraction in each case, so that the ratio is unaffected. It is not actually

necessary to pass a beam from the right. The same geometry results by rotating the spray 180° about the

collection axis, i.e. so that the spray points upward.

This emalysis has general application to any problem where the attenuation is due partially to scattering,

and the "absorptivity" at every point in a cross section is desired, for example in studying soot generation. The

632
"absorptivity" at every point along a line through the region can be found by differentiating the transmittances

along the line. This analysis is best termed "scattering tomography".

The transmittance along the collection path can be found by rotating the spray 90° about the spray axis,

as shown in Figure 7, making the former collection path (t^) align with the beam, allowing its transmittance to be determined. Denote the peak of the measured pulse height histogram by Pj^^ , where k=JL or r for the

beam from the left or right, respectively, and ^ = 0° or 90° for the spray at the original orientation or rotated

90° about the vertical, respectively, tj denotes the transmissivity along the four paths shown and T^ and

are the transmittances along the two perpendicular paths across the spray. The corrected pulse height peak is found as: P = P^qo /(tj tg ). Expressing t^ and t^ with the scattering tomography analysis above eventually

results in an identity (note T J =^^^2 '^'^^^2 ~^3^4)*

P = rtit2t3t4 PPl ^

L Ti T2

Referring to Figure 7, the possibilities for the numerator are:

(h^3P)(^2U^) = Pi 0° Pi 180° ' PX 0° Pr270° ' Pr90° 180° ' o*" Pr90° Pr270°

(tit4P)(t2t3P) = Pjj PrO° , Pjd 90° ^Ji 270° ' Prl80° PrO° ' or Prl80° P^ 270°

These are summarized as two procedures: (1) after the first measurement, rotate the spray 180° about

the spray axis, or (2) after the first measurement, rotate the spray 90° about the spray axis and 180° about the

collection axis (to pass beam from right). The first procedure is more convenient since the spray remains

pointing downward. It is interesting that including the collection attenuation leads to a simpler procedure than

considering the input attenuation alone. Beretta [2] states the same result for the case of an axially symmetric

spray. As seen here, the result is not restricted to axial symmetry.

A related problem is that the beam attenuation is not constant, due to statistical fluctuations in the

number of drops along the path to the probe volume. The collected light also fluctuates, but it is not significant

due to the large collection cone. A drop scatters most of the light incident upon its cross-sectional area out of the beam. An equal

amount of light is diffracted, but most of it remains within the confines of the beam. Scattering causes

localized shadows in the beam profile, giving ripples on the pulses. However, the Gaussian fitting largely

negates the ripples, so that only temporal fluctuations in the spatially averaged intensity need be considered.

Assuming a uniform beam intensity and a mono-disperse spray, the fraction of beam energy scattered

by a single drop is (d/D) ^ , The total fraction attenuated is: A = l-exp[-N(d/D) ^

], where N is the number of

drops in the beam. According to Poisson statistics, the attenuated fraction fluctuates by an rms fraction A/J N,

so the rms beam intensity fluctuation is:

1^^5/1 = (d/D) A[-ln(l-A)]-^

For both very low and very high attenuations there is little fractional intensity fluctuation. The

maximum fluctuation occurs at an attenuation of 71.5%, in which case Ifj„5/I = 0.638(d/D). Given the square

root relation, the corresponding diameter fluctuations are half this. It is desirable to converge the main beam

as it approaches the probe volume, so that its average diameter through the spray is increased, decreasing the

fluctuations.

Hess [11] measured intensity fluctuations of +/-14%, at a beam attenuation of 10.7%. The beam

diameter was about 600 /imD at the probe volume, and somewhat larger in the spray. Since the attenuating

drops were far upstream, a factor of 2 should be added to account for diffraction. The analysis would then

predict an rms fluctuation of J 2(d/D)(0.318). The average drop size in the spray was not stated. To agree

with the above analysis, the drops would have to have been an average of 187 /im in diameter.

633
1

e>0 e«90

FlQ. 7. Input ond colloctlon ottonuoLlon

main beam

locater beam

DelervanWDB20

40 psi, 1=1 52mm,

on was, 4/4/91

Time (ms)

Fig. 8 Ptiotomultiplier signals in spray

-0.5

--0.5

4/12/91 Delavan WDB 20-45 deg

40 psi, z~1 52mm, on aotis 3041 counts

log-normal fit: y=Cx exp^-0.5[ln(x/x0)/s] 2} xO = 20.45

s - 0.572 C-2936

peal< - 1 4.8 umD SMD - 46.3 umD

rel span = 1.19

140-

0 9 18 26 34 42 52 62 72 82 (umD)

Fig. 9 (a) Drop size distribution in spray

6.9 10.4 13.9 1 7.3 20.8 24.3

Fig. 9 (b) Speed distribution in spray

(m/s)

EXPERIMENTAL RESULTS

Measurements in a Delavan WDB20-45° nozzle spraying water at 276 kPa, 685 cc/min, 152 mm from

the nozzle on the centerline, are shown in Figures 8 and 9. Similar shaped drop size distributions have been

reported [6,14,15].

To insure that the rapid cutoff on the low side of the drop size histogram was not due to instrimient

sensitivity, the time traces were monitored. The smallest drops detected had locater pulses well above the

trigger threshold, so that drops smaller than the 5 /zmD limit in Figure 9a could have been detected. The

limiting drop size is probably due to the Umited kinetic energy available to atomize the hquid.

The log-normal function gives the best fit to both the diameter and speed histograms, as shown in

Figure 9:

f(x) = (C/x)exp{-^[hi(x/xO)/s]2} ; normaUze: C = l/[s(27r)^]

The parameters xO and s control the peak and width of the distribution. An upper limit modification to the

log-normal function has been suggested [16], however, the addition of a third parameter confuses the data

analysis. Such a limit is better added at the modeling stage.

634
The Nukiyama-Tanasawa (N-T) function has also been used:

f(x) = C XTM exp[-(x/xO)P] ; normalize: C = p/{xO ^ (m + 1) T [(m + l)/p]

}

where F () is the gamma function. When exponent m is restricted to (p-1) it is termed the WeibuU function

[17]. One difficulty is that the N-T function does not maintain a similar shape. For wide distributions it is skewed left, appearing similar to the log-normal function, whereas for narrow distributions it becomes skewed

right. Also, if m< 0 it predicts an infinite number of drops as the diameter approaches zero. Another popular function is the Rosm-Rammler, usually expressed as a cummulative volimie fraction:

Fy(x) = l-exp[-(x/xO)P]

however, ^ ^

so that it is seen to be simply the Nukiyama-Tanasawa function with m = p-4, a fact noted by Bevans [18].

For application to combustion models, an efficient procedure should exist to generate random variates

from the selected distribution function. When the cumulative distribution function (cdf) can be analytically

inverted, random variates are generated simply as X = F'^ (R), where F"^ () is the inverse cdf and R is a random number from 0 to 1 [19]. Examples are the exponential function, used in Los Alamos' KIVA code

[20], and the Weibull function. The log-normal cdf cannot be inverted, however efficient methods exist for

generating random variates from it [19, p 132].

The Delavan nozzle used was the subject of a "round-robin" test, using 6 different drop sizing

instruments [21]. At the test conditions of Figure 9a they reported Sauter Mean Diameters (SMD) from 65 to 180 fimD and "relative spans" from 0.95 to 1.7. The present data gives an SMD of 46.3 /imD and 1.19 relative

span. However, the SMD is a poor basis for comparing distributions from single particle counters since it is overly sensitive to the tail, where the counts cU"e very low.

The beam attenuation through the center of the spray was 19.4%, with an rms fluctuation of 2.42%.

The previous attenuation analysis calculates an equivalent monodisperse drop size of 16.4 ^mD, for the

400 ;zmD (minimum) main beam, which is close to the peak value of 14.8 jUmD in Figure 9a. Doubling the

pressure to 552 kPa hardly changed the peak diameter (15.8 /zmD) but narrowed the distribution (SMD =

36.3 /imD), while the peak velocity increased from 8.93 m/s to 11.0 m/s.

CONCLUSIONS

A low-cost instrument has been developed to allow drop size and speed measurements in dense sprays.

The design overcomes previous difficulties with Pulse Intensity methods. A simple "scattering tomography"

zmalysis accounts for beam attenuation through the spray. Drop size and speed distributions in the Delavan

spray are best fit by a log-normal distribution.

ACKNOWLEDGEMENTS

The research followed a 1986 UES/AFOSR Summer Faculty Research Program (SFRP) tenure at the

Air Force Phillips Lab and a 1988 SFRP tenure at Arnold Engineering Development Center. Michael Powell

at the Phillips Lab assisted with the initial design and Chad Limbaugh at Arnold Engineering supported tests

in a model rocket injector. Research sponsored by the Air Force Office of Scientific Research, under

Contract F49620-88-C-0053.

635
REFERENCES

1. Azzopardi, B.J., "Measurement of Drop Sizes", Int. Journal of Heat and Mass Transfer, Vol 22,

pp. 1245-1279 (1979).

2. Beretta, F., Cavaliere, A., and D'Alessio, A., "Drop Size and Concentration in a Spray by Sideward Laser

Light Scattering Measurements", Combustion Science and Technology, Vol 36, pp 19-37 (1984).

3. Presser, C, Santoro, R.J., and Semerjian, H.G., "Velocity and Droplet Size Measurements in a Fuel Spray",

AL\A 24th Aerospace Sciences Meeting, AL\A-86-0297, (Jan 1986).

4. Bachalo, W.D. and Houser, M.D., "Phase/Doppler Spray Analyzer for Simultaneous Measurements of

Drop Size and Velocity Distributions", Optical Engineering, Vol 23, no 5, pp 583-590, (Sept/Oct 1984).

5. Holve, D. and S.A. Self, "Optical Particle Sizing for in situ Measurements, Part 11", AppUed Optics 18,

p. 1646 (1979)

6. Mizutani, Y., Kodama, H., and Miyasaka, K., "Doppler-Mie Combination Technique for Determination

of Size-Velocity Correlation of Spray Droplets", Combustion and Flame, Vol 44, pp 85-95 (1982).

7. Hess, C.F. "Nonintrusive Optical Single-Particle Counter for measuring the Size and Velocity of Droplets in

a Spray", AppUed Optics, Vol. 23, no 23, pp 4375-4382 (Dec 1984).

8. Eauret, P.R. et al, "Measurement of Drop Size and Velocity in Vaporizing Sprays, in Proceedings, Second

Int'l Conference on Liquid Atomization and Spray Systems. Madison, Wisconsin, p 261 (June 1982).

9. Wang, J.C.F. and Hencken, K.R., "In situ Particle Size Measurements using a Two-Color Laser Scattering

Technique", Applied Optics, Vol 25, no 5, pp 653-657 (March 1986).

10. Umhauer, H., "Particle Size Dist. Analysis by Scattering Light Meas. using an Optically Defined

Measurement Volume", Journal of Aerosol Science, Vol 14, no 6, p 765 (1983).

11. Hess, C.F., "An Instrument to Measure the Size and Velocity of Particles in Particle Laden Flovi's", 21st

Joint Propulsion Conference, Moterey, California, AIAA-85-1443 (July 1985).

12. Sakai, T., Sadakata, M., Saito, M., et al, "Uniform Size Droplets by Longitudinal Vibration of Newtonian

and non-Newtonian Fluids", Second Int'l Conference on Liquid Atomization and Spray Systems, Madison,

Wisconsin, pp 37-45 (June 1982).

13. Keng, Weayer, and Orr. "A Study of Vibrating Capillary Atomizers", Georgia Institute of Technology

Engineering Experiment Station Report, Grant B-279 (April 1967).

14. BoUe, L. and Moureau, J.C., "Spray Cooling of Hot Surfaces" in Multiphase Science and Technology,

VoU. (G.F. Hewitt, J.M. Delhaye, and N. Zuber, Eds.), pp. 1-10, Hemisphere Publ., Washington (1982).

15. Hiroyasu, H., Kadota, T., and M. Arai, "Supplementary Comments: Fuel Spray Characterization in Diesel

Engines" in Combustion ModeHng in Reciprocating Engines (Mattayi, J.N. and Ammna, C.A., Eds.),

pp. 369-405, Plenum Press, N.Y. (1980).

16. Dobbins, R.A. "Light Scattering and Transmission Prop, of Sprays", Phd. Thesis, Princeton Uniy. (1960).

17. Derman, C, Gleser, L., and Olkin, I., A Guide to Probability Theory and Applications, Holt, Rinehart, and

Winston (1973).

18. Beyans, R.S., "Mathematical Expressions for Drop Size Distributions in Sprays", Conference on Fuel

Sprays. Uniy. of Michigan (March 1949).

19. Dagpunar, J. Principles of Random Variate Generation, Clarendon Press (1988).

20. Amsden, A.A., J.D. Ramshaw, et al. "KIVA: A Computer Program for 2 and 3 - Dimensional Fluid Flows

with Chemical Reactions and Fuel Sprays", Los Alamos National Laboratory Report LA-10245-MS

(Feb 1985).

21. Dodge, L.G., "Comparison of Drop-Size Measurements for Similar Atomizers", Special Report No.

SwRI-8858/2, Office of Nayal Reseach, (December 1986).

636
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 70

MEASUREMENT OF DROPLET SIZES BY LIQUID NITROGEN

FREEZING METHOD

T. Karasawa*, S. Shiga' and T. Kurabayashi^

*Dept. of Mechanical Engineering, Gunma University, Kiryu, Gunma, Japan

^Hiroshima Institute of Technology, Hiroshima, Japan

ABSTRACT

A liquid nitrogen freezing method for accurate measurement of droplet sizes

has been developed since 1971 and recently the final setup has been completed

through the efforts for overcoming some problems and difficulties. Droplet

sizes of water spray are measured by both the freezing method and the

immersion sampling method and the results are compared. It is revealed that

the number of droplets caught in the freezing method is several times as much

as that of the immersion sampling method. The reason of the decrease of

droplets number in the immersion sampling method is considered to be caused by

the disappearing of smaller droplets due to evaporation. Therefore the Sauter

mean diameter obtained by the immersion sampling method tends to be larger

than that by the freezing method. Through the comparison experiment with

immersion sampling method, it has been revealed the present apparatus can

afford highly accurate value of droplet sizes.

INTRODUCTION

It is strongly depired to obtain accurate droplet sizes in the fields of

liquid atomization technique. While various measuring methods have been used,

few of them seems to be able to afford the reliable value of droplet sizes due

to individual shortcoming essential to each method. The freezing method, in

which liquid droplets are solidified by low temperature atmosphere possesses

several advantages to obtain their accurate sizes. Since the first study on

this method by Longwell(l) in 1943, several works in this field have been made

by different investigators (2)-(7). In recent ten years, however, no

successful results has reported.

The authors started the research on this freezing method in 1971. A

cylindrical freezing chamber is air-proved to avoid generation of cloud in the

chamber. Liquid nitrogen fed into the double wall space of the chamber

vaporizes and forms cooling atmosphere. Liquid to be tested is injected

downward through a nozzle and its frozen droplets fall onto a photographic

film set on a bottom plate of the chamber. In order to avoid the severe

problems of frosting and optical difficulty, a simple photographic technique

called photogram is employed to get shadows of the droplets. In 1978, at the

637
first ICLASS, history of development of the research from the first trial

setup to the sixth setup was reported and availability of this method for

accurate measurement was suggested on the basis of the experimental results.

In 1979 and 1980, remarkable improvements had been made in the method

and the seventh trial setup was designed and developed. Prior to the

construction of this device, distance neccessary for droplets to freeze was

theoretically estimated and evaporation of droplets were also calculated. The

most important improvement in the setup was employment of 100 feet long film

in a film suply system, which enables to obtain a significantly increased

number of data in one series of experiment using liquid nitrogen filled in a

150 liter Dewer tank. The film suply system was also newly designed to protect

the film against sudden exposure to the atmosphere with large temperature

difference from room to freezing chamber, to avoid the temperature shock which

deteriorates film emulsion. Using the setup, adequate exposure time to

compensate the decrease of light sensitibity of the film emulsion due to

lowered temperature, and adequate axial temperature gradient in the test

chamber to form spherical frozen particles were searched experimentally. All

these results were already reported (8).

Through further improvement made in recent years, the final setup has

been completed for accurate measurement of droplet sizes. In this paper

procedure of the improvements is described at first, and the accuracy of the

freezing method is examined by comparing the droplet sizes obtained by this

method with those by the another method. This time, the immersion sampling

method is chosen as a representative of the other methods, which is widely

used at present and is often employed as a standard method when comparison is

required

.

EXPERIMENTAL SETUP AND PROCEDURE

The arrangement of experimental devices is shown in Fig.l. The freezing

chamber (32) is composed of a double wall type cylinder and a top and bottom

end plates to keep airtight for avoiding generation of fog in the chamber. The

inside of the chamber is kept dark. A length of 2 meter of the chamber was

determined considering the analytical result mentioned above, and the inner

diameter is 600 mm. The cylinder walls are made of 0.5 mm stainless steel

plate. The thickness of the previous body wall was 1.0 or 1.2 mm, and this

reducton of the wall thickness produced decreased heat capacity of the body,

nitrogen consumption economy and easy control of ambient gas temperature.

Liquid nitrogen in the Dewer tank (14) is poured into the gap between the

inner and outer walls of the chamber cylinder and the evaporated nitrogen

forms freezing atmosphere in the chamber. The amount of nitrogen supplied is

measured by using a spring balance (13). The chamber is thermally insulated

by foamed polystyrene with a thickness of 150 mm. Distribution of temperature in the freezing chamber is measured by thermocouples (29) set at seven

heights, which can be moved radially by rotating a support bar (30). Thermocouples are also inserted in other parts, injection nozzle (5), film boxes

(17) and (28), etc.. All of these sixteen thermocouples are connected to a

data-logger(l ) and their measurement time is controled by a micro computor.

When the temperature distribution in the chamber achieves required

profile, a new part of a 35 mm roll type long film is pulled out from the film

box (28) through the temperature buffer (25) at adequate speed for avoiding

heat shock, and is set on a film guide (23). On the guide the spread film

forms a collective area of frozen particles with a width of 24 mm and a length

of 600 mm. These operations are made automatically by a rolling system

638
Fig.l The eighth apparatus

consisted of a photo-detect counter (27), an inverter motor and an electronic

control circuit. Liquid to be tested is then injected downward into the

freezing chamber. When the smallest frozen particles are considered to reach

the film completely, shadows of the frozen particles are recorded on the film

by the technique called photogram, at which direct exposure is made to the

light from the bulb (2) for adequate duration determined depending on the

coldness of the film. As soon as the exposure is finished, the exposed part

of the film is rolled up into the film box (17), and the frozen particles on

the film are replaced by a suction unit (15). The photogramed particles are

printed with magnification of 20 and then measured.

Since the experiments for comparison of the two methods have to be

carried out under the same conditions as far as possible, the chamber of the

freezing method is also used to carry out the experiment of immersion sampling

method and the sampling cells are placed at the same position as that of the

film supply system in the freezing method. All the experiments are carried

out on water spray injected through a hole nozzle with 0.3 mm diameter.

Injection pressure is changed in four steps, i.e. 12, 10, 8 and 6 MPa, and the

injection period is changed also in four steps, i.e. 150, 90, 40 and 30

millisecond. The injection period of 90 ms is set as a standard injection

period. The quantities of water injected are chosen as small enough not to

cause the coalescence of droplets at the immersion liquid surface and large

enough to offer reasonable number of particles.

639
RESULTS AND DISCUSSIONS

Influence of Exposure Time to Low Temperature Atmosphere on its Negative

It is known that the application of sudden change of temperature to the film

sometimes causes unevenness in density and spots or patches with irregular

shape on its negative. However the details of these phenomena have not been

clear. Therefore the following tests were carried out. The elapsed time from

the setting of a film in alow temperature atmosphere of -100 C to the light

exposure was changed as 0, 30 and 60 minutes, and then the film was developed.

The result showed that the negative without elapsed time is normal, but a lot

of fine spots appeared on the negative with 30 minutes elapsed time and the

number of the spots increase in the case of 60 minutes. However since the

contour of the spot is not clear, while that of the shadow of the frozen

particle is quite clear regardless of its size, both the spots and the

particles can be easily distinguished. Therefore, it is considered that the

elapsed time of 30 minutes which generates relatively small number of the

spots does not affect on the result of measurement. To avoid these

undesiarable phenomena, in the present experiments the film is introduced into

the chamber immediately before the liquid injection and rolled up into the

film box as soon as the exposure is finished.

Comparison of the Number of Droplets Captured

In Fig. 2 radial profiles of number density of droplet caught in the freezing

xio

10

eu

to

o.o

E3

Freezing method Pi=12MPa

O

Freezing method PhlOMPo

A Freezing method

PI=12MPa«25;im)

Freezing method

PI=10MPa(<25/im)

D—cr

Freezing method

PI=8MPo

PI|8MPaK25/im)

Freezing method

'PI=6MPa

' ^^--PI=6MPa{<25/jm)

Immersion method .PI=l2MPo

_____j/^=6MPa

w

0 100 200 300

Distance from nozzle axis r mm

Fig. 2 Variation of the number of

droplets caught by both

the freezing method and

immersion method

^^j^i^^y^y^HA^^'C • •

• • ;

•

.. •••,nVvV '^^^^^^ ••• •••• Jo

r«» too fioo

I

.* ' •^ .v,» <: •

- IT ; ^ r-soo

^.•v.;-'-:-v-;:iv-'';

r»2oo • .• • •

Freezing method Immersion method

Fig. 3 Photographs of droplets

640
method is compared with those in the immersion sampling method for four

injection pressure levels. Some representative photographs of droplets in the

two methods are shown in Fig. 3. The figure shows that the number of droplets

caught in the freezing method is nine times as much as that in the immersion

sampling method, and more than half of the number is occupied by the fine

droplets smaller than 25 micrometer. These facts are caused by evaporation

of the droplets during their flights in the ambient gas, and the increased

evaporation rate as the diameter becomes small (8). In the freezing method

droplets are prevented from evaporation due to the extremely low temperature

atmosphere and almost all droplets can remain and are captured. Fig. 2 also

shows the same tendency as above.

Comparison of Sauter Mean Diameter in Two Methods

Fig. 4 shows the radial distributions of local Sauter mean diameter obtained by

freezing method and immersion sampling method. In Fig. 4(a) it is shown that

the Sauter mean diameters by the immersion sampling method are much larger

than those by the freezing method over whole range of radial derection. A

similar tendency was observed in the case of injection pressure of 10 MPa.

The reason of these facts is that in the immersion sampling method the number

of finer droplets is decreased by evaporation as mentioned above. Radial

profiles of the mean diameter in (a) are more gentle than those in (b) and

(c). This tendency may be caused by the fact that in the high pressure

injection the size of droplets becomes smaller even at the center of the

spray, and the ralatively larger droplets in the vicinity of spray axis are

apt to be shifted toward periphery due to the high injection velocity.

In figure (b) in the case of injection pressure of 8 MPa, Sauter mean

diameters by the two methods come to nearly the same at the vicinity of spray

center. This is caused by that although the number of droplets obtained by

the freezing method is larger than that of the immersion sampling method, the

total number of droplets generated decreases and the number of large droplets

over a diameter of 200 micrometer becomes to increase as the injection

pressure decreases.

In figure (c) Sauter mean diameters by the immersion sampling method

becomes larger again than those by the freezing method. As the injection

pressure decreases the dispersion of spray becomes weak and the sizes of

droplets increase. These tendencies bring larger possibility of coalescence

of droplets when they are captured by immersion liquid. This fact shifts the

value of mean diameter near the spray center. In the periphery of spray the

mean diameters are shifted down nearly to the values of freezing method. In

this region the increase of mean diameter due to the evaporation of fine

droplets becomes small, the number of fine droplets decreases under the lower

injection pressure.

Comparison of Size Distribution

In Fig. 5 number distributions of droplets obtained by the freezing method are

compared with those by the immersion sampling method. Difference is seen in

the region of fine droplet under the condition of high injection pressure.

Influence of Injection Period on Sauter Mean Diameter

Injected liquid quantity was changed by changing the injection period to

reveal their influence on Sauter mean diameter, and the results are shown in

Fig. 6. As seen in Fig. 6(a) the mean diameter by the freezing method with

injection periods of 30 and 1 0 millisecond is larger than those with 90 and

150 millisecond. This may be caused by the relatively longer transient period

641
200

E

n150

01E

o

g'OO

0}Eu

01 »-o

oo

o

-I 50

PI=l2MPo

t=90msec

Y=2000mm

Wafer

o Freezing mefhod

A Immpr^lnn

sompUng mefhod

A

O A

o o

• 1 1

200

0 100 200

DIsfonce from nozzle oxis r mm

(a) Pi » 12 MI'a

Fig. 4 Comparison of local SMD byfreezing method with that

by immersion method

300

|l50

100

«-

as *-

0)E

a

'oc

D

V

E

L.V

+-3

O

V)

o 50

200

E

Nl50

u

0) +-w

E

o

T3glOO

41E

0) +-o

tn

§ 50

Pi=8MPo

t=90msec

Y=2000mm

Woter

o Freezing meffiod

A Immersion

sompling method

0 100 200

Distance from nozzle oxIs r mm

(b) PI = 8 MPa

300

Pi=6MPo

t=90msec

Y=2000mm

Woler

o Freezing method

A Immersion

sampling method

o

0 100 200 300

Distance from nozzle oxIs r mm

(c) PI - 6 MPa

642
of nozzle needle motion in the case of shorter injection period where

insufficient atomization occurs. In the immersion sampling method these

differences are not recognized. Based on this result, it can be said that

the freezing method is more sensible than the immersion sampling method.

Diameter d /im Diameter d fun

(a) Pi = 6 MPa (b) Pi = 12 MPa

Fig. 5 Distribution of droplet size

200r

E

~150|-

o

E

_o

•a

DlOOl01£

<u

-+-O

to o O

DO

o

8

50-

Freezing method

Pi=12MPa

Y=2000mm

Water

D t=30msec

A 40

O 90

V 150

200

0 100 200

Distance from nozzle axis r mm

300

E

~ 150

E

_oT3§100

03E

3

D

oo

o 50

V

A A

a>o

immersion

sampling method

Pi=12MPG

Y=2000mm

Water

o t=30msec

A 40

O 90

V 150

2 0

0 100 200

Distance from nozzle axis r mm

300

(a) Freezing method (b) Immersion method

Fig. 6 Influence of injection period on local SMD

643
CONCLUSIONS

A new apparatus for accurate measurement of spray droplet sizes based on the

liquid nitrogen freezing method has been constructed after the constructions

of seven trial setups. Using this apparatus comparison tests with the

immersion sampling method were carried out and the results are summerized as:

In the present apparatus improvements were made on some points for more

reliable measurement, easy operation and nitrogen economy.

The immersion sampling method presents larger mean droplet diameter

mainly due to disappearing of fine droplets by evaporation and coalescence of

droplets occurs when they are captured by immersion liquid.

In the freezing method the evaporation and coalescence of droplets are

almost avoided and, therefore, the method can afford highly accurate spray

diameter.

Since the freezing method is based on a quite simple principle and has

little possibility of error, the present setup is considered to be a standard

measuring apparatus.

ACKNOWLEDGMENTS

The authors wish to thank Mr . Hisao Nakamura for his enthusiastic helps

for constructions and tests of the apparatus.

This work was partly supported by the Tanikawa Foundation, Osaka, Japan.

REFERENCES

1 Longwell J P, Doctorial dissertation, Massachusetts Institute of Technology, 1943.

2 Taylor E H and Harman D B, 'Measuring Drop Sizes in Sprays', Industrial and

Engineering Chemistry, Vol.46, No. 7, pp. 1455-1457 , July 1954.

3 Tanasawa Y, 'Automobile Engines, Lecture', Toyota Gijutu, Vol.5, No. 7,

pp. 24-34, 1952.

4 Joyce J R, 'The Atomization of Liquid Fuels for Combustion', Inst. Fuels

J., Vol.22, No. 124, pp. 150-156, 1949.

5 Nelson P A and Stevens W F, 'Size Distribution of Droplets from Centrifugal

Spray Nozzles', J. Am. Inst. Chem. Eng., Vol.7, No.l, pp. 80-86, Mar. 1961.

6 Street P J and Danaford V E J, 'A Technique for Determining Droplet Size

Distributions Using Liquid Nitrogen', J. Inst. Petroleum, Vol.54, No. 536,

pp. 241-242, Aug. 1968.

7 Rao K V L, 'Liquid Nitrogen Cooled Sampling Probe for the Measurement of

Spray Drop Size Distribution in Moving Liquid-Air Sprays', Proc. 1st

ICLASS, Tokyo, pp. 293-300, Aug. 1978.

8 Karasawa T, Kaeriyama H and Kurabayashi T, 'Freezing Method by Liquid

Nitrogen for Measuring Spray Droplet Sizes', J. of the Fuel Society of

Japan, Vol.62, No. 679, pp. 901-912, Nov. 1983.

644
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 71

LDV AND PDA ANALYSIS OF SPRAY-WALL INTERACTION

G. Brunello', S. Callera*, A. Coghe^, G.E. Cossali^ and F. Gamma*

*CNPM-CNR, Peschiera Borromeo (MI), Italy

^Politecnico di Milano, Milano, Italy

*Universitk "La Sapienza", Roma, Italy

ABSTRACT

A diesel spray impacting on a wall in a close chamber, under

ambient air conditions, has been analysed by means of Laser

Doppler Velocimetry (LDV) and Phase Doppler Anemometry (PDA) in

order to measure droplet size and velocity in regions close to

the wall. Spray behaviour has been characterized in terms of

different injection periods and different spatial regions, and

proper data analysis allowed to identify the main features of

the droplet-wall interaction phenomenon.

INTRODUCTION

Recent theoretical predictions, experimental observations and

engine tests [1] evidenced that the interaction of a diesel

sprays with the piston wall might be beneficial to engine

combustion because of the influence on fuel vaporization and

heterogeneous mixture formation processes. Many experiments have

been performed in the past by different groups [2,3,4,5] to

investigate the spray-wall interaction, generally using highspeed photography, and the first submodel developed to describe

the spray-wall interaction process has been presented very

recently [6]. Spray-wall interaction characteristics seem to be

controlled mainly by the drop Weber number although other

parameters as wall temperature, fuel boiling temperature, drop

impinging angle, swirl intensity, air density and temperature,

Leidenfrost temperature, wall surface roughness have also been

proven to be important. Analysis of spray-wall interaction is

furthermore complicated, due to the interaction of spray

droplets with entrained air flow [7], which makes unrealistic to

extrapolate data obtained in experiments with single droplets.

Development and validation of a spray impingement model requires

a large amount of experimental data on the droplets of a real

spray, like size and velocity in the region close to the wall.

The present research work was aimed to a detailed analysis of

the spray wall impingement process, by using the Phase Doppler

Anemometer (PDA) to obtain simultaneous measurements of spray

drop size and velocity in the close proximity of the wall.

645
EXPERIMENTAL SET-UP

The experiment was performed in a closed cylindrical chamber (89

mm i.d.,300 mm height) equipped with two glass windows

positioned to allow the use of the PDA in the 70 deg forward

scattering configuration to reduce the effective probe volume

dimensions ( about 1mm in length and o.l5 mm diameter) and the

sensitivity of the phase shift to refractive index of the drops

[9]. A standard diesel injector (Bosch single hole nozzle, 0.25

mm diameter) injecting at 45 deg with respect to the nozzle

axis, was mounted off -axis on the top wall of the chamber. The

injector was fed by a Bosch high-pressure pump driven at 500 rpm

by an electrical motor and diesel oil was used as fuel. The

impact wall, an aluminium disc (50 mm diameter) with a polished

surface, was centred on the chamber axis and positioned at a

distance of 20 mm from the nozzle tip. (see fig.l). The

impingement distance (L) and the impingement angle (45 deg) were

kept constant. The nozzle opening pressure was fixed at 20 MPa

and the injection pressure history is shown in Fig. 2. Operating

experience indicated that the injection system provides very

reproducible pressure time histories. The ambient gas was air at

room temperature and pressure. The operating conditions of the

experiment are summarized in Table 1.

Table 1 - Operating conditions

Nozzle diameter, d

Max. fuel inj. pressure

Pump speed

Volume of fuel injected

Duration of injection

(mm*/in j

)

(ms)

(mm)

(MPa)

(RPM)

70

500

36

0.25

3.4

Fig.l Schematic of the impingement

experiment (all dimensions in mm).

Fig. 2 The liquid injection pressure.

646
Drop size and velocity were measured by the Phase Doppler

Anemometer (Dantec PDA); the light source was an Argon-ion laser

(5 W nominal power) and the transmitting optics comprised a 40

MHz Bragg cell for frequency shifting and directional

sensitivity and a front lens of 300 mm focal length. The

receiving optics was the 57X10 Dantec unit equipped with three

preamplified photomultipliers . The covariance processor of the

PDA instrument was capable of 1 \is time resolution in data

acquisition and the data transfer rate, to a PC, was higher than

100 kHz. The frequency bandwidth (36 MHz) and the optical

configuration of the transmitting optics defined a measurable

velocity range from -25 to 125 m/s and a measurable size range

from 2 to 80|j.m which resulted the most appropriate in the

present case. Normal and tangential velocity components were

measured at the same point at different times and drop size was

measured together with the velocity component normal to the flat

plate.

Measurements were performed at several locations on three lines

at distances of 0.5, 1 and 2 mm from the surface (see Fig.l),

laying on the symmetry plane normal to the wall. The

investigated region extended from the impingement area to the

boundary of the impact disc, in the downstream direction. Data

acquisition, triggered by the needle lift opening at 20 MPa,

covered a period of 5 ms for each injection and 5,000 to 10,000

data were collected over many consecutive injections for each

measurement point. An encoder pulse with 1 \is resolution allowed

a subsequent reconstruction of the size and velocity time

history of each individual injection and also the application of

time-ensemble averaging procedures.

RESULTS AND DISCUSSION

A typical result for the drop size and two velocity components,

measured downstream the nominal spray impinging location, is

shown in Fig. 3, where all the individual measurements collected

over a large number of injections are reported versus time, the

origin, t = 0, being defined by the 20 MPa triggering level on

the injection pressure trace. As in free sprays [8] three main

periods are distinguishable: the first period with strong

unsteadiness due to the passage of the spray head, a main period

where most of the fuel is injected, and a late period

characterized by the trailing edge, where the smaller droplets,

left behind by the passage of the previous structures, and the

bigger ones, generated during the late injection at low

pressure, are travelling together at low velocity. Close to the

surface (h=0.5 mm in Fig. 3 ), after the tip arrival, tangential

velocity stabilizes, with high rms, around a value larger than

the tip velocity and than decays after completion of injection.

The corresponding normal (i.e. perpendicular to the wall)

velocity component shows that only droplets traversing the

control volume with negative velocities (i.e. back from the

wall) were detected (Fig. 3 ). The size distribution shows an

upper bound of approximately 30 |jim and few data at higher

values. At the same location (X = 6 mm) and larger distance from

647
the wall (h=2 mm) the tangential component (Fig. 4 ) shows a

transient behaviour in correspondence of the early and late

injection periods, while the velocity of a large fraction of

droplets is almost zero during the main injection period,

suggesting that the wall spray does not extend beyond 2 mm from

the wall. At the same point, it appears clearly the coexistence

of droplets traversing the control volume with positive and

negative normal velocity components (fig. 4 ). Comparison with

Fig. 3 shows that the droplet sizes appear lower at h = 2 mm,

during the main injection period.

The relative steadyness of normal and tangential velocity

components during the main injection period, although the

variability is high, suggested the following data analysis:

tangential velocities were ensemble averaged over a period from

1.5 to 2.5 ms after injection start (AIS) and the results

plotted versus distance from the wall, at X = 14 mm, and the

same was done with the size data. The results (Figs. 5 and 6)

show that the maximum mean velocity and mean diameter were

reached both at the closest distance from the wall with a steep

gradient, suggesting that the most of the tangential droplet

momentum remains concentrated close to wall during the main

injection period.

80.08

60.80

40.80

20.80

0.00

20.00

10.80

0.00

-10.00

-20.00

H — O . S

0 ime 2000 sooo /Snoirm. voloc±tiy C m/ 3 )

4000 5800

.--5'.., l:-v :.> •:»>%: /vr-y

80.88

60.88

48.88

20.00

a. 80

20.88

18.80

8.88

-18.88

-28.88

^Diatwteic [unl H — 2 nun

9 1888 2000 3000

noT-m. velocity Cm/s)

4800 5888

.> j>'..'».-..

2888 3000 4000 5888

1888 2888 3888 4888 snaa

128.88

180.00

80.00

60.88

-28.88

tians * velocity Cm/s>

2880 3808

time C >is )

4oee soee

128.88 -I tang. velocity Cm/s>

\ 108.88 .1

88.88

68.88

48.88 J

i 1 28.88 \

8.88

-28.88 J- 1088 2888 3880

time C>is)

4000 5880

Fig. 3 Time evolution of drop

size and velocity at X = 6 ram.

Fig. 4 Time evolution of drop

size and velocity at X = 6 ram.

648
02468 10 02468 10 distance. H [mm] distance, H [mm]

Fig. 5 Fig. 6

The time evolution of the droplet velocity vector at larger

distances from the wall allows to get insights of the air

entrainment process. Each velocity component has been ensemble

averaged over time windows of 10 )us and the results at X = 14 mm

and h = 3.5 mm are presented in Fig. 7. At the very beginning of

the spray passage the velocity vector is almost parallel to the

wall, slightly upward tilted, then , after performing a complete

revolution, it reachs the downward orientation, at 1 ms AIS,

with magnitude reduced to few m/s, and it maintains such

orientation during all the main injection period, indicating the

existence of entrainment by the wall spray. At lower distances

from the wall the velocity vector slightly oscillates up and

down around the parallel orientation, particularly during the

early period. The presence of periodic oscillations of the

normal velocity component was clearly identified at h = 2.5mm on

the ensemble averaged time history, just after the tip passage,

as shown in Fig. 8. It is remarkable the coherence of this

structure which reproduces itself almost in phase over a large

number of injections, with a characteristic period of about 0.25

ms which corresponds to a frequency of 4 kHz.

The tip arrival time at any location downstream the impact

nominal point (X = 0) can be detected from the tangential and/or

normal velocity histories, as the time at which the first high

velocity is measured. The results are reported in Fig. 9 as

penetration versus time for two distances from the wall: h = 0.5

and 2 mm, from wich the mean tip velocity was deduced.

Comparison with the results obtained on the free spray [8] shows

a decrease of the tip velocity to about 50% of the free spray

value. The spray front is not normal to the wall as it stems by

the fact that at 0.5 mm the leading droplets reach the

measurement location systematically before the corresponding

droplets at 2 mm. The measured time delay is independent of

location, suggesting that the front shape does not change with

distance from the impingement location and implying a wall/front

angle of about 30 deg.

649
10 m/s

t—O

1

20

10

u

Q>>

O 0

o

c

-10

-20

a

—1 1 1 1 1 1 A ....

0.5 Fig. 7 Fig. 8

1 1.5 time [ms|

The analysis of all velocity probability distribution functions

(PDF), for both the normal and tangential velocity components,

revealed the presence, in many cases, of bimodal distributions,

indicating the coexistence of two droplet populations. Although

it was easy to identify the two groups of droplets,

characterized by opposite signs of the normal velocity

component, as the impinging and reflected ones, it was puzzling

to which of the two populations were associated the two peaks in

the tangential component pdf, as the two velocity components

were measured separately and thus no correspondence between

individual data exists. Anyway, an attempt was made to

reconstruct the velocity profiles along the X orientation, by

using the ensemble averages of each population, for both

components, and typical results are shown in Figs. 10, for the

time window of 0.2 ms after tip arrival.

It can be noticed that one of the two curves representing the

tangential component shows a similar profile and peaks at the

same location where the normal velocity profile of the impacting

60

^

40

c

o

c

a> 20

////

In 2 mm

60

40

20

150

125

100

75

•o 50 >

25

0

-25

Vr

o s VIWr

Wl

•ft

'C^ • -o -

0.1 0.2 0.3 time [ms]

0.4 0.5 0 6 12 distance, x [mm]

18

Fig. 9 Fig. 10

650
droplets peaks too. This must be considered as sufficient to

regard such part of the tangential velocity profile as relative

to the impinging droplets. In fact, due to the 45 deg spray

impact angle, the normal and tangential drop velocity components

should be of comparable magnitude inside the impingement region

(-6 < X < +2). Downstream, in the wall region, the normal

component is only negative and the tangential velocity reaches

the extreme values. In the transition region, incoming droplets

whose tangential velocity rapidly decreases, coexist with

reflected or deflected droplets whose tangential velocity

component rapidly increases toward the high values,

characteristic of the wall spray. It must be stressed that the

measurements in the impingement region (-6 < X < +2) resulted

more difficult and in some cases, especially during the main

injection period, actually impossible, due to the higher drop

number density. Fig. 11 shows a vectorial plots of the droplet

flow field, based on the same data of Fig. 10 , from which the

characteristics of each wall region and droplet population might

be better understood.

Fig. 11 Vectorial plot of the

droplet flow field 2 mm from

the impact wall (same data of

figure 10 )

.

CONCLUSIONS

The experimental study of a diesel spray impinging at 45 deg on

a flat surface under ambient gas conditions revealed the

following;

1) The time behaviour of the wall spray strongly resembles that

of a free spray; three different periods may be recognised: i)

an early period lasting about 0.2 ms, characterised by the

passage of the spray head and by high velocity unsteadiness; ii)

a main period lasting about 70% of the injection duration,

characterised by an almost steady spray, although the

variability is high; iii) the late period, characterised by the

presence of low velocity droplets.

2) The wall region my be subdivided into three regions with

different characteristics: a) the impingement region

characterized by still high normal velocity component and high

droplet density; b) the wall spray region characterized by high

tangential velocity component and comparatively low normal

component, where the droplet density is lower and allows direct

measurements of drop size and velocity; c) a transition region

651
in between, with characteristics of both of the previously mentioned regions.

3) Estimated tip velocity is about 50% of that of the impinging

free spray; the spray front appears to move with an angle of

about 30 deg respect to the wall, and such value is not changing

with distance from the impingement location.

4) Droplet tangential momentum remains concentrated close to the

wall, the maximum mean tangential velocity as well as the

maximum mean droplet size were found at 0.5 - 1 mm from the

wall.

5) Double peak PDFs, found at many measurement locations,

suggest the existence of two droplet populations characterized

by different trajectories. The data analysis here reported

suggests that such populations represent impinging and reflected

droplets, the latter probably coming from the impact of the

former on the wall.

REFERENCES

[1] Binder K. and Pfeffer V.; "New results of research into

mixture formation in the D.I. Diesel engine"; Intn. Symp.

COMODIA 90, Kyoto, Japan, September 1990.

[2] Katsura N. et al.; "Characteristics of a Diesel spray

impinging on a flat wall"; SAE Paper No. 890264, 1989.

[3] Fujimoto H. et al . ; "Characteristics of a Diesel spray

impinging on a flat wall"; Intn. Symp. COMODIA 90, Kyoto,

Japan, September 1990.

[4] Hiroyasu H. et al.; "Computed tomographic study on internal

structure of a Diesel spray impinging on a flat wall";

Intn. Symp. COMODIA 90, Kyoto, Japan, September 1990.

[5] Coghe, A. , Cossali, G. , Gamma, F.; "LDV Analysis of SprayWall interaction"; 6th Annual Conf. ILASS-Europe, Pisa, 1990

[6] Naber J.D. and Reitz R.D.; "Modeling Engine Spray/Wall

Impingement"; SAE Paper No. 880107, 1988.

[7] Cossali G.E., Brunello G. , Coghe A.; "LDV characterization

of air entrainment in transient diesel sprays"; SAE paper

910178, Detroit 1991.

[8] Brunello G. et al.; "Simultaneous evaluation of tip and drop

velocities in unsteady Diesel sprays"; Workshop on Fluid

Dynamics and Combustion in Reciprocating Engines, CNRIstituto Motori, Capri, Italy, April 1990.

[9] Pitcher G. , Wigley G. and Saffman M. ; "Sensitivity of

dropsize measurement by Phase Doppler Anemometry to

refractive index changes in combusting fuel sprays"; Fifth

Intn. Symp. on Application of Laser Techniques to Fluid

Mechanics, Lisbon, July 1990.

[10] Matsuoka S.; "Combustion in the Diesel Engine"; in : "Internal Combustion Engineering, Science and Technology"

Ed. J.H. Weaving, Elsevier Applied Science, London 1990.

Support for this work was provided by the Cornmission of the European

Curranunities, under Program "Joule" (Contract N. JOUE-0012-D (AM)).

652
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 72

APPLICATION OF POISSON STATISTICS TO THE PROBLEM OF SIZE AND

VOLUME FLUX MEASUREMENT BY PHASE-DOPPLER ANEMOMETRY

C.F. Edwards and K.D. Marx

Combustion Research Facility

Sandia National Laboratories

Livermore, CA, U.S.A.

ABSTRACT

This paper outlines the effects of Poisson particle statistics on the ability of a "single-particle" diagnostic to make

accurate size and volume flux measurements. Although the analysis presented is quite general, a simplified flowfield and

probe volume have been used to illustrate its consequences. Even this simple example suffices to show that the action of the

Poisson filters places restrictions on the quality of measurements by techniques such as phase-DoppIer anemometry. The

keys to the filters are the expected values of the number of particles in the probe voliune and the number of particles entering

-2 the probe region during the residence time of a preceding particle. Only if these values are kept below order 10 can the

measured joint distribution function and particle rate be assumed to reflect the true distribution and rate of the flow.

INTRODUCTION

Phase-Doppler Anemometry is a technique via which both the size and velocity of a spherical particle may be

determined with good accuracy, high speed and good spatial resolution. Because of this capability it is possible to compute

particle size distributions, size-velocity correlations and even liquid volxmie flux rate for spray applications. However, some

controversy has arisen recently regarding its ability to provide accurate volume flux measurements in test flows designed

specifically for that purpose. In other flows which are of more practical interest, there is no way to know how accurate the

measured volimie flux is, so this unknown must be carried as a caveat with the data set when it is used.

This issue was brought clearly into focus at a recent ILASS-Americas meeting in Hartford, CT. At this meeting a paper was presented by Dr. L. Dodge and co-workers at Southwest Research Institute which showed a potential

imderaccoimting of the liquid volume flux in a test spray. Other papers were presented which reported data for liquid volume

flux, but which were unable to quantify the bands of imcertainty associated with the data.

During the discussion of these papers. Prof. D. Senser of Purdue University suggested that the major difficulty might

simply be a result of the requirement that there be only a single particle in the probe volume. He also noted that in flows

where a strong size-velocity correlation exists, errors might be introduced into the measured size distribution function.

This paper investigates the potential for accuracy in such measurements in light of Prof. Senser's suggestion. It outlines the sampling process of an ideal phase-Doppler system in terms of the Poisson statistics of particles in the probe

volimie. The results show that even for an ideal processor, Poisson statistics impose severe constraints on the ability of any

"single particle" diagnostic to accurately measure size and voliune flux.

Since the need for brevity precludes our giving a systematic derivation of the analysis along with examples of the concomitant effects, we attempt here to provide only the backbone of the analysis and a single illustration in a simplified

flowfield. The reader who is interested in a more systematic exposition, evaluation of assumptions, and illustration of effects,

should contact the authors to obtain a copy of the full manuscript [1].

653
AN IDEAL PHASE-DOPPLER SYSTEM

Consider the flow of particles through the probe volume of an ideal phase-Doppler system. The system which we

postulate is not connected with any hardware, but is a region of space in which particles of any size are detected with perfect

accuracy and in which they exist only as points (inflnitesimal spatial extent). Since the idealized particles are of infinitesimal

spatial extent, no two can occupy the same position at the same time and no interactions between particles are possible (e.g.,

collisions).

Anticipating that the probe volume of an actual phase-Doppler system is a function of the particle size, the probe

volume considered here is formed by combining probe volimies of various sizes, AV((j>), corresponding to each particle

diameter (j). Each of the probe volumes senses only those particles which are of size ^ associated with it. The probe volumes

may be of arbitrary shape but all are subject to the constraint that AV cannot be reentrant. All of the probe volumes are assumed to be sufficiently small that the particle statistics are constant over the dimensions of the composite probe volume

and the velocity and size of a particle are constant during its transit through the composite probe volume. Each of the component probe volumes has associated with it an area, Ap(y,^), which is its cross sectional area projected in the direction of

a particle traveling with vector velocity v. The question to be answered then is a simple one. Given a known flux of particles through this region and the

restriction that there must be only one particle sensed in the composite probe volume at a given time for a measurement to be

considered valid, what will be the relationship between the validated and true fluxes through the region?

The answer to this question is that the validated flux of particles will reflect the true flux but only after having passed

through two stages of filters. Since the characteristics of these filters are specified by spatial and temporal Poisson processes, we will refer to these as Poisson filters. The first stage filter (Fig. la) is formed by a spatial Poisson process. Since a particle can only be measured if it is the only one in the probe volimie, then at the instant of entry of any particle into the probe volume there must be sufficient

interparticle separation that the previous particle has already left the probe volume. If the separation is not sufficient, the

particle will be invalidated.

Consider a flow with only a single class of monodisperse particles of size (j), traveling through the probe volume.

Denote this class of particles by the subscript i. Since all particles within class i are the same size, then the composite probe

volume is the same as the component probe volume and is given by AVj- = AV(({)j-). The probability that a particle in this class A

will pass through the spatial filter is the same as the probability that the interparticle volume, Vj-, is greater than the probe

A volume, AV,-. If the probability distribution of interparticle volume for this class is denoted by Pxi(Si) (normalized to unity),

then the probability of passage through the spatial filter is given by

AV,.

Pxi = P(yi > AV/) = 1 - P(V,- < AV,) = 1 - fpxi(^i)dVi (1)

0

Once a particle has passed through the spatial Poisson filter — that is, it has an empty probe volume at the time of its entrance — the particle must still pass through another Poisson filter before it can be validated. This filter corresponds to the

particle being able to exit the probe volimie before another particle enters and thereby invalidates it (Fig. lb). It is apparent

that the probability that the particle will be able to do this will depend both upon its residence time in the probe volume and

upon the distribution of interparticle arrival times at the probe volume.

If the residence time of a particle in the probe voliune is denoted by T, and the probability distribution of interparticle

arrival times at the probe volume for class / is denoted by Pti^i) (normalized to unity), then the probability of passage

through this temporal filter for a flow with a single class i is given by

T

P,i = P(!ti>x) = l-P<Ji<x) = I- Jpti(Ji)<^i (2)

0

654
In order to complete this description we need expressions for the distribution functions PxiC^i) ^nd PiiOi)- These

expressions may be found by application of Poisson statistics to our point particle flux. The interparticle time distribution for

a continuous, homogeneous Poisson process is given by Snyder [2] as

Ptfii) = ^n-^'^"'' (3)

In a similar fashion, the interparticle volume distribution for a continuous, homogeneous process is given by

A PxA) = Ki^'^'"^' (4)

where X^^ and Xj^,- are the temporal and spatial process intensities. Note that, by definition, X^j is the mean number of

particles per unit volume, or mmiber density p,-, of the /th class of the spray.

Substituting the Poisson expressions for interparticle time and volume into the expressions for the spatial and

temporal Hlters yields the probability of measurement of a given single particle of class / with residence time z. What is

required, however, is the response of the ideal phase-Doppler system to a distribution of droplets (classes) within the spray. This may be accomplished by introducing a particle distribution function in the two quantities measured by the phase-Doppler

instrument: velocity and size. This distribution function may be cast in either spatial (particles per imit volimie) or temporal

(particles per unit area per unit time) terms. Both formulations have advantages and disadvantages. However, since the

phase-Doppler instrument is temporally-based (measuring particle flow through the probe volume) we will adopt the temporal

distribution function in the following discussion.

Consider the true particle flux through the probe volume. The number of particles entering the probe volume per imit

time in the phase space element dv may be written as

^;Ap(v.<l))/i(v.<}))d())dv (5)

tf where is the composite flux rate, /j(v,(t)) is the temporal joint distribution function in size and velocity (normalized to

unity), and dv and d^ are differential elements of phase space centered about the velocity v and size ((>.

Correspondence between the temporal and spatial approaches gives the number density in terms of temporal variables

as

p=X"t/S^ (6)

where 5^ is the spatial mean particle speed.

Composite probabilities and may be formed in terms of this distribution fimction by letting each class i

comprise the population of particles within a differential element d^ wide, centered about particle size and noting that the

probabilities of each class are independent. Under these conditions, the composite probabilities are given by the products of

the component probabilities, and sums over classes become integrals over the distribution fimctions:

n r. FT -Ki AV; Pi AVi .pAV^

{ {

'P' (8)

where the spatial mean probe volume AV^^ and temporal mean probe area Apj are given by

655
/.(v.(l))rfvd(t) (9)

Apt = jjAp{ymt(ym^<^

(|> V

(10)

and f is spatial size distribution function.

Note that the spatial filter probability does not depend upon the size or velocity class of the particle being considered

but only on the composite characteristics of the particle flow. As such, this filter rejects particles uniformly over the

size/velocity space of the particle flux, and therefore introduces no bias into the "measured" size/velocity distribution.

This is not true of the temporal filter, however. In this case the filter probability depends upon the residence time of

the particle under consideration. Since the residence time is correlated inversely with the particle velocity and directly with

the probe volume depth, the temporal filter will preferentially suppress the populations of particles with low velocities or

large component probe volumes and will therefore introduce a bias into the measured distribution function.

Armed with the expressions for and , we can now return to consideration of the relationship between the flux of

particles which is measured by our ideal system and the true flux of particles. Combining the expression for the true flux with

the expression for P^ and substituting for p in terms of temporal variables, the number of particles passing through the spatial

filter per imit time in <iv is given by

These are the particles which are eligible for measurement if they can pass through the temporal filter. The function

Pj gives the probability of passage through the temporal filter as a function of particle residence time in the probe volume.

As noted previously, this residence time will certainly depend upon the velocity and may also depend upon the particle size. However, specification of particle velocity and size alone is not sufficient to determine particle residence time. This is because particles with identical size and velocity but traveling through different parts of the probe volume will yield different

residence times since the depth of the probe volume is not constant but is a function of the position where the particle enters.

In order to accoimt for this distribution of residence times among the particles that populate dv d^, we introduce an

auxiliary distribution function of residence time, ^(T;v,(j)), with parametric dependence upon v and ()), and normalized to

unity. This distribution is used, in effect, to spread the distribution function /^(v,()>) into a new dimension t.

Combining this with P^ gives the number of particles per unit time which pass through the composite Poisson filter in

d\ about V, about ()), and dx about x

But we are only interested in the distribution of particles over size/velocity space. This may be recovered by

integrating over the particle residence time, which gives the number of particles passing through the composite filter per unit

time in c/v <f(|> as

X

For an ideal phase-Doppler system, subject only to the single particle constraint, this is the measured number of

particles per imit time indv d^ , which can be expressed as

X't A^(v.(t))€"^^^^'

^ ^^ft(v,<f)d^dv (11)

(12)

(13)

(14)

656
where is the measured particle rate and/^(v,<t>) is the measured joint distribution function with unit normalization. The

relationship between the measured and true distributions is then given by

4n(v.t» = -hf e^^^'''^%v.(^)J.-^^ V \(x;y,^)ch (15)

1

For convenience, we will return to using the notation Pj^ for the spatial filter probability and we introduce the mean

temporal filter probability for a particle of velocity v and size ^ as

tf)di (16)

The relationship between the measured particle rate and the true flux rate may be determined by integrating over v and

(j) and using the normalization constraint on ff^(y,^):

hm = X;P^//A^(v.(l))P/v.<t>y,(v.(t))dvd(t) (17)

0 V

The relationship between the measured and true distribution functions can best be expressed in terms of an efficiency

for each distribution. We define the Poisson efficiency as the number of particles measured per unit time in some element of

phase space divided by the true number of particles that passed through the probe volimie per imit time in that same element.

Using this definition, the joint distribution efficiency is given by

X^A^(v,(t))/^(v,(l))

Similar efficiencies can be defined for the size distribution and velocity distribution by considering the number of particles

measured, divided by the true number of particles. We denote the size and velocity efficiencies by Ti(())) and ti(v)

respectively.

ILLUSTRATION OF THE FILTER EFFECTS IN A UNIDIRECTIONAL FLOW

The effect of the Poisson filters on the data measured by an ideal phase-Doppler system may be illustrated by

considering a simplified system consisting of a unidirectional flow of particles through a constant-depth probe volume of

arbitrary cross section with area Ap , and depth L in the flow direction. The probe volume is the same for particles of all

sizes so that AVjj(())) = AV = L Ap. Under these conditions, the residence time of the particle is a function of velocity only and

is given by T = L / v. The spatial and mean temporal probabilities for this simplified system may then be written as

P^ = , P^(v,(t)) = Pt(y) = e

t ' (19,20)

where X.^ = "k^Ap.

It is apparent that the probability of passage through the spatial filter depends only on the total number density of the

flow and the probe volume. In fact, it is the product of these two, i.e., the expected value of the number of particles in the probe volume, that determines the filter probability. In a similar fashion, the temporal filter depends upon the expected value

of the number of particles in the probe volume during residence time i. Since both expressions have the same fimctional

form, the effects of both filters may be simunarized on a single plot as a function of either expected value.

657
Figure 2 shows the probability of rejection (1 - Pj^ or 1 - P^) as a function of either expected value. This figure shows

that for values as low as 10 \ a 10% rejection rate occurs, while for values just over 2, greater than 90% rejection will result. -2 Only if the expected value is kept below ~10 can the results be assiuned to be unaffected.

The joint distribution efficiency of our simpliHed flow is given by the product of the spatial and temporal filter

probabilities (one minus the values in Fig. 2). The size and velocity distribution efHciencies, however, depend upon integrals

of the single-particle temporal filter probability over some prescribed joint distribution function /f(v,(t)). In order to illustrate

these efficiencies, we use a joint distribution function representative of the data measured in the decelerating flowfield of a kerosene spray flame [3].

Figure 3a gives the measured size/velocity correlation of the spray and Fig. 3b shows the impact of the temporal

Poisson filter. In a decelerating flowfield like this, one expects to find that particles up to some maximum size will follow the gas phase behavior. Above this size droplets remain ballistic to the flow, increasingly so with size. As shown in the figure,

droplets of 10 |im and below are equilibrated with the gas phase while those above 10 (im have excess velocities. The dashed

lines in the figure give the mean velocity +/- the RMS, to indicate the width of the velocity distribution. (Note that for the

measured data, the RMS is essentially independent of size at this location.)

Figure 3b shows the size distribution efficiency T|(<|)) for various particle flux rates. Since the true number density of

the flow is not known, has been set equal to one. In this case, T\{<^) shows only the effect of the temporal filter on the

measured flow. These results were obtained by assigning a Gaussian velocity distribution (clipped at three standard

deviations) with mean and RMS as shown in Fig. 3a, and computing f\{(^) for flow through the 200 \xm probe volume depth.

3

The results show that for particle flux rates of the order 10 Is and below, there is little effect of the filter (a few percent at 4

low velocities). At 10 Is, however, more than 10% of the high velocity component of the spray is rejected, and 20% of the

equilibrated spray is lost. When the particle flux is increased to 10^ /s, almost 70% of the high velocity spray is lost, and more than 85% of the equilibrated spray is not accounted for. The dashed line in the figure indicates the filter behavior for

the experimentally measured particle rate of 3700 Is. Even at this low rate almost 10% of the spray is lost to the temporal

Poisson filter.

It should be noted that high particle flux rate is not the only condition xmder which the temporal Poisson filter can have a substantial impact on measured data. An alternative case of practical interest is that of intermediate flux rate but long

particle residence time. This may be achieved by considering the same decelerating flow system depicted in Fig. 3, but with a much lower carrier velocity.

The data shown in Fig. 3 were taken for an air-assist atomizer with mass-based atomizing-air to fuel ratio of one. The

result of this high atomizing air flowrate is that the carrier stream for this flow achieves a significant velocity (~10 m/s) even

as the spray decelerates. Instead, consider the case where there is no initial carrier stream, but the spray is into quiescent air. This may be simulated by reducing the velocity of the equilibrated particles (0 < 10 jim) from 10 m/s, as shown in Fig. 3a, to ~1 m/s, and then allowing the velocity to increase linearly to 20 m/s at 40 jlm. The result of passing a spray with this 3

characteristic through the temporal Poisson filter is shown in Fig. 4. Even at a very low particle flux rate (10 /s) a 4

substantial amount of the low velocity component of the spray is lost. At 10 Is, while high velocity particles have just

started to be affected, low velocity particles have already experienced an order of magnitude loss in numbers through the

filter. For the highest flux case shown (10^ Is), 2/3 of the highest velocity particles are rejected, while the low velocity

particles are virtually eliminated.

As mentioned in the introduction, the original motivation for investigating the effects of the single-particle

measurement constraint was the controversy surrounding the ability of phase-Doppler instruments to measure volume flux. Although there are many more effects which need to be considered in order to perfonn a complete analysis of the volume flux problem (e.g., probe volume corrections, etc.), the consideration of the Poisson filter effect provides a baseline analysis since no real system can out-perform the ideal system constructed here. The effect of the first stage of the Poisson filter (spatial) on the measured size distribution and volume flux is easy to

see. Since all classes of particles are individually, as well as collectively, Poisson distributed in space, the first stage filter

exhibits no preference for any particular marked class. Particles rejected by this filter are therefore distributed in the same

proportion as the true size/velocity distribution and as a result, no bias is imparted to the measured distributions. Similarly,

658
the loss of liquid volume flux is spread proportionately over the true distribution. The result is that when X% of the spray is

lost in the filter, X% of the volume flux is also lost

The effect of the second stage Alter (temporal) on the measuired size distribution and volume flux must be computed

for each individual case. The curves of Figs. S and 6 show the size distributions for the two cases previously discussed.

These were computed iising the measured size distribution from the kerosene spray flame and the size distribution efflciency

of Figs. 3b and 4. Note that in the case with the higher carrier velocity (Fig. S), there is little effect until the particle rate 5 3

approaches 10 /s, while in the more severe case (Fig. 6), signiflcant bias begins to appear even at 10 /s.

Also shown in these flgures is the fraction of the liquid volume flux that passes through the temporal filter. It is apparent that even in the milder of the two cases, a signiflcant fraction of the volume flux may be lost. Note that if the loss through the spatial filter were also included, these values would be reduced further.

SUMMARY

In this paper we have outlined the effects of Poisson particle statistics on the ability of a "single-particle" diagnostic to make accurate size and volume flux measurements. Although the analysis presented is quite general, a simplifled flowfleld

and probe volume have been used to illustrate its consequences. Even this simple example sufflces to show that the action of

the Poisson Alters places restrictions (possibly severe) on the quality of measurements by techniques such as phase-Doppler

anemometry. The keys to the Alters are the expected values of the number of particles in the probe volume pAVj^, and the

I* number of particles entering the probe region during residence time T, which is given by Kf Apf T . Only if these values are

-2 kept below order 10 can the measured joint distribution function and particle rate be assimied to reflect the true distribution

and rate of the flow.

ACKNOWLEDGEMENT

This work was performed at the Combustion Research Facility, Sandia National Laboratories, and was supported by

the U.S. Department of Energy, OfAce of Industrial Processes, Advanced Industrial Concepts Division.

REFERENCES

1. Edwards, C. F. and Marx, K. D.: manuscript in preparation, to be submitted to Atomization and Sprays, (1991).

2. Snyder, D. L.: Random Point Processes. Wiley, (1975).

3. Edwards, C. F., Rudoff, R. C, and Bachalo, W. D.: "Measurement of Correlated Droplet Size and Velocity Statistics, Size

Distributions, and Volume Flux in a Steady Spray Flame," Proceedings of the Fifth International Symposium on the

Applications of Laser Techniques to Fluid Mechanics. Lisbon, Portugal, (1990).

Fig. 1 Illustration of the spatial (a) and temporal (b) Poisson filter processes.

659
•au

u

.01 .02 .05 .1 .2 .5 1 pAV or Xth/v

Fig. 2 Probability ofrejection by the spatial or temporal

filters vs. expected number ofparticles in the probe volume.

15 20 25 30 35 40

Diameter, 0 (fim)

Fig. 3a Size/velocity correlation used to illustrate the effect of the Poissonfilters.

©-

1.0

0.9

0.8

0.7

0.6

0.5

lo'/s-

"S 0.4

g 0.3 s

0.2

0.1

0.0

Measured arriva^^\

rate (3700 /s)

10* /s

1 1 1 1 1 1 1 0 10 15 20 25 30 35 40

Diameter, (|> ()xm)

Fig. 3b Measuredfraction as afunction ofparticle sizefor

the size/velocity correlation ofFig. 3a.

©-

^—

/

d

o•3I

1.0

0.9 —

0.8

0.7 /

lo'/s-i

0.6

0.5 10* /s —

y

0.4

0.3

0.2 10* /s

0.1

0.0 1

1 J^^^ 1 1 1 0 10 15 20 25 30 35

Diameter, (|im)

Fig. 4 Measuredfraction as afunction ofparticle size for

the reduced carrier velocity case.

40

I

c

o3

s

s

True Distribution, 10 /s

Liquid Volume Flux K

Measured

(/s) Fraction

10' 0.99

10* 0.87

lO' 0.26

0.030

0.025

0.020 -

0.015

0.010 -

0.005

0.0000 5 10 15 20 25

Diameto- ((im)

Fig. 5 Measured and true size distributionfunctionsfor the

size/velocity correlation ofFig. 3a.

S

I

(fo

•ai

s

5

CO

0.030

0.025

0.020-

0.015 -

0.010

0.005 -

0.000

Liquid Volume Flux \

Measured

(/s) Fraction

10' 0.97

10* 0.78

10' 0.15

Diameter (|j.m)

Fig. 6 Measured and true size distributionfunctions for the

reduced carrier velocity case.

660
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 73

CLASSIFICATION OF ROUND LIQUID JET DISINTEGRATION IN

COAXIAL AIR STREAMS

Z. Farago* and N. Chigier^

DLR, German Aerospace Research Establishment, Hardthausen, Germany

^Mechanical Engineering Dept., Carnegie Mellon University, Pittsburgh, PA, U.S.A.

ABSTRACT

From an analysis of over one thousand high speed spark photographs, the modes of round

liquid jet disintegration in a coaxial air stream were classified over a liquid Reynolds number range

of 200 to 20,000 and over an aerodynamic Weber number range of 0.001 to 600. The observed

disintegration modes are compared to the modes of spherical water drop and thin water sheet

disruption in an air stream.

INTRODUCTION

In order to understand the basic mechanisms of free liquid jet disruption in a gas stream,

experiments on different types of research atomizers have recently been carried out at Camegie

Mellon University. Mansour and Chigier [1] examined the disintegration of a two-dimensional

liquid sheet between two parallel gas streams. Parametric experiments on coaxial airblast jet

atomization were carried out by Farago and Chigier [2] in which the aerodynamic Weber number

was varied over a range of six orders of magnitude, while the Reynolds number was varied over a range of two orders of magnitude. Using these experiments, intact liquid lengths of coaxial

atomizers were measured by Eroglu et al. [3].

After a thorough analysis of more than 1600 photographs of the author's experiments, a new

chart of morphological classification of the disintegration of round liquid jets was developed and is

presented in this paper. Different modes of liquid jet and spherical liquid drop disruption observed

by other investigators are compared with the present results.

The classifications of the modes of jet disruption developed by Ohnesorge [4], Miesse [5] and

Reitz [6] show the modes of atomization as a function of the Reynolds and Ohnesorge numbers.

However, these representations do not include the relative velocity between the gas and the liquid

stream for twin fluid atomizers. Other investigators like Wolfe [7] or Buschulte [8,9] used the

aerodynamic Weber number as a single variable for their presentations. Kennedy [10] presented

his experiments on the disintegration of spherical water drops in an air stream as a function of the

air velocity and the nondimensional number of Wea/Rcw^-^. Dimensional analysis of disintegration

of liquid jets in quiescent air, in parallel gas flow and of spherical drops in a gas stream shows that

at least four independent non-dimensional groups are needed to guarantee the similarity between

different experiments (Farago [11]). The analysis presented in this paper shows that jet

661
disintegration modes are a function of only two parameters: the liquid Reynolds number and the

aerodynamic Weber number. This presentation is limited to the aerodynamic atomization of liquid

with water and air both at ambient pressure and temperature.

EXPERIMENTAL PROCEDURE

The nozzle exit dimensions are shown in Fig. 1. The atomizer consists of a central water jet

surrounded by a coaxial air jet. Both water and air mass flows can be varied by changing their

supply pressures. Water is stored in a pressurized tank with a capacity of 10 liters. The (fiiving

force for the water supply is pressurized air. The highest water pressure which could be achieved

was 620 kPa. This water supply provides a flow, free of pulsations and vibrations, with a maximum emerging jet velocity of about 20 m/s. The inlet pressure of the atomizing air could be

varied from zero to 25 kPa, the latter providing an air velocity of about 180 m/s. The atomizer and

the exhaust system were connected to the same three-directional (z-r-phi) positioner so that the

axial distance between the nozzle exit and the exhaust inlet always remained the same (120 mm).

The accuracy of the positioner is 10 micrometer in the axial (z) and the radial (r) directions and 0.1

degree in the phi-direction. When the exit velocity of the atomizing air was kept at zero for

experiments on emerging round liquid jets in the "quiescent" air, the exhaust system generated an

environmental air velocity of about 0.5 m/s.

Photographs of jet disintegration and spray formation were taken with a Canon Fl camera

using a f=200 mm macro lens. Different extension tubes and bellows were attached to the camera

while taking photographs. An Electro-Optics EG&G 549 microflash was used to provide a 0.5

microsecond flash duration and 5x10^ beam candlepower light intensity.

EXPERIMENTAL RESULTS AND DISCUSSION

The water flow rate was changed from 0.65 to 52 kg/h, corresponding to a velocity range of

0.24 to 19.6 m/s and a Reynolds number range of 240 to 18,960. The air flow rate was varied

from quiescent conditions to 68 kg/h maximum. The maximum air flow condition corresponds to

an air velocity of 183 m/s and an air Reynolds number of 104,400. For the coaxial airblast jet

atomization, the lowest air mass flow rate was 8.5 kg/h with the corresponding air velocity and

Reynolds number of 22.9 m/s and 13,065 respectively. For aerodynamic atomization, the highest

aerodynamic Weber number was achieved by a low emerging water velocity of 0.95 m/s combined

with the maximum air exit velocity of 183 m/s. Correspondingly, the maximum aerodynamic

Weber number was 550. The lowest Weber number for coaxial airblast jet atomization was 0.4

corresponding to an air and water exit velocity of 18.2 and 22.9 m/s respectively. An ambient air

velocity of 0.5 rti/s was imposed by the exhaust system even when there was no air flow from the

atomizer. The lowest Weber numbers were achieved by emerging water jets into the quiescent air.

For the water velocity of 0.24 m/s, the corresponding Weber number was 0.001.

Detailed analysis of the photographs shows that jet disintegration can be divided into three

1) Rayleigh breakup. The dominant criteria for assigning jet disintegration for this category were:

i) the mean drop diameter is of the order of the jet diameter; the maximum drop diameter is

about twice the jet diameter, ii) The liquid jet disintegrates to drops without any liquid

membrane or liquid fiber ligament formation. This mechanism can be divided into two

10.36

Fig. 1 Sketch of the nozzle exit

categories:

662
subgroups such as: la) Axisymmetric Rayleigh breakup

lb) Non-axisymmetric Rayleigh breakup.

2) Jet disintegration via membrane type ligaments. The round jet develops into a thin liquid sheet

(membrane) which forms Kelvin-Helmholtz waves and breaks into drops as described by

Crapper, Dombrowski et al. [12]. Jet disintegration is completed within one or two wave

lengths. The diameter of the drops is considerably smaller than the diameter of the jet.

3) Jet disintegration via fiber type ligaments. Fibers are formed, peel off the jet, and the fibers

break into drops via the Rayleigh mechanism. The liquid core is accelerated and develops

waves. It may break into drops via the non-axisymmetric Rayleigh mechanism, or it may

break into ligaments. The ligaments may then break into smaller drops or ligaments by one of

the three disintegration modes. The completion of jet disintegration needs many wave lengths.

The drop diameter is an order of magnitude smaller than the diameter of the jet.

All three categories can be divided into two sub-modes:

A) Pulsating jet disruption as the normal sub-mode of atomization

B) Super-pulsating jet disruption which is connected to an extremely high periodical change

between low and high density regions in the spray.

More photographs and more details of this experiment are presented in reference 15.

Fig. 2 Disintegration of water jet in a coaxial air stream

663
Axisvmmetric Ravleigh Breakup

Studies of axisymmetric Rayleigh jet disintegration for water emerging into quiescent air were

made for Ohnesorge number 0.0037. According to Reitz's classification, the first-wind-induced

disintegration mechanism should appear at the liquid Reynolds number of Rcw = 1500. The

second wind induced disintegration mode should begin at Rcw = 5000. In contrast to Reitz's

classification, no significant change in the Rayleigh breakup from the first to the second wind

induced region was observed. The present experiments show the well known dependence of the

intact liquid length on the jet velocity which, for the first time, was described by Haenlein [13] at

similar experimental conditions.

The influence of gravity on the initially horizontal jet was clearly seen at low exit velocities.

Nevertheless, the shape of the jet, the ligaments and the drops were nearly axisymmetric for the

whole flow range. The influence of gravity on the jet flow direction is highly reduced for a horizontal water jet with a coaxial air stream, as can be seen in Fig. 2. Though this disintegration

mode is classified as the axisymmetric Rayleigh mechanism, the disintegrating jet may begin to

show a very slight tendency to form a helical axis. Nevertheless, the jet, the ligaments and the

drops are still nearly axisymmetric around this slightly helical axis. The jets with the corkscrew

shape before the jet disruption shown in Fig. 2 are typical examples of the first wind-induced

disintegration according to Reitz's classification. They are still assigned to the axisymmetric

Rayleigh mechanism in this presentation. The boundary between the axisymmetric and nonaxisymmetric Rayleigh type jet disintegration modes is poorly defined.

Non-Axisymmetric Rayleigh Breakup

The non-axisymmetric Rayleigh mode of jet breakup forms the boundary between the classical

Rayleigh breakup and the membrane-type breakup. Figure 3 shows photographs of this jet

disintegration. The liquid jet first accelerates rapidly in the high speed gas stream. Accordingly,

it's diameter is significantly reduced. The direction of the thin capillary liquid jet is influenced by

the interaction between the liquid jet and the turbulent structures of the surrounding air stream. The

wave formation on the round liquid jet may be caused by both the Kelvin-Helmholz instability and

the liquid-gas interaction. The aerodynamic forces reduce the jet diameter, but the breakup itself

may be due to the Rayleigh mechanism. The whole jet breaks into drops without intermediate

membrane formation or fibers peeling off This mode of jet disintegration may also be the final

step of the membrane type and the fiber type atomization. To understand this mechanism of round

liquid jet breakup, a combination is needed of the description of the jet deformation by

aerodynamic forces and the instability analysis of the deformed round liquid jet.

Membrane Tvpe Breakup

The dominant criteria for the membrane type breakup are: The whole round jet forms a thin

sheet (membrane) before breaking into drops. The liquid matter accumulates at the edge of the thin

sheet, forming a round liquid frame. The diameter of the round ft-ame is smaller than the diameter

of the emerging intact liquid jet. The liquid frame breaks into drops via the non-axisymmetric

Rayleigh mechanism. The mean drop diameter is one order smaller than the jet diameter. The

maximum drop diameter is also smaller than the jet diameter.

Figure 4 shows photographs demonstrating this type of liquid jet disintegration. The

morphology of the membrane type disintegration of round jets is very similar to that of thin liquid

sheets, previously described by Dombrowski and Mansour, among others. The very similar

disintegration mode of spherical liquid drops in an air stream is known in the literature as

"umbrella", "bag" or "hat" - type disintegration presented by Lane [14], Wolfe [7], Kennedy [10]

and others.

664
u^^, = 0 225 iii/.s

Fig. 3 Non-axisymmetric rayleigh breakup of water jet in a coaxial air stream

Fig. 4 Membrane type breakup of water jet in a coaxial air stream

665
Fiber Type Breakup

For many applications, such as the atomization of liquid oxygen in the space shuttle main

engine, the fiber type liquid jet breakup is the most important mode of liquid atomization. The

atomization begins with the formation of fibers and their peeling off the main liquid core. The

fibers break into droplets by the non-axisymmetric Rayleigh jet disintegration mode. In this stage

of the atomization, very small drops are produced. The diameter of those droplets is a few orders

of magnitude smaller than the diameter of the emerging liquid jet. Further downstream, the main

liquid core becomes wavy and breaks into ligaments. The length of these ligaments is about two to

five times the initial jet diameter. From the ligaments, new fibers are peeled off. The diameter of

the newly formed fibers increases with increasing axial distance from the nozzle exit. Even further

downstream, the ligaments form clouds of drops and liquid fibers, which leads to a periodic

change from high to low density spray. Generally, the size of the drops generated further

downstream is bigger than those formed closer to the nozzle exit. For calculation of drop size,

breakup time, and axial distance to complete the atomization, measurements need to be made of the

velocity field of the ligaments and of the surrounding gas stream.

Pulsating Disintegration Sub-Mode

All the photographs in Figs. 2 through 4 show the pulsating jet disintegration sub-mode which

can be considered as the "normal" sub-mode of jet disintegration. As can be seen in all the

photographs, the ligament and drop formation is connected to unstable wave formation of the

disintegrating liquid jet. Therefore, atomization is always a pulsating process.

Super-Pulsating Disintegration Sub-Mode

For air velocities of 183 m/s and liquid velocities between 1 and 6 m/s, photographs show the

"super-pulsating" jet disintegration sub-mode. This type of jet disruption appears when a low

liquid mass flow rate is associated with a very high flow rate of the atomizing gas. Reducing the

liquid flow rate which induces the change from a continuous liquid flow (in the quiescent air) to

dripping flow conditions can be considered an analogy to the change from the pulsating to the

super-pulsating sub-mode of airblast atomization. The intact liquid length is extremely short for

this sub-mode. In some cases, it amounts only to one or two jet diameters. Generally, the

periodic change of liquid flow rate, drop size and number density is significantly higher for the

super-pulsating sub-mode than for the "normal" pulsating sub-mode. When the air mass flow rate

increases, the corresponding liquid mass flow also increases at which point the super-pulsating

sub-mode changes to the normal pulsating conditions. Although the boundary between the two

sub-modes is not clearly defined, the tendency is clear and obvious. In the transient region from

the normal-pulsating to the super-pulsating sub-modes, the jet disintegration shows either the

morphology corresponding to the normal-pulsating mode or the shape of the super-pulsating

disruption. Especially at very low liquid flow rates, the spray pulsation, which is defined as the

ratio in number density or in liquid flow rate between the dense and the diluted spray, is extremely

high.

Classification of Jet Disintegration Modes

Figure 5 shows a plot of the different jet disintegration modes described above as a function of

the aerodynamic Weber number and the liquid Reynolds number at the nozzle exit. For each

experimental point in Fig. 5, four to ten photographs were analyzed. If one of the photographs

demonstrated different disintegration modes than the others at the same flow condition, both

observed modes were overlapped in Fig. 5. This case often occurred at the transition range

between two disintegration modes.

According to this chart, the axisymmetric Rayleigh type disintegration is the dominant mode of

jet disruption for the aerodynamic Weber number range of about We < 15. In the range of about

15 < We < 25, the non-axisymmetric disintegration mode is the most typical. However, at the

transition from the axisymmetric to the non-axisymmetric disintegration mode, the corresponding

Weber number slightly increases with the increasing Reynolds number. The membrane type jet

disintegration appears in the aerodynamic Weber number range of about 25 < We < 70. This range

666
u

«

JO

1«

I

a

n3.

o AiciBymmetric Rayleigh type disintegration o Non-axieyrametric Rayleigh type disintegration o Membrane type disintegration A Fiber type disintegration + Super—pulsating sub—mode

O

Membrane

type region

Aerodynamic Weber Number

Fig. 5 Modes of liquid water jet disintegration in a coaxial air stream

seems to reduce with the increasing liquid Reynolds number. At higher Weber numbers, the round

jet disintegrates by the fiber type disruption mode. If the ratio of the Reynolds number to the

square root of the Weber number is lower than about Rew / SQRT (Wea) < 100, the superpulsating sub-mode dominates for the jet disintegration.

CONCLUSION

The photographic studies show similarities between the disintegration modes of round water

jets in a coaxial air stream and the disintegration of water drops in an air stream previously

investigated by Lane [14], Wolfe [7], Kennedy et al. [10]. At similar aerodynamic Weber

numbers, the disintegration modes are also similar. Additionally, the disintegration of thin water

sheets, previously examined by Dombrowski [12], Mansour [1] and others, shows similar events

at similar Weber numbers. A comparison of the classification of round liquid jet disintegration

modes shows some disagreement with the first and second wind induced atomization classification

presented by Ohnesorge [4], Reitz [6], and others.

For technical applications, the most important mode of atomization is the "fiber-type" jet

disintegration. Spark light photographs show high pulsation for this mode of atomization,

especially for the flow condition corresponding to Rcw / SQRT (Wca) < 100. For the fiber type

atomization, the greatest amount of liquid is atomized in the secondary atomization region. Here

too, the biggest drops are generated. However, most theoretical studies deal only with the onset of

the instabilities and the primary atomization region based on wave instabilities at the axisymmetric

intact liquid surface. For calculations in the critical and most important secondary atomization

region, detailed measurements are needed of ligament size, ligament velocities, and relative

velocities.

NOMENCLATURE

di coaxial tube inner diameter [m]

dh air gap hydraulic diameter [m]

Oh Ohnesorge Number, Oh = Wem^-S / Rcw

Rca Reynolds Number of air, Rca = Ua (dh/Va)

Rcw Reynolds Number of water, Rcw = Uw (di/Vw)

Ua & Uvv air and water mean velocity [m/s]

667
Ur relative velocity between gas and liquid [m/s]

Wca aerodynamic Weber Number, Wea = Ur^ pa diMw

Wcm mechanical Weber Number, Wem = Uw^Pw di/cTw

pw air and water density [kg/m^]

Gw water surface tension [kg/s^]

Va & Vw air and water kinematic viscosity [m^/s]

ACKNOWLEDGEMENT

The authors wish to acknowledge financial support ft-om NASA Marshall Space Center, grant

number NAG8- 126.

REFERENCES

1 . Mansour, A., Chigier, N., Physics of Fluids. Vol. 2, No. 5, 1990.

2. Farag6, Z., Chigier, N., ASME Paper No. 90-GT-81, International Gas Turbine and

Aeroengine Congress, Brussels, Belgium, June 1990.

3. Eroglu, H., Chigier, N. and Farago, Z., Physics of Fluids A to be published, 1991.

4. Ohnesorge, W.V., Z.V.D.I., Vol. 81, No. 16, 1937.

5. Miesse, C.C., Ind. Eng. Chem., Vol. 49, No. 9, 1955.

6. Reitz, R.D., Ph.D. thesis, Princeton University, 1978.

7. Wolfe, H.E., Unclassified AD473340, U.S. Defense Documentation Center, Cameron

Station, Alexandria, VA, 1964.

8. Buschulte, W., DLR-FB 73-79, 1973.

9. Buschulte, W., ISABE, 9th Int. Symp. of Air Breathing Engines, Athens, 1989.

10. Kennedy, J.B., Roberts, J., ILASS-Americas, 4th Annual Conference, Hartford, CT, May

1990.

11. Farago, Z., Chigier, N., ILASS-Americas, 4th Annual Conference, Hartford, CT, May 1990.

12. Crapper, G.D., Dombrowski, N., Jepson, W.P., and Pyott, G.A.D., cited in An Album of

Fluid Motion. Parabolic Press, Stanford, CA, 1982.

13. Haenlein, A. ,Forsch. Geb. Ingenieurwesens A, Vol. 2, No. 4, 1931.

14. Lane, W.R., Ind. Eng. Chem.. Vol. 43, 1951.

15. Farago, Z. and Chigier, N., Atomization and Sprays. Vol. 2, 1992.

668
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 74

EXPERIMENTAL INVESTIGATION OF TWO-PHASE FLOW IN THE NEAR

FIELD OF AN AIRBLAST ATOMIZER

P. Hebrard, P. Trichet and X. Bardey*

ONERA/CERT/DERMES, Toulouse, France

*SNECMA, Centre de Villaroche, Moissy Cramayel, France

ABSTRACT

A detailed study of the two-phase flow produced by a gas turbine airblast atomizer is performed. Twocomponent Phase Doppler Interferometry is utilized to characterize the two-phase flow fields produced by the

atomizer: continuous and dispersed phases. Measurements of the mean axial and transverse velocities for each

phase are conducted. In addition, the droplet size distribution and the volume flux are measured. The results

demonstrate the influence that the spray has on the gas flow and also the swirl influence on the drop size

distribution.

INTRODUCTION

The spray behavior of an atomizer has a direct effect on the combustor performance. Detailed

measurements of both continuous and dispersed phases in liquid fiiel sprays are essential for a good

understanding of the physical processes of fiiel-air mixing and transport phenomena. The present work

addresses these questions by utilizing recently developed diagnostic techniques which provide nonintrusive

measurements of droplet size distribution, droplet velocity, droplet mass flux and dilute phase velocity inside a small measuring volume in the flow field. In addition, the experimental data obtained are to be utilized as a data

base for the development and validation of numerical codes [1].

The approach taken in this study is to apply nonintrusive diagnostic techniques to characterize an

airblast atomizer under nonreacting conditions. Measurements are obtained at three axial sections 4.5, 16.4 and

36.4 mm downstream of the atomizer and include mean axial and transverse velocities of the gas phase within

the isothermal spray, mean axial and transverse droplet velocities, local drop size distribution and volume flux.

EXPERIMENTS

Atomizer

The atomizer utilized in the present study is shown in Fig. la. The atomizer is of the airblast type and

is designed for use in gas turbine engines.

Test Conditions

The operating conditions selected for the present study are a baseline case corresponding to a "cruise"

operating level: the fuel flowrate is 4.1 g/s, and the pressure drop through the atomizer is 4 %. Water is employed as fuel simulation fluid in the present study.

669
a) Airblast Atomizer

b) Schematic of the spray test rig

Figure 1 Experimental set up for spray analysis

670
Facility

The atomizer is mounted in an adapter (airbox) which provides both fuel and atomizing air. The

particular configuration and orientation employed in this study are presented in Fig. lb.

Diagnostics

The anemometer used in this study is a two-components dual-beam Phase Doppler Particle Analyzer

(Aerometrics PDPA System 200) [2]. Two orthogonal probe volumes are provided by separating the 488.0 nm

and 514.5 rmi lines from an Argon laser. Data acquisition, processing and storage are conducted on a Compaq

deskpro 386/20 e Computer.

The PDPA provides spatially and temporally resolved measurements of drop size (D^^j, T^20' ^30' ^32^

velocity (mean, rms), correlations (size-velocity, velocity-velocity) and volume distributions. In addition, the

instrument is used to discriminate the two phases in the spray [3]. By optimizing the sensitivity of the

instrument to small particles and extracting particles which are small enough to track the flow, the continuous

phase velocity can be measured.

Measurements are made across three axial planes: 4.5, 16.4 and 36.4 mm. For each of these planes,

data are collected at each point of a grid (see Fig. 2b). Through most of the spay, 4000 data points are

collected for each location - Exceptions were made when the data rate was very low: in these regions, the flow was sampled for a minimum of 1000 seconds.

RESULTS AND DISCUSSION

Droplet Size Distribution

Surface plots showing the local variation of Sauter mean diameter SMD = DJ2 for each of the three cross sections are shown in Fig. 3. Close to the nozzle, the droplet size distribution exhibits a dome shaped

repartition. The inner and outer air flows of the airblast atomizer produce a near uniform drop size distribution,

but the air jets issuing from the cup holes produce a secondary atomization attested by the near circular

depression. Some large drops are observed in the core evidencing the fact that large drops are generated by the

irmer swirling air.

Further downstream, the distribution broadens as one moves along the spray. The swirling flow

recirculates the small drops back within the central region of the flow. Large indentations on the boundary of

the spray are due to the cup hole jets. At the lower section, the size distribution is more uniform on the

periphery of the spray. The spray cone angle is widened by the swirl effect and small drops are in the inner

core. These size distributions are characteristic of sprays with swirl [4].

Droplet Velocity

Surface plots showing the local variation of droplet mean velocity for each of the three cross sections are presented in Fig. 4 (axial velocities) and Fig. 5 (transverse velocities). Close to the nozzle, the axial

velocity distribution exhibits two concentric maxima; these peaks are due to the inner and outer swirling air. Further downsteam, these two distributions merge to form an unique axial flow field. At the lower position the

velocity distribution widens. Between the three cross sections, we can note increasing velocities indicating

droplet acceleration. In the core of the swirling flow, a size-velocity correlation exists.

Transverse velocity distributions shown in Fig. 5 indicate that, close to the nozzle, there are two

swirling flows; the near circular dip is due to the cup hole jets. Along the y = 0 axis, we have a diametral

evolution of tangential velocity and along x = 0 we have the radial velocity. Further downstream, the huge

swirl flow occupies the core of the spray and on the outer edge of the spray, the transverse velocity is nearly

671
a) Spray visualization by laser light sheet

-Y

+ + + + + + +

+ + + + + -L.«+ -f + + +

f A + + + + +

+/+ + + + + +

+ + + + + +

d

-ts\ ^ —'-icrl -ti OBl

~ r»Tio + + + + + +/+

"Tims

4- + + + + if

Z= 16.4 mm

b) Net of measurement locations ( grid si7e 5mmX5mm )

Figure 2 Shape of spray area in the spray cone

672
673
674
675
constant. At the lower position, we find a widening of the velocity distribution, and the swirling core is.still present. As for the axial velocity, the transverse velocity increases downstream. However, contrary to the axial

velocity, the transverse velocities demonstrate no significant dependence upon size, indicating little momentum

transfer between phases in this velocity component.

Droplet Volume Flux

The phase Doppler instrument can also be used to measure the volume flux of the fuel. It calculates the cross section of the measurement volume and utilizes size and velocity measurements to obtain the local volume

flux [5]. Surface plots showing the local variation of volume flux for each of the three cross sections are shown

in Fig. 6. Nozzles generally display asymmetries [6]. The swirl component of the flow produces a rotation of

this asymmetry in a "cork screw" manner in and out of the measurement planes. The swirling air has the effect

of widening the spray cone angle, showing low concentration of drops in the core.

Continuous Phase in the Presence of the Spray

The PDPA has been used to discriminate the gas and dispersed phases of the flow. Comparisons have

been made between the two phases, for the cross section located at 16.4 mm downstream of the atomizer (data

not shown). There are several interesting points which can be pointed out. The axial gas velocity is found to be

slightly higher than the drop velocity. The outside boundary of the spray presents indentations, proof of the cup

hole jets' interactions. On the transverse velocity, many peaks have been observed which can be attributed to the same cause. Only at the lowest position, we find negative values on the axis for the gas velocity, indicating

a recirculating flow.

Figure 7 presents a diametral scan of the mean axial continuous phase velocities. We can note that the

presence of the spray seems to reduce the width of the profile and also that, in this section, the reverse flow

near the axis cannot be measured.

CONCLUSION

Several measurements have been made throughout each of three cross sections in an airblast atomizer

spray. These results are presented as surface plots showing how the liquid volume flux, SMD and velocities vary

with position within each of these cross sections, corresponding to 4.5, 16.4 and 36.4 mm downstream of the

nozzle orifice.

This study demonstrates that substantial information can be extracted from the complex environment of

a spray, using non intrusive techniques. For the present atomizer and operating conditions it is observed that:

a) the conditions near the atomizer differ greatly from the conditions further downstream; b) the distributions of

size, velocity and volume flux are strongly influenced by the dynamics imposed by the swirling field; c) the

presence of fuel spray injection substantially changes the character of the flow field in the near field of the

atomizer; and d) between the three cross sections, the drop velocity increases with increasing axial distance,

indicating a momentum change.

Future work will be extended to the third velocity component, then leading to insight into : velocity of

selected size groups, estimated droplet flight path, momentum transfer from the gas to the drops, assessment of

droplet drag coefficients, and turbulence damping by the drop. These additional data are necessary for fiirther development and verification of two-phase flow numerical codes.

ACKNOWLEDGEMENTS

This work was supported by a SNECMA contract. We would like to thank Dr P. MILLAN for his

expert technical assistance. The participation of J. FORNE in the collection of data and development of data

reduction software is appreciated.

676
677
30.00 -

\ 23.00 -

16.00

9.00

2.00

3-

•5.00 , -35.00

+ SNECMA (LDA)

CERT (PDPA)

21.00 -7.00 7.00

RADIAL LOCATION (in.Ti)

21.00

Figure 7 Continuous phase mean velocities , downstream

location: 16.4 mm.

SNECMA LDA results: without spray

CERT PDPA results: with isothermal spray

35.00

REFERENCES

1. Hebrard, P., Lavergne, G., Trichet, P., "Experimental Study of Boundary Conditions for Spray

Systems in Combustion Chambers", Sixth Annual Conference ILASS Europe. Pise, Italy (July 1990).

2. Bachalo, W.D., Houser, M.J., "Development of the Phase/Doppler Spray Analyzer for Liquid Drop

Size and Velocity Characterizations", AlAA/SAE/ASME Twentieth Joint Propulsion Conference.

Cinciimati, Ohio (June 1984).

3. Mc Donnell, V.G., Samuelsen, G.S., "Application of two Component Phase Doppler Interferometry to the Measurement of Particle Size, Mass Flux, and Velocities in two phase-flows". Twenty second

Symposium on Combustion, pp. 1961-1971, The Combustion Institute, Pittsburgh, PA (1988).

4. Brena de la Rosa, A., Bachalo. W.D., Rudoff, RC "Spray Characterization and Turbulence Properties

in an Isothermal Spray with Swirl". ASME paper 89 GT 273, Gas Turbine and Aero Engine Congress. Toronto, Canada (Jime 1989).

5. Bachalo. W.S., Rudoff, R.C., Brena de la Rosa, A., "Mass Flux Measurements of a High Number

Density Spray System using the Phase Doppler Particle Analyzer", AIAA 88-236, AIAA Twenty Sixth

Aerospace Science Meeting. Reno, Nevada (January 1988).

6. Ros^ord, T.J., Russsell, S., "Influences on Fuel Spray Circumferential Uniformity",

AlAA/SAE/ASME/ASEE Twenty-third Joint Propulsion Conference. San Diego, CA (July 1987).

678
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 75

LIQUID SHEET INSTABILITY IN A COFLOWING AIR STREAM

H. Eroglu and N. Chigier

Mechanical Engineering Department

Carnegie Mellon University

Pittsburgh, PA, U.S.A.

ABSTRACT

Wave characteristics of a liquid sheet injected from a two-dimensional airblast atomizer were

investigated using high speed photography and frequency measurements. Photography was used

to determine the spatial wave amplitude growth rates and average wavelengths. Wave frequencies

were measured by means of a laser beam attenuation technique. Wave propagation speeds were

calculated from the measured average wavelengths and frequencies. The effects of liquid viscosity

and surface tension on the frequencies were investigated by mixing glycerine and Triton X-100

respectively with water. The average wavelengths decreased with both liquid mass flow rate and

air velocity. However, the dominant frequency increased with these parameters. The spatial wave

amplitude growth rate decreased with liquid mass flow rate and increased with air velocity.

INTRODUCTION

Study of the disintegration of a thin liquid sheet in a coflowing air stream can help in the

understanding of the basic mechanisms of atomization in practical systems without the added

complexity of a three-dimensional configuration.

Liquid sheet disintegration has been studied previously by several researchers [1-5]. More

recently, computational methods have been developed. Among the recent computational methods,

Lin and Lian [6] studied the instability of a viscous liquid jet in a gas, using a linear model. They

assumed that the disturbances on the sheet surface grow spatially, but they did not consider the

temporal growth rates.

Among the recent experimental investigations of liquid sheet instability, Lee and Yang [7]

measured sheet thickness variations both as a function of time and distance in the flow direction.

They performed their experiments on liquid films bounded by a solid surface at one side and a high

velocity air flow on the other. As a result, the liquid surface waves were found to be influenced

considerably by the air velocity. Aral and Hashimoto [8] investigated the disintegration of a thin

liquid sheet in a co-current air stream. The liquid sheet vibration and breakup frequencies were

observed to be equal to each other. The wave frequency and amplitude were correlated with Weber

and Reynolds numbers. Their frequency measurements revealed that the wave frequency increases

with liquid and air velocities. Mansour and Chigier [9] studied the instability of a liquid sheet

issuing from a two-dimensional airblast atomizer. They reported that the Uquid sheet vibrates in

the sinuous mode at low liquid velocities, and dilational waves are dominant at high liquid

679
velocities. At intermediate liquid sheet velocities, both sinuous and dilational modes are superimposed on the liquid sheet.

This paper describes the wave characteristics in terms of average wavelengths, spatial

amplitude growth rates, wave frequencies and propagation speeds for the same two-dimensional

airblast atomizer used in reference [9]. The effects of liquid viscosity and surface tension on the

wave frequencies are also investigated.

EXPERIMENTAL TECHNIQUE

Figure 1 shows the two-dimensional airblast atomizer. The liquid sheet emerges from a

central slit which is 0.254 mm wide and has an aspect ratio of 120:1. The air streams have a convergence angle of about 48 degrees and impinge on both sides of the liquid sheet. The slit

width for the discharging air streams is 1.45 mm. The two-dimensional atomizer was mounted on

an XY positioner with an accuracy of 1 mm in all directions. The liquid and air supply systems

provide a maximum liquid flow rate of 2270 kg/h and air pressure of 600 kPa.

Fig. 1 2-Dimensional Airblast Atomizer

A 4x5 inch plate camera was used with bellows extensions. This camera provides a narrow

depth of field and up to three times magnification of the disintegrating liquid on the film plane. An

EG&G 549 microflash was used with a 0.5 \is pulse duration. A back lighting technique was

found to be the most effective lighting arrangement for obtaining high contrast images. Four to

seven photographs of the liquid sheet were taken at each flow condition for determining average

values of wave amplitudes and lengths. The accuracy of these average values is estimated to be

within 10%.

Figure 2 shows the optical configuration for the frequency measurements. An Aerotech model

LS2P Helium-Neon laser with 2 mW output power is used as the light source. The oscillating

liquid sheet is positioned such that the line of the laser beam coincides with the y direction of Fig.

2. The photodiode generates a current proportional to the incident light intensity which varies as the liquid sheet crossed the laser light and causes its attenuation. The current from the photodiode

is converted into voltage by way of a simple operational amplifier feedback circuit. This voltage is

fed into a Nicolet digital oscilloscope to visualize the waveform. The analog voltage is also

transferred to an IBM PC/XT computer at the rate of 20 kHz via an analog-to-digital converter.

Fast Fourier transform analysis is applied to the data in order to determine the dominant

frequencies. Dominant frequencies are finally displayed in the frequency domain as shown in Fig.

3. The best signal quality was obtained when the laser beam was positioned close to the wave

crests as suggested by Mansour and Chigier [9]. The uncertainty of the frequency measurements

at low liquid flow rates was about 3% for a 95% confidence level. However, it increased to 15%

at very high flow rates, due to the broadening of the frequency band widths. The experiments at

high liquid flow rates were repeated four or five times, and the frequencies were averaged.

680
® 0

1) H9-Ne Laser

2) Collimalffig Lens

3} Nozzle

4) Optical Filter 5) Focusing Lens

6) Phoiodiode

71 Current to Vohage Converter

8) A-D Converter

9) IBM XT

Fig. 2 Frequency Measurement System

o

p.

PllOOl)

20000

P1000

22000

20000

18000

16000

14000

12000

10000

8000

6000

4000 -

2O0O

m = 1 1H kg'h. V ^ y) nl/s

Frequency (Hz)

Fig. 3 Typical Spectral Power Density at riiji = 1 18 kg/h, Vair = 58 m/s

RESULTS AND DISCUSSION

The wave amplitudes and lengths were measured from the photographs over an air velocity

range of 17.73 - 74.67 m/s and up to a maximum liquid mass flow rate of 450 kg/h. The wave

frequencies were measured using a laser beam attenuation technique. Wave propagation speeds

were calculated from the measured wavelengths and frequencies. Finally, the effects of liquid

viscosity (3-12 cps) and surface tension (32-55 dynes/cm) on the sheet wave frequencies were

investigated using solutions of glycerine and Triton X-100 respectively with water.

Wave Amplitudes and Lengths

Figure 4 shows the spatial wave growth for a liquid sheet of mass flow rate 81.76 kg/h, at an

air velocity of 17.73 m/s. Although both sinuous and dilational waves are superimposed on the

liquid sheet, the growth of the sinuous wave in the downstream flow direction is much larger. The

higher growth rate of sinuous waves was also reported by Hagerty and Shea [3]. Wave

amplitudes were obtained by measuring the distances from the sheet centerline to the crests of the

major disturbances as shown in Fig. 4. Figure 5 illustrates the variation of wave amplitude with

the distance from the nozzle exit. In this figure, the wave amplitude is normalized by the slit width

(t) for the emerging liquid sheet. Although the variation in Fig. 5 is approximately linear,

exponential or other higher order variations were also observed in some of the photographs. Since

the slope of the linear amplitude variation is constant, it was possible to define a constant spatial

amplitude growth rate for this pair of liquid and air flow conditions. The rate of increase of wave

681
Fig. 4 Amplitude Measurements from a Photograph

•d 4

Air Volocily. liquid Mass f low H. Va = 17.73 m/s m=81.76 Kg/li

Distance from nozzle exit, X (mm)

Liquid mass flow rale (kg/h)

, , ^ ^ Fig. 6 Variation of Average Wave Growth Rate

Fig. 5 Variation of Wave Amphtude along the Jet "

^^^^ liquid Mass How Rate

Axis (Vair = 17.73 m/s, m^ = 81.76 kg/h) (Vair = 17.73, 30.90, 43.35, 64.67, 74.67 m/s)

amplitude with distance downstream is plotted as a function of increase in liquid mass flow rate

and air velocity in Fig. 6. This figure shows that the spatial amplitude growth rate decreases with

liquid mass flow rate and increases with air velocity. High spatial amplitude growth rates occur

when the liquid inertia is low at low liquid flow rates. At that time, the specific energy of air per

unit volume of liquid is high, and this energy is sufficient to cause sinuous wave formation with

rapid amplitude growth. The relative velocity between the air and the liquid increases with

decreasing liquid mass flow rate when the air velocity remains unchanged. TTie disturbances are known to be amplified faster when the aerodynamic force increases at the air-liquid interface due to

higher relative velocities [101. Figure 6 also illustrates that the spatial amplitude growth rate

increases with air velocity when the liquid mass flow rate is kept constant. At very high liquid

flow rates, dilational surface waves were observed on the liquid sheet surface with insignificant

amplitude growth. The specific energy of air per unit volume of liquid at these test conditions was

not enough to cause any sinusoidal oscillations, especially at low air discharge velocities. At

intermediate liquid mass flow rates, both sinuous and dilational waves were superimposed. With

increasing downstream distance, the sinuous waves grew faster. As a result, more gradual

variation of spatial amplitude growth rates occurred at the intermediate liquid mass flow rates.

The disturbances on the liquid-air interface were found to be rather irregularly spaced. The

spacings between the crests of the waves were measured using the same method as that of Eroglu

682
and Chigier [1 1] for the round jets from airblast coaxial atomizers. These distances were accepted

as the wavelengths. Since the wavelengths vary in the axial (x) direction, average wavelengths

were calculated and normalized with respect to the central slit width (t) for the liquid phase. The

normalized average wavelengths were plotted as a function of liquid mass flow rate at four

different air velocities as shown in Fig. 7. Average wavelengths were found to decrease with

increasing liquid mass flow rate and air velocity. This decrease is more pronounced at lower liquid

mass flow rates, where the liquid sheet develops predominantly sinuous waves with large

wavelengths. The more gradual decrease or negligible change of wavelengths at high liquid mass

flow rates occurs when the dilational waves are the dominant disturbances. The dilational wave

development is usually accompanied by turbulent fluctuations in the liquid sheet. It seems that the

small scale turbulent structures inside the liquid sheet control the initial phase of the wave

development at high liquid mass flow rates and cause short wavelength surface disturbances to

develop.

Wave Frequencies

A preliminary survey was performed in order to determine the optimum location for

measurement of the wave frequencies in the downstream flow direction (x) of the liquid sheet.

Figure 8 shows the result of this survey at the liquid mass flow rate of 56.78 kg/h, and air velocity

of 26 m/s. The dominant wave frequency was measured as 400 Hz at this pair of flow conditions

and did not change along the downstream flow direction. However, it was interesting to notice

that the maximum spectral power density (spd) first increased to a maximum at about 7 mm

distance from the nozzle exit and then started decreasing. A comparison of Fig. 8 with the

photographs taken at similar flow conditions indicated that the peak of the curve corresponds to the

position of maximum wave amplitude before the start of sheet disintegration. The maximum spd

continued decreasing in the downstream flow direction until the disintegration was completed. The

initial increase of maximum spd at upstream locations is a direct result of the emergence of a

Axial Distance (mm)

L.qu.d mass How rate ( kg/h

) p.^ ^ SpcctTal Powcr Dcttsity Variatlon along

Fig. 7 Variation of Normalized Average jj^^ j^j yj^^is

Wavelength with Liquid Mas Flow Rate

(Vair = 17.73, 30,90, 43.35, 64.67 m/s) (Vair = 26 m/s, = 56.78 kg/h, f = 400 Hz)

dominant frequency, and the disappearance of the secondary modes. Since the wave frequency

does not change in the x direction, the optimum measurement location is selected for signal clarity.

Figure 9 shows the oscillation frequencies of the liquid sheet as a function of liquid mass flow

rate at six different air velocities. In general, the wave frequency increases with liquid mass flow

rate and air velocity. This behavior is consistent with the results of Lee and Yang [7] and Aral and

683
Hashimoto [8] on the liquid sheet instability with co-current air flow. Further analysis of Fig. 9

and the frequency jumps separating the regions A, B, and C for the different oscillation modes are explained in detail by Mansour and Chigier [9].

The results presented in this paper were obtained from the instantaneous photographs and

frequency measurements from points in a fixed (Eulerian) reference frame for a naturally vibrating

liquid sheet in a coflowing air stream. The photographic results gave average wavelengths and

spatial amplitude growth rates although it was not possible to identify the amplitude growth

corresponding to an optimum wavelength within a band of several wavelengths. The frequency

measurements in a fixed (Eulerian) reference frame provided the number of waves passing through

a point per second. Since it is very difficult to measure the temporal amplitude growth rates in a

moving (Lagrangian) reference frame with the same velocity as the liquid sheet velocity, this study

only provides information for the analysis of spatial instability.

Wave Propagation Speed

The wave propagation speeds were calculated by taking the product of average wavelengths

and the wave frequencies. Figure 10 shows the variation of wave propagation speed with liquid

sheet velocity at three different air velocities (30.90, 43.35, 64.67 m/s). Values are less than 1 m/s

for liquid sheet velocities lower than 15 rn/s. The wave propagation speed seems to be insensitive

to both liquid and air velocity variations. This behavior is consistent with the wavelength decrease

shown in Fig. 7 and the corresponding frequency increase shown in Fig. 9.

Ail Vo(ociiies

o 30.90 ni/s

43 35 ni/s

A 55 00 m/s

« 64 67 m/s

98 07 m/s

o 119 62 m/s

}^ c

/

/

/ V.

900 o o 00 00 °o

Air Velocities

o Vair=30.90 tWs

Vair=43-35 nvs

Vair=84.87 m/s

Liquid Mass Flow Rate (kg/h) Liquid sheet velocity (m/s)

Fig. 9 Liquid Sheet Wave Frequency versus Fig. 10 Wave Propagation Speed versus Liquid Mass Row Rate, [9] Liquid Sheet Velocity

(Vair = 30.50, 43.35, 55.0, 64.67, 98.07, 1 19.62 m/s) (Vair = 30.90, 43.35, 64.67 m/s)

Effects of Liquid Viscosity and Surface Tension

The viscosity of the liquid phase was varied by adding 96% pure glycerine obtained from

Dow Chemical Company into water at several volume percentages. The volume percentages of

glycerine-water solutions were 66, 60, 40 and 0% respectively. The corresponding liquid

viscosities changed from a maximum of 12 centipoise (66% glycerine) to a 3 centipoise (pure

water) as measured by using a Brookfield viscometer.

The results of this investigation are shown in Fig. 11. The graphs show that the dominant

wave frequencies decrease with viscosity. This decrease is more pronounced at higher liquid mass

flow rates and air velocities. Such a result is expected since the viscosity is generally accepted to

dampen the wave motion and give rise to increases in the wavelengths which are normally

associated with frequency decrease.

684
Liquid mass How rale. Air valoctly

m = 56 78 Kg/hr. V =. 26 nvs

• m = 90 85 kg/hi. V » 32 nvs

m= 118.)0k9«ii,V = S8m/s

C (<:p«)

Fig. 11 Effect of Viscosity Variation on Liquid Sheet Wave Frequency

Triton X-100 non-ionic surfactant was added to water in order to vary its surface tension. The

surface tension values of the liquid were obtained using the Du Novy ring method for different

amounts of surfactant addition. The frequency measurements were made for the surface tension

values 55, 50, 43, 41, 36 and 32 dynes/cm. The same liquid and air flow conditions as those for

the viscosity investigation were selected. Frequency measurement uncertainties are estimated to be

less than 5% for a 95% confidence level. The frequency measurement results shown in Fig. 12

indicate that the effect of surface tension on the dominant wave frequencies is insignificant for the

tested liquid and air flow conditions.

Liquid mass flow rale. Air velocity

a m 56 78 kgmr. V - 26 ITVS

• m • 90.es kgmr. V=32 m/s

m> IIS IOkgAir. V»S8m's

O (dyne/cm)

Fig. 12 Effect of Surface Tension on Liquid Sheet Wave Frequency

CONCLUSIONS

Wave characteristics of a liquid sheet with impinging air streams on the two opposite

interfaces were investigated by means of high speed photography and frequency measurements.

685
Spatial amplitude growth rates were found to decrease with liquid mass flow rate and increase

with air velocity. Average wavelengths decreased with increasing liquid mass flow rate and air

velocity. Small scale turbulent structures in the liquid sheet are probably responsible for the short

wavelengths encountered at high liquid mass flow rates.

The dominant wave frequency of the vibrating liquid sheet increased with both liquid mass

flow rate and air velocity. It decreased with increasing viscosity. The frequency appears to be

insensitive to the variation of surface tension from 32 to 55 dynes/cm.

The wave propagation velocity was in general less than 1 m/s, and there was litde change with

liquid and air velocity variations.

ACKNOWLEDGEMENTS

The authors would like to thank Mr. Eric Dion for performing some of the frequency

measurements. The financial support from NASA Marshall Space Center (grant No. NAG8-126),

and the atomizer from Parker Hannifin Corporation are gratefully acknowledged.

REFERENCES

1. York J.L., Stubbs, H.F. and Tek, M.R., "The Mechanism of Disintegration of Liquid

Sheets", Trans. ASME. Vol. 75, pp. 1279-1286, 1953.

2. Squire, H.B., "Investigation of the Instability of a Moving Liquid Film", Br. J. Appl. Phys. . Vol. 4, pp. 167-169, 1953.

3. Hagerty, W.W. and Shea, J.F., "A Study of the Stability of Plane Fluid Sheets", J. Appl.

Mech.. Vol. 22, No. 4, pp. 509-514, 1955.

4. Dombrowski, N. and Johns, W.R., "The Aerodynamic Instability and Disintegration of

Viscous Liquid Sheets", Chem. Eng. Sci. . Vol. 18, pp. 203-214, 1963.

5. Eraser, R.P., Dombrowski, N. and Routley, J.H., "The Atomization of a Liquid Sheet by an

Impinging Air Stream", Chem. Eng. Sci. . Vol. 18, pp. 339-353, 1963.

6. Lin, S.P. and Lian, Z.W., "Mechanisms of the Breakup of Liquid Jets", AIAA Journal. Vol.

28, No. 1, pp. 120-126, 1990.

7. Lee, G.Y. and Yang, O.Y., "Mechanism for the Disintegration of Liquid Film Surface Waves

by a Parallel air Flow", proceedings of the 4th International Conference on Liquid Atomization

and Spray Systems, pp. 193-199, Sendai, Japan, 1988.

8. Aral, T. and Hashimoto, H., "Disintegration of a Thin Liquid Sheet in a Cocurrent Gas

Stream", proceedings of the 3rd International Conference on Liquid Atomization and Spray

Systems, London, pp. VIB/1/1 - VIB/1/7, 1985.

9. Mansour, A. and Chigier, N., "Dynamic Similarity Between Liquid Sheets and Hard Spring

Systems",, submitted to Phvsics of Fluids. 1990.

10. Lefebvre, A.H., Atomization and Spravs. Hemisphere Publishing Company, New York,

1989.

1 1 . Eroglu, H. and Chigier, N., "Wave Characteristics of Liquid Jets from Airblast Coaxial

Atomizers" submitted to Atomization and Spravs. 1990.

686
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 76

SPRAY GAS-PHASE INTERACTION DOWNSTREAM OF A CO-AXIAL

COUNTER-SWIRLING DOME SWIRL CUP

H.Y. Wang, W. Sowa, V.G. McDonell and G.S. Samuelsen

UCI Combustion Laboratory

Dept. of Mechanical and Environmental Engineering

University of California

Irvine, CA, U.S.A.

ABSTRACT

The flow field downstream of a 3x scale GE CFM56 combustor swirl cup is

characterized with and without the presence of a spray by using phase Doppler

interferometry . The gas-phase mean axial velocities reveal a closed, on-axis

recirculation zone in each case. For the case without the spray, the

recirculation zone is comparatively larger. Anisotropic turbulence persists

within the regions where the measurements were conducted. Velocity statistics

are determined for the single-phase (gas phase in absence of spray), the

continuous-phase (gas phase in presence of spray), and the discrete-phase

(liquid phase in presence of spray). Differences between the cases are

revealed and discussed. Overall, the effect of the spray on the gas-phase

mean velocity is relatively small and can, in the present case, be reasonably

neglected in modeling the flow.

INTRODUCTION

Co-axial, counter-swirling air streams have been investigated to

understand differences in combustion characteristics relative to co-swirl

combustion [1,2,3,4]. Isothermal, co-axial, counter-swirling systems, for

example, have been studied by several investigators [2,3,4] who suggest that

co-swirl does not produce a recirculation zone. Other investigators find that

a recirculation zone exists for both co-swirl and counter-swirl [5,6]. The

different conclusions are attributed to different swirl number and other

factors (e.g., the method employed to generate swirl since the swirl number is

not the only parameter that characterizes swirling flow)

.

Droplet laden swirling flows are not well understood. Some efforts have

been conducted to characterize sprays with different primary air swirling

conditions [7,8,9,10], but none investigate the effect of a spray on the gasphase flow field. In the research reported herein, a complex flow field

generated by a practical hardware is characterized for both the single-phase

and the two-phase flow, and the effect of the spray on the gas-phase is

identified.

687
The purpose of this paper is to establish the effect of the spray on the

gas-phase flow field for a liquid spray injected into a swirl cup air streeun.

EXPERIMENT

Facilities

The 3x model combustor swirl cup is shown schematically in Fig. 1. A

90° cone-angle simplex nozzle is mounted in the center of the swirl cup. The

venturi separates the primary and secondary swirlers. The primary swirl

stream rotates clockwise looking from the top of the nozzle, while the

secondary swirl stream rotates counterclockwise.

radial velocity xao, y=0, r=160 mm

Fig. 1 Swirl cup and velocity histograms

(1) Simulated fuel nozzle

(2) Primary swirler component

(3) Primary venturi

(4) Secondary swirler component

(5) Conical sleeve

The swirl cup was cemented onto an aluminium plate to provide symmetric

boundary conditions to the flow field. The aluminum plate, in turn, was

fastened to a PVC cylindrical plenum. A 0.006 m cell-diameter polycarbonate

honeycomb (0.106 m thick) was placed 0.051 m above the top of the swirl cup to

provide a uniform velocity profile at the entrance plane to the swirlers.

Water was used to simulate the fuel.

688
A two-component phase Doppler interferometer (Aerometrics PDPA Model

3100-s) was used to measure (1) the single-phase velocities, (2) the

continuous-phase velocities in the presence of the spray, and (3) the size and

velocities of the discrete-phase. Table 1 provides details of the

configuration used in the present study. A TSI Model No. 3906 six- jet

atomizer was used as a seed generator. A 10% salt solution was used to

provide adequate seed yielding a satisfactory signal at a laser power of 1.0

W. In the presence of the spray, the continuous-phase and discrete-phase

velocities were measured separately. The continuous-phase velocities measured

in the presence of spray were determined by isolating the statistics to

droplets ranging in diameter from 2.44-4.14 ;jm, which was the smallest size

class detected by the PDPA. These droplets were assumed to follow the gasphase.

100

TABLE 1. Characteristics of PDPA 3100-S

Transmitter

0.5145^m line (U, D)^

Fringe Spacing

Waist

0.4880^;m line (V or W)

Fringe Spacing

Waist

Receiver

1000 mm f/9.3 Collection Lens

238 mm Focusing Lens

100 ^m Spatial Filter

30° Off-Axis Forward Collection Angle

2

U, V, W and D are axial, radial, tangential

velocity and droplet diameter respectively.

9.88 }jm

187.17 fjm

9.84 ^m

177.53 /im

80-

E^ 60

2 40

20

0

150

o Z= 51.2mm

• Z= 80.5mm

A Z = 1 10.0mm

6^ ''Um»m*»

£ 100

CM

Q 50+

0

-200 -100 0 100

Y mm

200

Fig. 2 Droplet size profiles

Test Conditions And Test Locations

Both the two-phase tests and the single-phase tests were conducted at an

air flow rate 0.154 kg/s. The liquid flow rate was 0.010 kg/s for the twophase tests. The pressure drop across the nozzle discharge was 827 kPa. For

the same air flow rate at the exit plane of primary cup, the swirl number was

0.47 for the primary cup and -0.55 for the secondary cup [6]. Measurements

were conducted at three axial locations (Z/Rp = 1.75, 2.75 and 3.75 where the

radius of the exit plane of the primary swirler cup, Rp, is 0.029 m and the

origin is the exit plane of the primary cup). In addition, data were acquired

along the center line to study the change in the length of the recirculation

zone

.

DISCUSSION OF RESULTS

The results are divided into two groups: discrete-phase and gas-phase

measurements in the presence (continuous-phase) and absence (single-phase) of

the spray.

689
Discrete-Phase Results

One feature of the swirl cup design is to provide uniform and fine

droplets. A portion of the droplets issue from the nozzle and convect

downstream directly and a portion impinge onto the inner surface of the

primary swirl component to form a liquid film which is then atomized due to

the strong shearing of the primary and secondary air streams at the edge of

the primary cup. Fig. 2 shows the length mean diameter D^g Sauter

mean diameter 032* These distributions indicate the presence of small

droplets along the centerline. At the periphery of the spray, D^g D32

increase sharply. This is attributed to (1) larger droplets from the atomizer

that have sufficient radial momentum to penetrate through the swirling air

streams, and (2) the breakup of liquid film at the edge of swirl cup that can

produce large droplets with a substantial radial momentum.

Continuous-Phase and Single-Phase Results

In this section, the gas-phase is compared with and without the spray.

In particular, the continuous-phase flow field in the two-phase test is

compared to the single-phase test. Both mean and fluctuating components are

considered for the axial, radial, and tangential components.

Mean Velocity Comparison . As shown in Fig. 3, the continuous-phase axial

velocities are generally similar to those of the single-phase with one notable

exception. In the recirculation zone, the continuous-phase axial velocities

are less negative indicating a recirculation zone of reduced strength. The

axial momentum transfer from the discrete-phase to the continuous-phase is

likely responsible in the recirculation region as droplets propelled

downstrecun oppose the reverse flow of the continuous-phase.

30

20

10

0

en -10 \-20

30

20 0

_i

Ld 10

>

_l 0

<

X

-10

< -20

< 30

LaJ :2 20

o Continuous— phase

• Single— phose

po o 8 e

o

I Z=1 10.0mm 'SSsggS*

80.5mm

o

8 ^

8 8

10

0

•10

-20

Z= 51.2mm

8 t„

v

-200 -100 0 100 200

Y mm

>-O

O

_l

UJ>

_J

<

<

<UJ

25

15

5

-5

-15

-25

25

15

5

-5

-15

-25

25

15

5

-5

-15

-25

o Continuous— phase

• Single— phase

Z = 1 10.0mm

°8 %•

Z= 80.5mm

•,oo8?*'

••#99'

Z= 51 .2mm

%

/

. ^ V

0

• >

-200 -100 0 100 200

Y mm

Fig. 3 Gas-phase mean axial velocity Fig. 4 Gas-phase mean radial velocity

690
The radial velocity data in Fig. 4 also reveal similar behavior between

the continuous-phase and single-phase.

Fig. 5 presents the comparison of mean tangential velocities at two

axial locations. The single-phase shows generally higher mean tangential

velocity. In the two-phase flow, most of the droplets that issue from the

nozzle, unlike the continuous-phase, do not have initial tangential velocity.

Their tangential velocity increases due to (1) the tangential momentum

transfer from the continuous-phase to the discrete-phase, and (2) the liquid

film breakup at the edge of the primary swirl cup. Therefore, the continuousphase tangential momentum is suppressed relative to the single-phase case.

20

'°

O

LjJ 5-20 ^ 20

I

10

1 0

1-10

-20

"Continuous- phase

•Single — phase

Z= 80.5mm

o8

Z= 51.2mm

• » ooo

•

-200 -100 0

X nnm

1 00 200

Fig. 5 Gas-phase mean tangential

velocity

20

15

10

5

o Continuous- phase •10-20 um

A20-30 fj,m * 30-40 //m

51-59 Aim *84S

73-88 Mm°.*r,

-200 -100 0 100 200

Distance from the centerline mm

Fig. 6 Rms of discrete-phase velocities

Fluctuating Velocity Comparison . To aid in interpretation of the

fluctuating velocity results, the profiles of the fluctuating velocity

components of the continuous-phase are presented in Fig. 6 along with the

"fluctuating velocities" of various size classes at axial location z = 51.2

mm. Note that the relative fluctuating velocity differences between the

continuous- and discrete-phase change radially.

Fluctuating gas-phase axial velocities are compared in Fig. 7. At the

first axial location, note the higher magnitude for the continuous phase

within the recirculation zone. This is attributed to the relative magnitude

of the fluctuating velocities between the continuous-phase and large droplets

(over 30 pm) as shown in Fig. 6. The continuous-phase velocity fluctuations

are enhanced when the velocity "fluctuation" of the large droplets is greater

than that of the continuous-phase. The greater velocity fluctuation of the

large droplets results from a variation in their ballistic injection instead

of the turbulent fluctuation in the common sense and is more dominant, as

shown, close to the injection place.

691
The radial and tangential velocity fluctuations for the continuous phase

are consistently lower than the corresponding single-phase measurements.

Again, the continuous-phase radial and tangential velocity fluctuations are

consistently greater than those of the large droplets and thus dampened (Fig.

6).

For the single-phase case, the fluctuating radial velocity reveals a

two-peak distribution at the axial location Z = 51.2 mm (close to the exit

plane of the swirl cup) (Fig. 8). The double peak distribution is suppressed

for the two-phase case. This is attributed to a momentxim exchange between the

gas-phase and the droplets, which have a single peak distribution and deanpen

the two-peak structure. Within the recirculation zone, both cases show a

similar behavior.

20

15

10

CO

^ 0

20 O

< 10

X

5

^

0

20

§15

^10

5+

0

o Continuous- phase

• Single— phase

Z=1 lO.Omnn

8 -9»«§«»88

Z= 80.5mm

Z= 51.2mm

/o 8 A-8 •

° 'o «

200 •100 0

Y mm

100 200

Fig. 7 Rms of gas-phase axial

velocity

20

15

w 10

E 5 ^ 0 o

0 20

_iUJ> 15

<

5

o

1

° < 20

0 15

_i^ 10

5

0

o Continuous— phase

• Single— phase

Z = 1 10.0mm

Z= 80.5mm

Z= 51 .2mm

8°°° '*

-200 -100 0

Y mm

100 200

Fig. 8 Rms of gas-phase radial

velocity

The greatest differences in gas-phase tangential velocity fluctuation

between the single-phase and continuous-phase is observed in and near the

shear layer (Fig. 9). This corresponds to the location at which the greatest

difference in tangential velocity fluctuation between the continuous-phase and

large droplets (over 30 ^m) occurs (Fig. 6). The degree of this difference

decreases as flow develops downstream due to mixing and entrainment.

Distribution along the Centerline . The mean and fluctuating gas-phase

axial velocity distributions along the centerline are presented in Fig. 10.

The recirculation zone closes closer to the inlet plane in the two-phase case.

Moreover, differences in the mean axial velocity between the single-phase and

the continuous-phase decreases downstreeun and coincides well with the decrease

in mean axial velocity difference between the continuous-phase and the larger

droplets [11]. This supports again the arguments used to interpret the mean

and fluctuating velocity differences between the single-phase and the

692
continuous-phase. The greatest fluctuation is observed at about Z = 160 mm

near the stagnation point of the recirculation zone.

VI \

E

(Jo_iUJ>

_l<

UJ

<

I-o

20

15

10

5

0

20

15

10

5

0

-200

"Continuous— phase

•Single-phase

Z= 80.5mm

% ooo

Z= 51 .2mm

o o

100 0

X mm

1 00 200

30

20

10

0

-10

-20

20

15

10

54

0

o Continuous — phase

• Single-phase

Along the centerline

0

o

o •

8*88« e • * * '

•

50 100 50 200 250

Z mm

Fig. 9 Rms of gas-phase tangential

velocity

Fig. 10 Mean and fluctuating gas-phase

axial velocity along the centerline

single-Phase and Discrete-Phase Results

Both the single-phase and the discrete-phase histograms are shown

respectively in Figure 1 for the axial velocity in the shear layer, the

tangential velocity in the shear layer, and the radial velocity around the

stagnation point. Velocity bimodal distributions are evident for the

discrete-phase at these locations, which reflects the flow field

characteristics superimposed on the gas-phase that are not present in the

single-phase case. The well defined bimodal velocity distributions of the

discrete-phase can be associated with either significant size-velocity

correlations, or intermittent flow field characteristics, or both. This

behavior which is pronounced in discrete-phase measurements is negligible or

relatively insignificant in the single-phase case. As a result, the two-phase

flow field in this practical hardware is both intriguing and complex.

CONCLUSIONS

(1) The differences in mean and fluctuating velocities between the singlephase and the continuous-phase correlate with differences in mean and

fluctuating velocities between the continuous-phase and the larger

droplets (over 30 ^m)

.

(2) The turbulent fluctuations for both the continuous-phase and the singlephase are anisotropic. The effect of the spray on the gas-phase mean

velocities is surprisingly small in this complex, two-phase flow.

(3) The two-phase case reveals a more intermittent behavior than the singlephase case at selected locations in the flow field such as in the shear

layer and at the stagnation point of the on-axis recirculation zone.

(4) The two-phase flow field and phase interaction in a practical hardware

can be complex. Intermittency, for example, plays an important role and

suggests, for example, the need for detailed transient analyses.

693
ACKNOWLEDGEMENTS

The authors acknowledge the financial support from GE Aircraft Engines,

and the assistance of C.T. Brown with testing protocol and H.D. Crum with

facility development.

REFERENCES

1. Yetter, R.A. and Gouldin, F.C., "Exhaust Gas Emissions of a Vortex

Breakdown Stabilized Combustor", Fall Meeting of the Western States

Section of the Combustion Institute (1976).

2. Vu, B.T. and Gouldin, F.C., "Flow Measurements in a Model Swirl Flow",

AIAA Journal, Vol. 20, pp. 652-659 (1982).

3. Habib, M.A. and Whitelaw, J.H., "Velocity Characteristics of Confined

Coaxial Jets with and without Swirl", Journal of Fluid Engineering, Vol.

102, pp. 47-53 (1980).

4. Gouldin, F.C., Depsky, J. and Lee, S.L., "Velocity Field Characteristics

of a Swirling Flow Combustor", AIAA Paper 83-0314 (1983).

5. Samimy, M. and Langenfeld, C.A., "Experimental Study of Isothermal

Swirling Flow in a Dump Combustor", AIAA Journal, Vol. 26, No. 12, pp.

1442-1449 (1988).

6. Mehta, J.M., Shin, H.W. and Wisler, D.C., "Mean Velocity and Turbulent

Flow Field Characteristics inside an Advanced Combustor Swirl Cup", AIAA

Paper 89-0215 (1989).

7. Cameron, CD., Brouwer, J. and Samuelsen, G.S., "A Comparison of Spray

Characterization in an Isothermal Chamber and in a Model Gas Turbine Can

Combustor", Second Int'l Conference on Liquid Atomization and Spray

System , pp. 145-152, Sendai, Japan (1988).

8. McDonell, V.G., Wood, G.P. and Samuelsen, G.S., "A Comparison of

Spatially Resolved Drop Size and Drop Velocity Measurements in an

Isothermal Chamber and a Swirl Stabilized Combustor", Twenty-First Symp.

(Int'l) on Combustion , pp. 685-694, The Combustion Institute, Munich,

West Germany (19 86).

9. De la Rosa, A.B., Bachalo, W.P. and Rudoff, R.C., "Spray

Characterization and Turbulence Properties in an Isothermal Spray with

Swirl," ASME Journal of Engineering For Gas Turbines and Power, Vol.

112, pp. 60-66 (1990).

10. Sankar, S.V., De la Rosa, A.B., Isakovic, A. and Bachalo, W.D., "A

Technique for Studying Mixing in Swirl Combustor," Sixth Int'l Sypm. on

Laser Anemometry Application to Fluid Mechanics . Lisbon, Portugal

( 1990)

.

11. Wang, H.Y., Sowa, W., McDonell, V.G. and Samuelsen, G.S., "Swirl Cup

Continuous and Discrete Phase Measurements in a 3x Scale Module", Report

UCI-ARTR-9 0-11-A, Combustion Laboratory, Univ. of Cal., Irvine (1990).

694
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 77

AIR BLAST ATOMIZATION: INSTABILITY OF AN ANNULAR LIQUID SHEET

SURROUNDED BY TWO AIR FLOWS OF DIFFERENT VELOCITIES

P. Camatte and M. Ledoux

URA CNRS 230/CORIA - University of Rouen

Mont Saint Aignan, France

ABSTRACT

Air blast atomization of a flat or annular sheet is studied under the

influence of internal and external air flows of two different velocities.

Non dimensional results are presented. Asymptotic cases of high velocity

atomizations, high thickness and high radius sheet are derived. First

experiments are also presented.

INTRODUCTION

Air blast atomization is of a wide use, especially in aircraft

engines where air is provided in excess by the compressor. In many

aircraft atomizers an annular liquid fuel sheet is disintegrated by to

cof lowing air flows, one internal the other external. These air flows are

generally swirling and at two different relative velocities with liquid

flow

.

Modelisation of such an atomizer involves two problems : 1) -instability of the annular sheet. 2) -breaking and rest ructuration into

drops. The present paper deals with the first problem, i.e. the

instability of an annular sheet surrounded by two coaxial air flows.

Few studies of this problem are found in the litterature [1] [2]

[3] [4] [5] all with the same velocity on both sides of the sheet. An

analysis is given here in the case of different velocities of internal and

external air flows

.

Linear analysis as given here is criticable in nature since break-up

processes are far from linearity. Nevertheless linear analysis has been

previously usued with success in designing pressure swirl atomizers [4]

.

It has also to be pointed out that non linear theory (weakly non linear or

based on wortex sheet analysis) need results of linear calculations to be

achieved. Furthermore interesting qualitative information can be drawn

from linear analysis.

695
ANALYSIS

The flow is shown on Fig.l in stable configuration. Both liquid and

gases are non-viscous (which can be accepted for low viscosity fluids as

water or kerosene). Two linear analyses of stability are given : l)-for a flat sheet (asymptotic case for a -> oo) 2) -for the real case (a finite) . Notations are defined on Fig.l.

General features

Both analyses are based on the same

physical grounds. Time varying velocity field

is decomposed into stable and instable part . In

each flow this last part derives from a potential (<p* , (p~ and (p^) t| (x,t,9) is the

deviation of the surface from stable shape.

Potentials are searched under the separate

form R T(9)e^"'^ R=R(y) in case 1 (plane) or

R=R(r) in case 2 (annular sheet) . From

following interface conditions iq will be found

of the form :

Fig.l The stable flow

= ti^ exp i ( kx + n8 - Wt ) (2.1)

k is the axial wavenumber : k = 2'n/X n is the azimuthal wavenumber

0) is a complex pulsation (J = (O^ + iw^

CO is related to axial propagation velocity c =kc

CO, t 0)^ when positive leads to instability a e

Ti^ is the initial amplitude at each side of the sheet ; one finds then two

values Ti'^ and -n" .

Continuity (2.2) and impulsion equation (2.3) under linearized form are

written :

Zkp"" = O ; Zkp" = 0 ; Z^iPj^ = 0 (2.2) —

du du W du 1 dp

+ U — + = (2.3)

dt dx r dO p 3x

in each fluid (In case 1, V E 0)

Condition at the interface give : —9<p df] dr\ V 8ti = — + U — + (2.4)

dr dt dx r 88

in each fluid at each interface.

Surface tension gives :

696
Pr ~ P = ± (J r^L r'g as (2.5)

at each interface, under non linearized form.

Results will be shown here under non dimensionalized forms. U^^ is the

average velocity of the two air flows Uj^= (\J* + U~)/2.

AU is the difference between the two air flows Au = U"" + U" (Au = 0

leads to U„ = = U" )

.

One introduces which is the maximum value of k leading to

instability in the case of symetrical plane sheet (Squire problem)

Pa U„

k =

c

So h = k^h, k = k/k^. ~ Pg ~ Au ~

p = — , Au = —, to =

(J

Pt U„ k^ U„ ^L M c M

By combining equations (2.2) to (2.5) a system of the following form

is found for and t]" :

FTi ^+ Gt]- = 0

G'< + F'ti" = 0

(2.6)

F,G,F',G' are depending on the system studied. Detailed calculations can

be obtained from the authors. (2.6) shows that : -The determinant FF'- GG' must be zero.

- and t]' cannot be independant . Their value can be arbitrary but

their ratio r\^/ iri" can be calculated from - G/F or - F'/G'.

FF' - GG' gives a dispersion equations of the form :

^4 ~3 ~2

0) + H3 (J + H2 (J + W + = O

where coefficients are function of k and n,

(2.7)

Case 1 : The flat sheet

a) U"^ , U~ , U^ are parallel flows. The dispersion equation (2.7) has the

form :

-W^coth 2kh-p(ku'^-t0) ^ +

crk- -tJ^coth 2kh-p(kU -(O) ^ +

sh-" 2kh

(2.8:

The ratio between amplitudes on two sides is :

'Ho - (J

° sh 2kh -W^coth 2kh-p (ku'' -w) ^ +

CJk^

697
sh(2kh) -CO^COth 2kh-p(kU -(J) ^ +

(2.9)

-W

These equations are interesting not only as an asymptotic result

(a -» oo) of case 2, but also the equations allow visible simplifications in

the asymptotic case kh -» oo.

b) Asymptotic case kh -> oo This case corresponds either to great velocities U"^ and U~ (small A,

great k) or to a very thick sheets (h great)

.

In both cases kh -> oo, sh 2kh oo and term

(0

vanishes

,

sh^ 2kh

(2.8) is then factorizing in two equations of degree two which can be

recognized as Kelvin-Helmholt z equations for simple interfaces. So each

interface is atomized independently. T|g / r\~ has then no definite values.

c) U"^ = U~ = U. This is the problem of Squire [6] . The equation is

factorized into two new equations of degree two, of the form :

A 0)^ + p (kU - 03)^

crk^ 0

where A = tank kh or coth kh . These two values of A correspond

respectively to values of / ti~ equal +1 (antisymetric perturbations) or

-1 (symetric case) . This corresponds to the problem analyzed by Hagerty

and Shea [7] . Corresponding solutions are shown on figures 2 and 3. It can be seen

that for increasing h solutions are asymptotic to a unique curve, which is

in fact the Kelvin Helmholtz solution. This is achieved for h of order 10.

0.10

0.08

0.06

0.04 -

0.02

/1U=0 ," = 1.29 10-

0.00

.015

0.0 0.4 0.8 1.2 0.0 0.4 0.8 1 .2

Fig. 2 First solution Growth rate versus wave number

Fig. 3 Second solution Growth rate versus wave number

d) Case when U""?; U

Solutions are presented on Fig. 4 and 5. Two different solutions

698
giving (j^> O are found. So the shape of the sheet have to be considered as

the sum of those two solutions. Some features have to be remarked.

-For thick sheets (h greater than 5) the ratio t|q / t)" (or inverse)

becomes very great. The perturbation from one side has a very low

influence on the other side (very low "penetration") . In this case the

values of / t|~ found here only give a qualitative indication since the

non-linearity is not represented. In fact solutions are asymptotic to

Kelvin-Helmholtz theory for each interface.

4ij=1 r = 1.29 10-3

First solution

0.1E+05

0. 1E + 04

0. 1E + 03

0. 1E + 02

0.02

0 .00

0. 1E + 01

I

/ //'

/ // / //

/ /

/

0.0 2.5

Fig. 4.a Growth rate versus wave numt>er

Fig. 4.b Amplitude ratio versus wave numt>er

-For very small h, amplitude rati,o is practically equal to 1 (or

negative for the second solution) (see h = 0.1 on fig. 3-4) . That is not

surprising since it is the case particularly of this sheets. Furthermore,

one solution is predominant

.

-In intermediate cases (e.g. h = 1-5) / t]~ can be significally

different from 1 or -1. CJ^ terms are still very different. Generally one

solution can be then be considered as predominant from the point of view

of atomization (see fig. 3-4) . In this case it is to be noticed that

diagrams are resolutly different from Kelvin-Helmholtz results (h « 1)

.

.0020

.0016

.0012

.0008 -

. 0004

. 0000 1!-^

0.0

0.0

0.1 0.2 0.3

-0 . 1 L

4U=1 P= 1.29 10-3

Second solution

Fig. 5. a Growth rate versus wave number h

Fig. 5.b Amplitude ratio versus wave number

699
Sj/k

e.2e

a=1

plane

X, sheet

Rg. 6.a Growth velocity for the maximum

(/I growth rate versus h

a.

7

a.

3

-S -4 -3 -2 -1 Log^Qh

Fig. 6.b Wave velocity for the maximum

growth rate versus h

&U =0 r =1.2910-3

Fig. 6.C Wave number for the maximum

growth rate versus h

Fig. 6.d twiaximum growth rate

versus h

1.00

0.8 xa=0.5

a = 1

_5 -4 -3 -2 -1 Ln h

Rg. 7 a Growth velocity for the maximum

growth rate versus h

2.50

^U=1 0=1.2910

-5 -4 -3 -2 -1 Ln h

Fig. 7.b Wave velocity for the maximum

growth.rate versus h

2.00

1.20

0.40

-5 -4 -3 -2 -1 Ln

Fig. 7.0 Wave numt^er for the maximum

growth rate versus h

-5 -4 -3 -2 -1

Fig. 7.d Maximum growth rate

versus h

700
Annular sheet

A new parameter ^is now involved, the radius a. It will appear

non-dimensionally as a = . No place can be found here to show the

analytic forms of J^he results^but some calculated values are shown for two

specific cases : ^ = 0 and Au = 1

.

It is to be remarked that results presented here do not involve the

hypothesis of very thin sheet (h << a) which was accepted in previous

papers ([1] [2] [3] [5]). It has been shown by the authors [8] that this

hypothesis could not stand when h/a > 10"^ [8] . ^

Two solutions are also found. As an example the solution of greated b)^ is

shown on figures 6 and 7. Some features can be observed.

1) -Case^AU = 0 (fig. 6) _

- (0^ has a jump around h equal 10"^ O)^ is proportional to c and the

corresponding^ values of c are that of the gas (small h) or that of the

liquid ^high h) . ^

- (j3^ is very small at lower h . There is a very small velocity

differg^nce between the wave and the gas

.

-(jj^ is small at higher h. The inertia of the sheet becomes important.

-For a>10 solution become asymptotic to the solution relative to the

flat sheet. This give a criteria to adopt then a simpler analysis

(case 1)

.

2) -Case Au = 1 (fig. 7) The distinction between aerodynamical (low h) and

inertial domains (high h) is found again on O) curves.

When a increases (a>10) the analysis relative to fat sheet is valid

again

.

Although no results are presented here, our model can also deal with

swirling flows

.

EXPERIMENTS

In order to validate some of these results, a generator of

cylindrical shet has been built. No place is found here to describe it but

some characteristics are given : -thickness h = 300pjn -radius a = 2,55mm -velocity of liquid = 2-4m/s

-velocity of internal air flow U~ = 0-40m/s.

Photography and acoustic frequency measurements are used. Experiments

are in advance now to study effects of swirl and external flows.

CONCLUSION

A new model of the instability of flat and annular sheet surrounded

by air flows of different velocities has been presented here. It has been

shown that two different perturbations where running along the sheet, both

with different amplitudes on each side. Asymptotic cases show that, for

high thicknesses or high velocities, the atomization (Kelvin-Helmholtz)

theory was valid. A criteria (a 10) is given for identifying an annular

sheet of great radius to a flat sheet. A first experimental validation is

presented.

This work was executed under CEE BRITE-EURAM contracts as associated

partner with Society SNECMA.

701
Nomenclature

a radius of annular sheet

h thickness of the sheet

k wavenumber (=2'Tr/A)

Ti perturbation of interface

CO complex pulsation

a surface tension

p density (L liquid, +, - gas)

subscript ~ indicate non dimensionalized numbers

BIBLIOGRAPHY

[1] G.D. Crapper, N. Dombrowski and G.A.D. Pyott, "Kelvin-Helmholt z wave

growth on cylindrical sheets". J. Fluid Mech . , vol. 68 (1975) . [2] Martinon, "Instabilit6s d'une nappe liquide cylindrique en presence

de deux ecoulements gazeux" . Journal de Mecanique Theorique et

Appliquee, vol.2 (1983).

[3] Guillie , "Etude experimentale et theorique de 1' injection

aerodynamique . Thdse de 3eme cycle a I'Ecole Nationale des Mines de

Paris (1987) . [4] CH. Dumouchel , "Problemes li6s a la conception d'un pulverisateur

mecanique hydrodynamique de chambre et instabilite de nappe". These

de I'Universite de Rouen (1989)

.

[5] Meyer J., Weihs D., J. Fluids Mech, vol.179, p. 531-545 (1987).

[6] Squire, Brit. Journ. of Appl. Phys . , vol.4 (1953).

[7] Hagerty and Shea, "A study of the stability of plane fluid sheets".

Journal of Applied Mechanics, vol.22 (1955)

.

[8] PH. Camatte, in ILASS (1989) proceedings, Pisa.

[9] Kendall, "Experiments on annular liquid jet instability and on the

formation of liquid shells" . Phys . of Fluids, vol.29 (1986).

702
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 78

LIQUID JET INSTABILITY EV COAXIAL AIR FLOW

H. Eroglu and N. Chigier

Mechanical Engineering Department

Carnegie Mellon University

Pittsburgh, PA, U.S.A.

ABSTRACT

Wave characteristics of liquid jets issuing from airblast coaxial atomizers have been

investigated. The wavelengths on the liquid jet surface were measured from high magnification

micro-photographs over a liquid jet velocity range of 1.1 - 18.2 m/s and for the air velocities 45.8,

56.12, 64.8 and 91.64 m/s. The wave frequencies were measured using a laser beam attenuation

technique. Both dilational and sinuous waves were observed. The wavelengths generally

increased in the downstream axial direction. Average wavelengths decreased with liquid and air

velocities, however, the average jet surface oscillation frequencies were found to increase with

these velocities. Wave propagation velocities were found to have much lower magnitudes than the

liquid jet velocity for all air velocities, and were insensitive to the variation of these velocities.

INTRODUCTION

The disintegration of liquid sheets and jets has been studied by many investigators.

Comprehensive reviews of disintegration mechanisms are reported by McCarthy and Molloy [1],

Reitz and Bracco [2], and Lefebvre [3]. The disintegration of low viscosity liquid jets issuing at

sufficiently high velocities from a nozzle was analyzed by Levich [4]. According to this analysis,

both long and short length (as compared to the jet radius) waves appear on the jet surface under the

effect of infinitely small disturbances. Because of the dynamic influence of air, both the long and

the short wavelengths become unstable and lead to the generation of a wide range of drop sizes.

Recent theoretical research has concentrated on the development of computational models [5-

71 for analysis of the disintegration of liquid sheets and jets. However, the number of

experimental investigations providing data on the wave characteristics of liquid sheets and jets is

very limited.

Among the recent experimental studies, Aral and Hashimoto [8] investigated the disintegration

of a thin liquid sheet in a co-current air stream. The vibration frequency of the liquid sheet was

measured using an electro-optical displacement device. The breakup frequency was obtained by

means of a stroboscope. The liquid sheet vibration frequency was found to be equal to the breakup

frequency. The liquid sheet vibration frequency was shown to be a function of Weber and

Reynolds numbers. These results also showed that the disintegration of the thin liquid sheet

depends on the wave motion of the liquid sheet. Therefore, the mean droplet diameter resulting

from the disintegration was correlated to the frequency of the vibrating liquid sheet.

703
Mansour and Chigier [9] studied the aerodynamic instability of liquid sheets issuing from a two-dimensional air-assisted atomizer. Still microphotography was used to determine the global

structure of the sprays, and a laser light attenuation technique was employed to measure the

oscillation frequencies. The measured vibrational frequencies were correlated with the resulting

spray angle. It was shown that the liquid sheet oscillations are dynamically similar to those in hard

spring systems. At low liquid flow rates, the sinuous breakup was found to be the dominant

mode. At high liquid flow rates, the oscillations were of the dilational type. At intermediate liquid

flow rates, both the sinuous and the dilational modes were superimposed on the liquid sheet.

This paper presents the results of the wavelength and frequency measurements for liquid jets

issuing from airblast coaxial atomizers.

EXPERIMENTAL CONFIGURATION

The atomizer consists of a central water jet surrounded by a coaxial air jet. Both water and air

supply pressures and mass flow rates can be varied. The house air supply provides a maximum

inlet pressure of 25 kPa for the atomizing air. Water is supplied from a tank with 10 liter capacity

and is pressurized with house air. This water supply provides a pulsation and vibration free water

flow and a maximum water pressure of 600 kPa if no air is used to assist the atomization. When

the maximum air flow rate is used for the atomization, the maximum possible water supply

pressure is reduced to 500 kPa.

The test atomizer and the exhaust system were connected to the same three-directional (z-r-phi)

positioner so that the axial distance between the nozzle exit and the exhaust inlet always remained

the same (120 mm). The accuracy of the positioner is 0.01 mm in the axial (z) and radial (r)

directions and 0. 1 degree in the phi direction.

The test nozzle has provisions for the accurate positioning of the central tube with respect to

the outer tube. A screen is used in order to reduce turbulence and also serves as a flow

straightener. The converging section of the outer tube provides a plug flow air velocity profile at

the nozzle exit. Figure 1 shows the nozzle exit geometry; the dimensions are in millimeters.

Photographs of the sprays were taken using a Canon Fl camera with a f = 200 mm macro

lens. Different extension tubes and bellows were attached to the camera. An Electro-Optics

EG&G 549 microflash was used to provide a 0.5 |is flash duration and 5x10^ beam candlepower

light intensity. More than 1500 photographs were taken at various air and water flow conditions.

The description of the photographs was previously reported by Farago and Chigier [10].

The schematic of the frequency measurement system is shown in Fig. 2. This system utilizes

the attenuation of a collimated laser beam passing through the liquid jet in the transverse direction

(y-direction). The intensity of the laser beam is measured using a photodiode which generates a current proportional to the incident radiation. This current is converted to a voltage in the range of

0-1.5 volts. A Tektronix T912 storage oscilloscope is used to visualize the signal. The wave

form is then transferred to an IBM PC/XT for data analysis and storage. An FFT routine is used to

transform the wave form from the time domain to the frequency domain. The power spectra of

oscillations are then displayed in the graphics form.

« 10 36

Fig. 1 Enlarged sketch of the nozzle exit

704
Initially, the axial location of the laser beam was brought close to the nozzle exit. Then, the

axial position of the beam with respect to the nozzle was varied until a clear wave form was

obtained on the oscilloscope screen. The actual waves on the surface of the jet were also observed

using a stroboscope. With the present arrangement, it was possible to provide frequency

information about the initial disturbances before the waves are transformed from the dominantly

1) He-Ne Laser

2) Collimating Lens

3) Nozzle

4) Optical Filter 5) Focusing Lens

6) Photodiode

7) Current to Voltage Converter

8) A-D Converter

9) IBM XT

Fig. 2 Frequency measurement system

dilational to the sinuous mode. This information is particularly important for the computational

analysis of liquid jet instability.

RESULTS AND DISCUSSION

Wave characteristics of liquid jets issuing from airblast coaxial atomizers were investigated in

terms of wavelengths and frequencies. Jet surface wavelengths were measured from the

photographs over the liquid jet velocity range 1.1 - 18.2 m/s and for the gas jet velocities 45.8,

56.12, 64.8 and 91.64 m/s. Liquid jet surface wave frequencies were measured using a laser

beam attenuation technique.

Wavelengths

A careful investigation of the photographs shows that, in general, two types of waves are encountered as the liquid jet velocity is varied while keeping the air velocity constant. The

dilational wave develops on the jet surface and may or may not be symmetrical with respect to the

jet axis. The sinuous wave has an assymetric, snakelike shape whose centerline deviates from the

nozzle axis. At low liquid jet velocities, the waves are predominantly sinuous, however, at high

Hquid velocities, the dilational waves are formed on the liquid jet surface.

Figure 3 shows the wave development of the liquid jet for the liquid and air velocities of 4.5

m/s and 56.12 m/s respectively. At such an intermediate liquid jet velocity, the initial disturbances

are of the dilational type. Sinuous wave formation starts further downstream. Dilational waves are superimposed on the sinuous waves. Due to the aerodynamic action, dilational waves lead to the

formation of ligaments mostly at the crests of the sinuous waves. These ligaments further

705
disintegrate by the shear of the surrounding air as shown in Fig. 3. This figure also shows that the

entire jet breaks up in a similar manner to that of liquid sheets where the half waves are torn off the

liquid sheet when the wave amplitude reaches a critical value.

Figure 4 shows the wave formation on the liquid jet at liquid and gas velocities of 9.56 m/s

and 56.12 m/s respectively. The upper figure was obtained by scanning the photograph of the

liquid jet at 100% and 300 dots per inch using an HP scanner and Applescan software. The outline

of the scanned image shown in the lower figure was then generated. The locations of wave crests

were marked on the same figure, and the distances between the marks were measured. These

distances were accepted as the wavelengths. The bottom figure shows the variation of the

normalized wavelength (X/d) in the axial direction (x/d) for both the upper and the lower surfaces.

The wavelengths increase almost linearly in the downstream direction. The magnitudes of the

wavelengths are very close to each other for the upper and the lower surfaces at the same axial

location. Although the waves of Fig. 4 are predominantly of the dilational type, the initiation of

sinuous wave formation can also be observed.

Since the wavelengths vary in the axial direction, average wavelengths were calculated for the

dilational surface waves before the transition to sinuous mode takes place. These were normalized

with respect to the jet diameter at the nozzle exit and plotted as a function of liquid jet velocity for

four different gas velocities. Figure 5 shows this plot. In general, the wavelengths decrease with

increase in liquid jet velocity and gas velocity. This decrease is more pronounced for the liquid jet

velocities at which the transition from laminar to turbulent flow occurs. After transition to

turbulence has been completed, the wavelengths seem to be insensitive to the increase in the liquid

jet velocity. This may be due to the predominant effect of the small-scale turbulent structures

inside the liquid jet controlling the initial phase of the wave development.

Figures 4 and 5 show that the dilational waves have lengths shorter than the perimeter of the

liquid jet (k < nd). It appears that the aerodynamic influence is the major destabilizing force for the

wave development rather than the surface tension force within the investigated range of flow

706
2.5

;5 2.01

O)

S 1.5- >

en

1.0-

0) ^ 0.5 i

0.0

0 = 45.8 m/s

£, Va = 56.12 m/s

0 V, , = 64.8 m/s

* Va , = 91.64 m/s

500 -

>>o

c

^ 300

cr

<u

—

I

10

200

100

1.1 m/s

Vgjr = 45.8 m/s

000 2 00 10,00

15 20

Axial Distance (x/d) Liquid Jet Velocity (m/s)

Fig. 5 Variation of average wavelength with liquid Fig. 6 Variation of average frequency

and gas jet velocities along the jet axis ^

(Vair = 45.8 m/s, Vj = 1.1 m/s)

conditions. It also seems that the turbulence in the liquid jet is responsible for the formation of

initial disturbances as shown in Fig. 4 at higher liquid jet velocities.

Frequencies

The primary interest in this paper is to provide information on the initial disturbances close to

the nozzle exit. The detection of the disturbances has less uncertainty, however, at downstream

axial locations (x/d) because of the more pronounced waves. The effect of axial measurement

location on the measured wave frequencies was investigated in an effort to determine the optimum

measurement location before the waves are transformed into the sinuous mode. Figure 6 shows

the result of this investigation for the liquid and gas jet velocities of 1.1 m/s and 45.8 m/s.

Frequency is shown to be rather insensitive to the axial measurement location as long as the type of

the wave remains the same.

Figures 7 and 8 show the power spectral densities of the liquid jet surface oscillations for an

air velocity of 56.12 m/s and the water velocities of 1.5 m/s and 12.76 m/s, respectively. Figure 7

shows a dominant oscillation frequency at the lower liquid jet velocity. However, at higher liquid

velocities (Fig. 8) several modes of oscillation appear, and the band width of oscillations increases

considerably. The liquid jet turbulence is probably the major cause for this widening of the

frequency band width. The measured frequencies also increase with the liquid jet velocity.

Because of the difficulty in determining the dominant modes of oscillation at high liquid jet

velocities, single representative frequencies were obtained at each pair of flow conditions by

wo

O 3

Q. O

Vgir = 56.12 m/s

V| = 1 .5 m/s

Frequency (Hz)

Fig. 7 Power spectral density at Vair = 56. 12 m/s, Vj = 1 .5 m/s

707
W

o

wer C 2.0 - to o

CL 2 -s -

Vair = 56.12 m/s

Vj =12.76 m/s

Frequency (Hz)

Fig. 8 Power spectral density at Vair = 56.12 m/s, Vj = 12.76 m/s

weighting the frequency values with corresponding power densities and averaging over the whole

power range. The repeatability of the measurements was found to be poor (15% uncertainty) at

high liquid jet velocities for which the frequency band widths widen considerably. Therefore, each

experiment was repeated four to five times, and the calculated average frequencies were further

averaged in order to increase the confidence in the data.

The variation of average frequencies as a function of nozzle exit liquid and gas jet velocities is

shown in Fig. 9. The average frequencies increase both with the liquid and the gas jet velocities.

This behavior is consistent with the results of Aral and Hashimoto [8] and Mansour and Chigier

[9] on the liquid sheet instability with co-current air flow. The frequency jumps separating the

regions for different modes of oscillation were also reported by Mansour and Chigier [9]. These

jumps, however, were not observed for the coaxial liquid jet of this study. This was probably due

to an insufficient number of variations of experimental conditions to identify this phenomenon.

Wave Propagation Speed

The wave propagation speeds were obtained by taking the product of average frequencies and

wavelengths over the range of velocities 1.1 - 18.2 m/s for water, and 45.8 - 91.64 m/s for air.

They were found to be in the order of 1 - 2 m/s and do not seem to be sensitive to the water and the

air velocity variations. This behavior is consistent with the wavelength decrease in Fig. 5 and the

corresponding frequency increase in Fig. 9 with increasing liquid and gas jet velocities.

NX

5000-

4000-

3000-

o

c

<1>

S" 2000

1000-

vair = 45.8 m/s

vair = 56.12 m/s

vair = 64.80 m/s

vair = 91.64 m/s

Vair = 129.50 m/s

10 20

Liquid Jet Velocity (m/s)

Fig. 9 Variation of average frequency with liquid and gas jet velocities

708
CONCLUSIONS

The wave characteristics of liquid jets issuing from airblast coaxial atomizers were

investigated. Two types of waves were encountered. The dominant wave type was sinuous at low

liquid jet velocities, and dilational at high liquid jet velocities. The wavelengths generally increased

in the downstream axial direction. Average wavelengths decreased with liquid and gas jet

velocities. The dilational waves of this study were found to have shorter lengths than the perimeter

of the liquid jet. The major destabilizing force is the aerodynamic interaction at the liquid/gas

interphase. The turbulence in the liquid jet is important in initiating the surface disturbances at high

liquid jet velocities. The average oscillation frequencies at the liquid jet surface appear to be

insensitive to the axial measurement location. The frequency band of the jet surface oscillations

increased with the liquid jet velocity. Average frequencies were found to increase with both the

liquid and the gas jet velocities. Wave propagation speed seems to be rather insensitive to the

liquid and the gas velocity variations.

NOMENCLATURE

d jet diameter

phi circumferential direction of the atomizer

r radial direction of the atomizer

Vair air jet velocity

Vj liquid jet velocity

X axial (Erection of the jet

y transverse direction of the jet z vertical direction of the jet; axial direction of the nozzle

X wavelength

ACKNOWLEDGEMENT

The authors wish to acknowledge financial support from NASA Marshall Space Center, grant

number NAGS- 126.

REFERENCES

1 . McCarthy, M.J. and Molloy, N.A., "Review of Stability of Liquid Jets and the Influence of

Nozzle Design", Chemical Engineering Journal, Vol. 7, pp. 1 - 20, 1974.

2. Reitz, R.D. and Bracco, F.V., "Mechanism of Atomization of a Liquid Jet", Phvsics of

Fluids. Vol. 25, No. 2, pp. 1730 - 1741, 1982.

3. Lefebvre, A.H., Atomization and Sprays. Hemisphere Publishing Company, New York,

1989.

4. Levich, V.G., Physicochemical Hydrodynamics. Prentice-Hall, New Jersey, 1962.

5. Lin, S.P. and Lian, Z.W., "Mechanisms of the Breakup of Liquid Jets", AIAA Journal, Vol.

28, No. 1, pp. 120 - 126, 1990.

6. Przekwas, A.J., Chuech, S.G. and Singhal, A.K., "Numerical Modeling for Primary

Atomization of Liquid Jets", AIAA Paper No. 89 - 0163, 1989.

7. Rangel, R.H. and Sirignano, W.A., "Nonlinear Growth of Kelvin-Helmholtz Instability:

Effect of Surface Tension and Density Ratio", Phvsics of Fluids. Vol. 31, No. 7, pp. 1845 - 1855, 1988.

709
8. Arai, T. and Hashimoto, H., "Disintegration of a Thin Liquid Sheet in a Cocurrent Gas

Stream", Proceedings of the 3rd International Conference on Liquid Atomization and Spray

Systems, pp. VIB/1/1-7, 1985.

9. Mansour, A. and Chigier, N., "Dynamic Similarity Between Liquid Sheets and Hard Spring

Systems", Physics of Fluids A, 1991.

10. Farago, Z. and Chigier, N., "Parametric Experiments on Coaxial Liquid Jet Atomization",

ASME Paper No. 90-GT-81, 1990.

710
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 79

EFFICIENT SPRAY MODELING ON VARIOUS APPLICATIONS

USING THE GROUP APPROACH

Q. Zhou and S.C. Yao

Department of Mechanical Engineering

Carnegie Mellon University

Pittsburgh, PA, U.S.A.

ABSTRACT

A "Group" approach for the modeling of sprays has been well established and applied to various

problems, e.g. particle laden jets, impacting sprays and spray heat transfer. In the group model, a

droplet group is considered to have a dimension which grows during its motion due to the turbulent

dispersion of droplets. The droplet dispersion within a group is described by a probability density

distribution function. All the applications have shown that the group model is very efficient in the

computational time, can give very realistic descriptions, and have very wide applications.

INTRODUCTION

Turbulent sprays have been widely applied to various industrial processes, such as spray combustion,

spray cooling, spray painting, spray drying etc. Computer modeling of turbulent sprays is becoming an

increasingly viable part of the design process with the objective to reduce the time and cost of process

development. In the efforts for developing spray models, the major focus is on the turbulent interaction

between gas phase and dispersed phase. Several approaches have been employed to model the turbulent

droplet dispersion in sprays [1]. However, because of their excessive computational requirements, the

spray models are still not feasible for general industrial applications.

Among the existing models, a rather advanced one is the Stochastic Separated Flow (SSF) model

which has been widely accepted for the modeling of turbulent sprays [1, 2]. The SSF model uses droplet

clusters to represent a spray and simulate the turbulent dispersion directly by considering the interaction

between pseudo-random turbulent eddies and droplets in a Lagrangian computation of cluster motion.

In the SSF approach, however, all the droplets within the cluster are assumed to be concentrated at one point during the flight. A droplet cluster is essentially a computational droplet. Therefore, in order

to obtain a statistically meaningful distribution, a significant number of computational droplets, and

consequently, very long computation times are needed.

A new group approach has been well established by the authors of this paper [3, 4, 5]. The group

model uses droplet groups to represent a spray and takes into account the turbulent droplet dispersion

within each group. It considers that each group has a dimension which grows during the flight. This

growth is due to the turbulent dispersion of droplets as the group travels in the Lagrangian coordinates.

The droplet dispersion within the group is described by a probability density function which is a Gaussian

711
distribution. Due to the growth of the group, only a small number of groups must be traced to provide

statistically reasonable results. Furthermore, the group approach closely simulates a real flow, because

in a real flow the droplets within a group are dispersed by turbulence instead of staying at one point.

The intent of this paper is not to elaborate the details of the group model, but to provide various

examples of its applications, the effectiveness of the model, and the efficiency of the computations using

this approach. The implementation of this model to different problems, such as particle laden jets,

impacting sprays, and spray heat transfer, has shown its generalized application to sprays. The concept

of this group approach can be incorporated into any computational program which presently employs

the Lagrangian approach to trace the dispersed phase, and widely applied to various sprays in industry.

This new modeling concept is expected to have more advantages when the geometrical scale is large,

where its higher computational efficiency becomes indispensable.

GROUP APPROACH

In the present study, a standard Eulerian approach is applied to the gas phase transport. The SSF

analysis (Lagrangian formulation) is used to trace the motion of the center of each group which is assumed to behave like a single droplet interacting with a succession of random distributed turbulent

eddies [1]. The major assumption in the modeling of droplet dispersion within a group is that the

turbulent flow is assumed to be locally isotropic and homogeneous within each group. It is also assumed

that all the droplets in a group move together in an absolute coordinate and the droplet dispersion is a

relative motion to the center of the group. Therefore, the droplet group grows from its center, and the

dispersion of droplets is observed from the center of the group. The relative droplet position within a group in respect to the group center at any time is random because of the turbulent gas velocity field. This kind of droplet dispersion can be closely represented by a normal Gaussian distribution [6]. The

typical probability density function for this distribution can be given for a 2-D field as:

where N is the total droplet number in this group, and X"^ and Y"^ are the overall mean square distances

in X and y directions. Further details of this group approach can be found in Ref. [3, 4, and 5].

APPLICATION OF GROUP APPROACH

Particle Laden Jets

In order to validate the new approach, the group modeling was applied to the analysis of a particleladen jet containing fly ash particles. The preliminary results of this study were compared with the

results of the SSF method and the experimental data. The experimental results were reported by Yuu

et al. in reference [7]. Their experimental set-up contained a jet nozzle which could provide uniform

exit properties. The nozzle diameter is 8mm. The particles are nearly monodisperse having a mass mean diameter of 20 /im. The flow is dilute with particle mass loadings of 0.1-0.4%. These loadings

are sufficiently small, so that the particles have a negligible effect on the mean and turbulent gas phase

properties.

The distributions of particle concentrations along the center line of the jet are shown in Figure 1 for

the present method, the SSF approach, and the experimental results. The SSF results are presented

for two different particle cluster numbers (20 and 500). For the smaller particle cluster number, the

particle concentration distribution is not smooth due to statistically insufficient data. In order to obtain a reasonable agreement with the experimental results, 500 particle clusters were required as shown in the

figure. Unlike the results of the SSF method, the results of the present approach are in sound agreement

712
c/co

1.4

1.2

1

.8

.6

.4

1

_

1

1

1 1

...O.. Experimental Data Group Approach N=15

SSF Data N=500

\^ O SSF Data N=20

S> + +

1 1 ] 10 15 x/d

20 25 30

Figure 1: Axial Particle Concentration Distributions at the center line

with the experimental results with only 15 groups. At x/d more than 15, slightly larger deviation is observed.

Figure 2 compares the radial distributions of particle concentration for the present approach and the

SSF method. Both methods have very similar results along radial direction also. Since the corresponding

experimental data of radial distribution and radial initial conditions are not clearly reported in reference

[7], the comparison was made only with the results of the SSF method.

0.0040

0.0035

0.0030

. 0.0025 C

0.0020

0.0015

0.0010

0.0005

0

.... Group Approach N=15

SSF Approach N=500

0 .1 .2 .3 .4 .5 .6 r(cm)

Figure 2: Radial Particle Concentration Distributions at x/d=20

A 32 X 29 grid was used in the calculations. The calculations were performed on a Sun 3861/250

computer(5 MIPS, 25 MHz). For the dispersed phase calculation, the computational time using SSF

approach with 500 groups was 509 seconds, while the present study only took 53 seconds for a similar

agreement with the experimental data. The computation time using the group model is about a factor

of 10 times less than the corresponding calculation using the SSF approach. This major reduction in

the computation time in the present approach is because of the significant decrease in the number of

droplet groups required for realistic predictions.

713
Impacting Sprays

Impacting sprays were selected as an example to illustrate the capability of this group model. This

selection was chosen because impacting sprays contain droplets of various sizes under severe flow conditions, and there are not many detailed investigations reported in literature despite its wide applications.

Impingement of axisymmetric water sprays from a 9.5 mm diameter nozzle normally on a flat plate was

studied using the group approach. The distance between the plate and the exit of the injector is 30.5

cm The droplets are assumed to be monodisperse. The spray flow is considered to be dilute with an

initial liquid volume flow rate fraction of 10-5

16 t=\

r(cm)

5 10 15

x(cni)

Figure 3: Gas Flow Streamlines

The gas phase mean flow field is shown in Figure 3. The mean flow streamlines demonstrate three

flow regimes which were defined in reference [8]. The free jet regime is upstream of jet impingement.

The strong interaction of the jet with the impingement surface produces a change in flow direction in

the impingement regime. The wall jet regime consists of radial flow along the surface beyond the point

at which the strong impingement to the wall occurs.

Figure 4 shows the trajectories of droplet group centers for the droplet diameter of 20 fim. It indicates

16

14

12

10

r(cm) 8

6

4

2

0

N=21

-'''/^^^^^'/.l^y^j^^'Jj

10 15

x(cin)

20 25 30

Figure 4: The Trajectories of Group Centers, 0=20 fim

714
that the droplets follow the mean gas flow very closely and the spread of the trajectories is similar to that

of the gas flow streamlines because of the small momenta of droplets. The irregularity of trajectories

and the significant dispersion of droplets reveal that the interaction between the turbulent eddies and

the droplets is very effective. The computations for the droplet diameters of 150 and 50 fim have also

been performed. It was found that larger diameter droplets, such as those with a 150^m diameter, tend

to continue in their initial direction with less influence by the gas flow field. As the droplet diameter

decreases, the momenta of droplets decreases. Therefore, the droplets follow the mean gas flow more

closely, the trajectories are more aff'ected by the flow turbulence, and the droplets are spread more

widely.

0.0025

0.0020

0.0015 C

0.0010

0.0005

"0 2 4 6 8 10 12 14 16

r(cm)

Figure 5: Radial Distributions of Droplet Deposition Flux

The radial distributions of the droplet deposition flux on the wall are shown in Figure 5 for the

droplet diameters 20, 50, 100, and 150//m respectively. This droplet deposition flux is defined as the

volume of total droplet deposition per unit wall area per unit time. A deposition is assumed to occur when a droplet approaches the wall surface within a distance of 0.01 cm. The figure demonstrates that

the smaller diameter droplets spread wider and impinge over all the wall surface which is consistent with

the implication of the trajectories. With increasing droplet diameters, the deposition flux distribution

increases at the central locations, and the radial extent of impingements decreases.

A comparison of the group approach with the SSF approach was also made in the present study.

Figure 6 shows that the present group approach with 21 groups yields results similar to the SSF approach

with 980 clusters; however, the SSF distribution is not as smooth. Therefore, much less group number

is needed for the group model to get more realistic results as compared to the SSF method.

The radial distributions of locally averaged impact angles for the four different droplet diameters are shown in Figure 7. The impact angle, a, is defined as the angle between the drop trajectory and the

direction normal to the wall. The impact angle in general increases with distance from the symmetry

axis. In the area near the symmetry axis, the impact angle increases as the droplet diameter decreases

because the smaller droplets follow the gcis flow more closely. The 150 fim diameter droplets have the

smallest impact angles due to their high momentum. The impact angles of 20 and 50 //m droplets

are smaller than expected in the area between 2 cm and 10 cm wall radii. It is noted that much

of these small diameter droplets do not impact the surface of the plate and go out following the gas

flow. For the droplets impacting the surface, they may impact at small angles due to the effect of gas

flow turbulence. The average impact angles obtained from these few impacts, therefore, gives lower

impacting angles within the 2 cm to 10 cm wall radius range as shown in Figure 7. The information

of radial distributions of overall droplet impingement velocities for different droplet diameters was also

provided in the computations. It also showed that the larger the droplets, the less influenced by the gas

...O.. dp - 20 fim

... +.. dp = 50 fim

...... dp = 100 fim

... X.. dp = 150 ftm

N=21

715
0.0025

0.0020

0.0015

0.0010

0.0005

Group Approach, dp = 150 fun, N=21

SSF Approach, dp = 150 /im, N=980

10 12 14 16

r(cm)

Figure 6: Radial Distributions of Droplet Deposition Flux

90

a(degree)

801-

70

60

50

40

30

20

10

0

„.o—o

J*

7-7-7T7

...O.. dp = 20 /xm

...+.. dp = 50 /im

...... dp = 100 /jm

... X.. dp = 150 //m

6 8 10

r(cm)

12 14 16

Figure 7: Impact Angle Distributions on the Wall

velocity distribution.

A 36 X 35 grid was used in the calculations. For the liquid phase calculation of 150;/m diameter

droplets, the SSF approach took about 900 seconds with 980 groups, while the group approach only took

about 46 seconds for similar results. The computation time using the group approach is about a factor

of 20 times less than the corresponding calculation using the SSF method. As expected, the reduction of

computation time is more significant when the geometrical scale is large. Using the group approach, the

information on spray trajectories, droplet deposition flux, impact velocity, and impact angle etc. can be

obtained easily. This understanding allows for advances on various impacting spray applications.

Spray Heat Transfer

The heat transfer to particles in a spray was also investigated using the group approach. It is found that

the group approach allows for very effective modeling of this heat transfer process in a very straightforward manner. The results are presented in Figure 8. The geometry and the initial kinematic conditions were kept similar to that in reference [7]. However, the cold particles with initial temperature 300K were

716
injected into hot air stream with initial temperature 800K, and were heated up by the surrounding gas

due to mixing. The temperatures of gas and particles are solved by energy conservation equations. The

calculations were performed for fly ash and alumina particles of the same size (20 fim). The density, p, and specific heat, Cp, for fly ash are 2.0 gm/cm^ and 1.214 j/(kgK), respectively. These properties are

3.9 gm/cm^ and 1.5 j/(kgK) for alumina particles.

Figure 8 shows the axial variation of gas and particle temperatures. The gas temperature decreases

rapidly up to about 8 cm distance from the nozzle exit. Its decresise is more gradual further downstream.

The average particle temperatures increase steeply for both particle materials immediately downstream of

the nozzle. These temperatures show a peak for both materials, and gradually decreaise. The calculated

results of the group model which used 15 groups for both materials are in very good agreement with the

results using the SSF method with 500 particle clusters.

A comparison of the average axial temperature distribution for the fly ash and alumina particles has

shown that the product of density and specific heat have a strong influence on the magnitude of particle

temperature. Alumina particles heat-up and cool down more slowly than fly ash due to their higher

thermal inertia. This results in a lower peak and higher downstream temperatures for alumina particles.

CONCLUSIONS

A group approach has been well established. All the applications have shown that:

• The group approach is "efficient" in the computational time because of very small number of

droplet group required.

• The group approach is "realistic", since it considers the turbulent droplet dispersion within each

group.

• The group approach can be "generalized" , and has very wide applications.

717
NOMENCLATURE

C droplet concentration

probability density distribution

number of droplets in a group

radial coordinate

time

temprature

axial coordinate

mean square distances

vertical distance

f(x,y,t)

N

r

tT

X

y

Subscripts

d droplet property

References

[1] G. M. Faeth. Mixing, Transport and Combustion in Sprays, Prog. Energy Combust. Sci. Vol.13,

293, 1987.

[2] A. D. Gosman. Aspects of Computer Simulation of Liquid-Fueled Combustors, AIAA Paper No.

81-0323, 1981.

[3] Q. Zhou, and S. C. Yao. An Efficient Modeling of Turbulent Sprays, Eastern States Section of the

Combustion Institute, Clean Water Beach, Florida, 1988.

[4] Q. Zhou and S. C. Yao. A Model of Particle Dispersion for Turbulent Sprays, presented at Symposium on Heat and Mass Transfer in Fire and Combustion Systems, ASME Winter Annual Meeting,

Dallas, Texas, Nov. 1990.

[5] Q. Zhou and S. C. Yao. Group Modeling of Impacting Spray Dynamics, accepted by International

Journal of Heat and Mass Transfer.

[6] N. A. Fuchs The Mechanics of Aerosols, Macmillan Company, New York, 1964.

[7] S. Yuu, N. Yasukouchi, Y. Hirosawa and T. Jotaki. Particle Turbulent Diffusion in a Dust Laden

Round Jet, AIChE Journal 24, 509, 1978.

[8] CD. Donaldson and R.S. Snedeker. A study of free jet impingement. Part 1. Mean properties of

free and impinging jets, J. Fluid Mech., vol. 45, part 2 pp281-319, 1971.

718
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 80

AXISYMMETRIC CALCULATIONS OF THREE-DROPLET INTERACTIONS

C.H. Chiang and W. A. Sirignano

Department of Mechanical and Aerospace Engineering

University of California

Irvine, CA, U.S.A.

ABSTRACT

The present study extends the previous droplet models [1, 2] to investigate numerically the system of

three droplets which are moving in tandem with respect to the free flow. The purposes of this study are to study the wake effect of the lead droplet on the downstream droplets and to examine the effects of initial

spacing on the total system. The effects of variable thermophysical properties, transient heating and internal

circulation of liquid, deceleration of the flow due to the drag of the droplet, boundary-layer blowing, and

moving interface due to surface regression as well as relative droplet motion are included. The results are compared with those of an isolated droplet [1] as well as those of the two-droplet system [2] to investigate

the effect of the presence of the third droplet. The interaction effects from the downstream or upstream

droplet are identified. The transport rates of droplets are reduced from the values for an isolated droplet,

and values for the downstream droplets are profoundly less than those for the lead droplet. The difference in

transport rates is large between the first two droplets; however, it becomes insignificant between the second

and the third droplets since both downstream droplets are situated in a very low convective environment.

The modifications to the transfer correlations for an isolated droplet needed to account for the interaction

effects are determined.

INTRODUCTION

In realistic spray situations, the fuel is usually introduced into the combustor as a stream of liquid that

breaks into droplets. The droplets subsequently vaporize in the convective gas stream to form the air-fuel

mixture. Typically, the fuel is of sufficiently low volatility that vaporization is an important controlling

factor in the estimation of combustion rates. Usually, in the dense-spray regions such as regions near the

fuel nozzle, the droplet spacing is so small that the interaction effects would modify the droplet behavior

significantly. In order to obtain the qualitative modification of the transfer coefficients in the practical

dense spray calculation, it is then necessary to consider the interactions among a stream of multiple moving

droplets. However, a detailed and accurate simulation of hundreds of droplets will be very time consuming

and very difficult to perform. In fact, the behavior of trailing droplets, which follow the first two or three

droplets, can be estimated from that of the first two or three droplets on account of the periodical nature

of linear droplet arrangements. This research addresses the interaction of three vaporizing droplets moving

coUinearly which represents a model of an injected stream of fuel droplets.

There is a lack of a detailed investigation of multiple-droplet-interactions involving variable properties,

transient heating and internal circulation of droplets in the literature. Even the simplified numerical computations of three-droplet dynamics are rarely found. Tal et al. [4] used a multisphere cylindrical cell model to

study hydrodynamics and heat transfer in assemblages of spheres. They found the hydrodynamic solution

and Nusselt number, defined by using average bulk temperature of the cell unit inlet, to be perodic. Tong and

Chen [5] have included vaporization in Tal's model to investigate the effects of droplet spacing on heat and

mass transfer of droplets in a liquid droplet array. A three-droplet array in a cylindrical duct has been used

to obtain correlations between Nusselt number and local ambient properties for each droplet. The cylindrical

cell model is somewhat idealized and some assumptions must be imposed on the cell boundaries. Hence, their

results must be verified by calculations from an advanced model. Kleinstreuer et al. [6] used a finite-element

microscale analysis to find the drag coefficients of interacting spheres in a linear array and a boundary-layer

analysis for vaporizing droplets to simulate coupled transfer processes for three interacting droplets in a

719
one-dimensional trajectory. Their solution is basically the combination of three single-droplet-solutions with an "effective approach stream temperature" to account for interactions.

In the present study, the momentum, heat and mass transfer of three interacting, vaporizing droplets

are taken into account and the relative motion due to different drag forces that the droplets have experienced

is included. We aim to understand the wake effects on the transport rates of the downstream droplets. The

primary emphasis will be placed upon the interaction effects due to different initial spacings.

The schematic flow configuration is presented in Figure 1 where the flow passing over three vaporizing

droplets moving in tandem is shown. The flow is laminar and axisymmetric with initially uniform ambient

conditions specified by U^,T^, Pg ^otP'oo^ Y/.oo = 0. The initial droplet spacing is also prescribed. Dij

is the non-dimensional droplet spacing (with respect to the initial radius of the first droplet) between the i^'^ droplet and the j"* droplet. The frame of reference is fixed to the center of the lead droplet. The problem can be viewed as an impulsively-started flow over a fixed droplet and two moving downstream droplets aligned

in tandem.

The full consideration of forced convection of the gas phase, transient deceleration of the flowfield

due to the drag force, internal circulation and transient heating of the liquid phase, variable properties,

and transport processes occurring at the vaporizing droplet interface required to solve the Navier-Stokes

equations, energy and species equations, combined with appropriate boundary conditions simultaneously.

Also, in order to consider the moving boundaries due to surface regression and relative droplet motion, a

general method of generating boundary-fitted coordinate systems is required.

The axisymmetric governing-equations in cylindrical coordinates and their corresponding finite-difference

equations, and numerical procedures are given in Chiang [3]. The computer codes employed in our previ- ous research [1, 2] have been modified to deal with the present three-droplet arrangement The modification

involves the grid generation routine as well as the routines which handle the relative motion and spacings

between droplets. The calling sequence and parameter transfer among subroutines have been adapted to accommodate new variables and the increase in memory size. Unless otherwise stated, the nomenclature remain the same as employed in previous papers [1, 2, 3]. RESULTS AND DISCUSSION

Six production runs simulating three droplets, with the same droplet size but with different initial

droplet spacings, moving collinearly have been performed. The values of physical parameters employed in

each case are the same as the base case of the two interacting droplet-study [2]. The values of initial spacings

in each case are given in legends of the figures.

The computations are performed on a CRAY Y-MP supercomputer. Figure 1 also shows the typical

grid distributions at the beginning and at the final computational time for the case of droplet coalescence.

It is noted that the computation is stopped when the droplet spacing is reduced below 2.6 or whenever the

downstream droplet approaches the outer computational boundary, since the grid generation routine may

generate overskewed grid system under these conditions. Often the computation terminated at a very early

stage of the droplet lifetime during which most of heat flux to the droplet surface would be transferred to

the interior of the droplet. Hence, the effective transfer number {Bh = Cp'^ fUmC^oo ~ '^^){^ ~ ^)/^'s) is

quite small. The amount of mass due to the evaporation is insignificant, as a result the evaporation rates of

droplets are not reported here. The time scale used in the following discussion is the gas-phase hydrodynamic

diffusion time scale.

The representative results to characterize the droplet behavior are summarized below.

The global contours of results from Case 1 (with D12=D23=12) and Case 4 (D12=:D23=6) are compared

in order to study the effect of initial spacing on the flow fleld. The vorticity distributions and isotherms for

the gas and liquid phases are presented in the top and bottom portions of each plot of Figure 2, respectively.

The results of Case 1 at two different times are portrayed in the upper two plots. At the early time,

droplets experience the convective effect such that high vorticity gradients occur at the front portion of the

droplet, and an asymmetric distribution is developed. The flow field surrounding each droplet resembles

that surrounding an isolated droplet [1]. At five gas-phase diffusion times, the D12 is reduced from 12 to

2.96. The vorticity convected downstream from the first droplet has directly touched the second droplet and

shifted the vortex center to the equatorial plane. The second droplet is well protected by the vorticity wake

of the lead droplet, while the third droplet, spaced 9.5 droplet radii away from the second droplet, behaves

720
qualitatively as an isolated droplet.

For the case of small initial droplet spacing (Case 4), the interaction effects are expected to be strong.

The velocities approaching the second and the third droplets are more or less similar but are considerably

smaller than that approaching the lead droplet. As a result, the convective effect on transport is subdued.

In both cases, the vorticity distribution for the lead droplet is distorted by the approaching of the second

droplet.

The isotherms demonstrate thermal-transport interactions among three droplets and the gas phase. The

three droplets show the same isolated-droplet-isothermal-pattern at the early time. The third trailing droplet

always exhibits the same temperature contour as that of an isolated droplet, even though the approaching

temperature and velocity have decreased from the values upstream of the lead droplet. The case with

small initial spacing shows quite different thermal transport mechanisms. The temperature gradients for

the downstream droplets are smaller than those for the lead droplet due to the action of the thermal wake.

The thermal boundary layer thickness increases in the downstream direction along the surface for the lead

droplet. This behavior is opposite for the third trailing droplet in the close spacing case. The variations of Nusselt number along the gas/liquid interface of the droplets for Case 1 and Case 4 are shown in Figure 3. In Case 1, the decrease of Nusselt number at the front stagnation portions of the trailing

droplets are caused by the cold fuel/air mixture convected downstream from the upstream droplets. The

qualitative variations of Nusselt numbers of the second droplet and the third droplet in Case 4 are totally

different from those in Ccise 1. The details have been discussed in Chiang[3]. The third droplet shows the

same Nusselt number distribution as for the second droplet except at the separation region, where the third

droplet possesses a higher value of Nusselt number due to the angular diffusion of heat flux from the free

stream. Similar behavior for Sherwood numbers are observed for the three droplets.

Figures 4 and 5 present the time variations of drag coefficients and Nusselt numbers, respectively. As

we expect, the transport rates of droplets are reduced from the values for an isolated droplet, and values

for the downstream droplets are profoundly less than those for the lead droplet. The transport rates for the

third droplet are lower than those for the second droplet, except when the lead droplet's interaction with the

second droplet becomes very strong. The results of Case 4 show that the second and third droplets exhibit

almost identical behaviors. The interaction between the downstream droplet and the lead droplet becomes

very significant for the small initial spacing case. The variations of trajectory with time and drag coefficients vs. instantaneous Reynolds number for

the cases of different initial spacings are illustrated in Figure 6 (Cases 1, 2, and 3) and Figure 7 (Cases 4,

5, and 6), respectively. Results of Case 1 with large initial spacing (D12=D23=12) indicate that the drag

coefficient is smaller for the second droplet than for the lead droplet and still smaller for the third droplet.

The major decrease in drag occurs in the first two droplets. In Case 2, the drag coefficients of the lead and

the second droplets are significantly reduced (comparing curves 7 and 10, and 8 and 11, respectively) due to

the strong interactions when the first two droplets are spaced only two diameters away. Similar trends occur

for the downstream droplet pairs of Case 3 (comparing curves 8 and 14, and 9 and 15, respectively). The

third droplet in Case 2 has a higher drag coefficient than that of the second droplet since the second droplet

is better shielded by the droplet before it. The D23 thus increases with time. Also, note that the drag

coefficient of the third droplet, which is spaced far away from the second droplet, seems to be independent

of D12 as indicated in curves 9 and 12. However, it strongly depends upon D23 as illustrated in curves 9,

12 and 15.

The comparison between results of Figures 6 and 7 leads to the clear conclusion that as initial spacings

decrease, the transport rates decrease correspondingly. The drag coefficients for the downstream droplets

are all collapsed together. However, at the final calculations when three droplets interact extensively with

neighbors, the second droplet experiences not only downstream interactions from the lead droplet but also

upstream interactions from the third droplet. As a result, the second droplet possesses the lowest drag coefficient. Note the interaction from the upstream droplet is considerably larger than that from the downstream

droplet. The lead droplet in Case 2 has the lowest drag coefficient (curve 10) among the lead droplets of

three cases since the interaction from the second droplet is the strongest (spacing D12 is the smallest). The

behavior of the spacing variation is qualitatively similar to that of large initial spacing Ccise. The variation

of D23, for the cases of approaching droplets, is always negligible except during the final calculation period.

In order to investigate the effect of D23 on the first and the second droplets with a given D12, the

comparisons of drag coefficients for the first two droplets with variable D23s are presented in Figures 8

721
and 9. The comparisons of results with the same D12 in the two-droplet-only arrangement are also presented.

For the cases of small initial-D12, the main droplet interactions occur between the first two droplets. The

presence of the third droplet does not affect the lead droplet but it reduces the Sherwood and Nusselt

numbers of the lead droplet by less than 5 % in magnitude. The drag coefficients and Sherwood numbers

for the second droplet decrease as the D23 decreases, though the percent change is small. For the Ccises of

large initial D12, the behavior of the first droplet becomes insensitive to the presence of the third droplet.

An interesting observation is that the presence of the third droplet at a sufficient distance from the second

droplet, with insignificant hydrodynamic interaction between the second and the third droplet, may increase

the drag coefficient of the second droplet (comparing curves 4 and 5 in Figures 9). The second droplet in

the two-droplet-only arrangement receives the interaction of recirculating-thermal-entrainment from the free

stream; as a result, the transfer number is high. Therefore, the surface blowing is enhanced and the friction

drag is reduced. The addition of the third droplet to the system moves this thermal effect to the third

droplet; hence, the transport rates of the second droplet are recovered. However, for the case of a small D23,

the strong interaction from the third droplet reduces the transport rates of the second droplet by a large

magnitude (curve 6).

A nonlinear regression model using least squares has been employed to find the correlations between

the transport rates of an interacting droplet and the corresponding transport rates when droplet is isolated.

The generalized form of numerical correlations for drag coefficient and Nusselt number of each droplet can be expressed as

where the constants are given in the Table 1. The correlations are valid for droplet spacing ranged from 2.7 to 12 and for Reynolds number ranged

from 90 to 130. Some correction factors, with different combinations of D12s and D23s, computed from

above correlations are presented in Table 2. The correlations basically predict the right trends of interacting

effects on droplet transport rates. The influence of D12 on the third droplet is insignificant as indicated by

the small exponents of D12.

CONCLUSIONS

The detailed behaviors of three vaporizing, interacting droplets for the cases of different initial spacings

have been investigated carefully. The results indicate that the interacting effects are strongly dependent

upon the initial droplet spacings. The general qualitative conclusions drawn from the two-droplet study [2] can be applied to two neighboring droplets in the three-droplet analysis.

Results for the cases of sufficiently large spacing (above approximately 6 droplet diameters) show that

the flow field of each droplet is qualitatively similar to that of an isolated droplet, although the transport

rates are reduced along the downstream direction due to the cascade effect of the wake. The major drops in

transport rates occur in the first two droplets. For the cases of small droplet spacings (less than approximately

3 droplet diameters), the flow field of downstream droplets can be significantly altered due to the interaction

effects from the upstream droplets. Usually, the second droplet has the lowest drag coefficient since it receives interactions from both neighboring droplets. However, the difference in transport rates between the

second and the third droplets is not significant since both droplets are fully protected by the wake of the

first droplet. The effect of D23 on the behavior of the first droplet is insignificant. However, depending

upon the values of D12 as well as D23, the effect of D23 on the behavior of the second droplet may become

significant. The correlations for heat transfer and droplet dynamics have been developed and are applicable

in a one-dimensional droplet spray calculation.

(1)

722
ACKNOWLEDGEMENTS

This work has been supported by the Air Force Office of Scientific Research under grant No. 90-0064

with Dr. Julian TishkofF acting as the technical monitor. The support of the San Diego Supercomputing

Center under a Block Grant of the Office of Academic Computing at UCI is greatly appreciated.

REFERENCES

[1] Chiang, C. H., Raju, M. S. and Sirignano, W. A. "Numerical Analysis of Convecting, Vaporizing Fuel

Droplet with Variable Properties", AIAA Aerospace Sciences Meeting, Paper 89-0834. (1989). Also to

appear in Int. J.' Heat Mass Transfer (1991).

[2] Chiang, C. H. and Sirignano, W. A., "Numerical Analysis of Interacting, Convecting, Vaporizing Fuel

Droplet with Variable Properties", AIAA Aerospace Sciences Meeting, Paper 90-0357 (1990). Also to

be submitted to Int. J. Heat Mass Transfer.

[3] Chiang, C. H., "Isolated And Interacting, Vaporizing Fuel Droplets: Field Calculation With Variable Properties", Ph.D. Dissertation, University of California, Irvine, Dept. of Mechanical Engineering

(1990).

[4] Tal (Thau), R., Lee, D. N.and Sirignano, W. A. "Hydrodynamics and Heat Transfer in Sphere of Particle

Assemblages - Cylindrical Cell Models", Int. J. Heat Mass Transfer, Vol 26 ,

, No. 9, pp. 1265-1273

(1983).

[5] Tong, A. Y. and Chen, S. J., "Heat Transfer Correlations for Vaporizing Liquid Droplet Arrays in A

High-Temperature Gas at Intermediate Reynolds Number", Int. J. Heat h Fluid Flow, Vol 9, No. 2,

pp. 118-130 (1988)

[6] Kleinstreuer, C, Chiang, H. and Wang, Y.-Y., "Mathematical Modelling of Interacting Vaporizing Fuel

Droplets", HTD-Vol. 106, Heat Transfer Phenomena in Radiation, Combustion and Fire, pp. 469-477

(1989)

Table 1: Constants for Cd and Nu Correlations

a b c

C'di

%r

2.092

2.084

0.798

0.324

1.218

0.811

-1.575

-0.482

0.012

-0.266

-0.175

-0.093

-0.393

-0.351

-0.320

-0.098

-0.398

-0.277

723
Table 2: Correction Factors for Cd and Nu with Different Combinations of D12 and D23

C'oi Cos Ati,

12.UU 1 O AA A OQjI u.yo4 A TX1 A Odd U.ooy A 7Afi u.#ud A <!TA U.D/D

A AA u.y/y U.DOD A KIT A QAA A A1 9 A AAfi

1 O AA 4.UU n 07ft u.y /D U.D14 f\ ATI A Ql^A U.O04 U.040 U.OOi

C AA 1 O AA Au.yoo A ftlQ U.OOO U.DO/ A fi/IO U.ODO U.DOO

6 on 6 00 0 938 0.532 0.541 0.831 0.563 0.582

6.00 4.00 0.928 0.460 0.477 0.824 0.487 0.532

4.00 12.00 0.911 0.554 0.634 0.824 0.644 0.642

4.00 6.00 0.883 0.431 0.543 0.812 0.531 0.566

4.00 4.00 0.863 0.344 0.480 0.804 0.449 0.514

2.70 12.00 0.835 0.461 0.636 0.805 0.619 0.628

2.70 6.00 0.784 0.312 0.545 0.791 0.498 0.550

2.70 4.00 0.746 0.207 0.482 0.783 0.410 0.496

FLOW CONFIGURATION - SCHENUTIC

Time = 0.00 R] = 1.00 . R2 = 7.00 , K3 = J.OO Spacin£l2 = 6.00 , Spacing23 = 6.00

Time = 2.40 Kl = 1.00 . K2 = J.OO , K.I = /.OC Sparing12= 2.78 , Spociin;2} = «.0^

Fig. 1: Flow field configuration and grid

distribution at two different times.

CAS AND LIQUID PHASE VORTlCITy

iMl KCE-OI Mm ^JJf^aoUft USUI

CAS AND LIQUID PHASE ISOTHERMS

t.8 15.7 E.7 29.6 36«

ruiu i.to

«;« I.OC , «7> , KJ' IJK .Sfadrngll' $.73 ,Sp»dMt7i m nj4

GAS AND LIQUID PHASE VORTICITY

Cnw l«K>v^ Uffi^l Mr. -lifi « Mu >JTM1

0.0.

CAS ANT5 UQUID PHASE ISOTHERMS

I i I I Sf 140 I9J 24.4 29.« Time • S.On

HI' 1.00 , 112 m 1.00 , K.im 1.00 .SparingUm IJH ,Sp<u:i«ti}m tJO

Imiaal KM, K3'l, KM, Sporimgll'll, S^euit23-I2 GAS AND LIQUID PHASE VORTlCITy

o.n

2.7.

5.4.

GAS AND LIQUID PHASE ISOTHERMS

Onw bvrvii S«M.r; M.- l.a£.c:H4i I JU .^i

R.9 i:- 174 21.7 26.0 Timr m 2.00 Kf(l)* , KH2I-: 9IJ9 , Kr;}jm 9IJf */• i.00 , /.«(/ , IM .S^irivM" J.7< ,^uu«2J> 5.W

Imilial KI.J. K2^l. *.<•/, Sfcnmt)2'*, Spcnitt2SM

Fig. 2: Vorticity and isotherm contour plot of gas and liquid

phases with different initial droplet spacings.

724
Vmim 2Jt C*M 1: Inlltal Oil «ia.OO, 023 «12M

C*M 4: InRtal 012 « 6.00, 021 m «.00 Lnd drapM, Cm* 1, R*(9 > M.«, t.p. M 2, leetl mki. M 140.0 LuddrepM, Cu«4, R*<l)> oaAt.p.at 130.3, toealmln. ri 141.4 SMond dnpM, Cm* 1, R«(l) • 80.7, t.p. 1 14S.0, loot mln. M isa7 SMonddnpM, C*w4,n«(l)>01J,i.p. •(1M.«,localinln.«l 0.0 Third dropM. CiM 1, R*(l) • tO.*, (.p. •« 100.1, local mln. at isaj

Third drapM, Caaa 4, R*(l)> 01.3, *.p.*l 107.1, bcalmki.al 0.0

9.5«

0. 18. 36. 54. 72. 90. 108. 126. 144. 162. 180. ANGULAR POSITION (DEGREE)

Figure 3: Surface Nunelt number distribution of the three dropleti with different initial dropicl (pacings.

1. Laad OropM, (InHtel D12<12, D23i12)

2. Sacend DropM, pnlllal D12s12, D23ts12)

3. Third Droptat, (Initial 012*12, D2ai12)

4. Laad Oroplat, (Initial 012:0,02330)

0. Sacond Oroplat, (Initial 012x0, 023iO)

0. Third OropM, (InMal 012:0, 023:0)

7. laolalad OropM

100

69.74 75.69 81.65 87.60 93.56 99.51 INSTANTANEOUS REYNOLDS NUMBER

Figure 4 Time variation of drag coeSicioits of the three dropteu

with differeit Initial droplet ipaclngi.

1. Laad DropM, (MiWal D12>1% D23e12)

2. SKond Diopla<, (Inlllal 012i>2, 023<12)

3. Thkd DropM. (Inlllal 012m12, 023<12)

4. Laad OropM, (InWal 012>«,023><)

0. Sacond Oreptal, (Inlllal 012«6, 023=6)

0. Third OropM, (Inlltal D12i6, D23>6)

7. iBotatad DropM

10.00

0.00 1.00 2.00 3.00 4.00 5.00 GAS HYDRODYNAMIC DIFFUSION TIME

Hgure 5 Time variation of Nusaelt numben of the three droplets Legends: with different initial droplet tpadngs. Case 1: Initial Rl=l, R2=l, R3 =l, D12=12, D23 =12, ReslOO

Case 2: Initial RI=1, R2=I, R3=l, D12=4, D23=12, Re^IOO

Case 3: Initial Rlrrl, R2=l, R3=l, D12=12, D23=4, Re=10a

I. D12, Case 1; 2. D23, Case 1; 3. D12, Case 2;

4. D23, Case 2; 5. D12, Case 3; 0. D23, Case 3;

7. On for the Lead Droplet, Case 1; 8. (j) for the Second Droplet, Case 1; 9. C) for the Third Droplet, Case 1; 10. C, for the Lead Droplet, Case 2; II. Ci, for the Second Droplet, Case 2; 12. Q for the Third Droplet, Case 2; 13. Cii for the Lead Droplet, Case 3; 14. Cj for the Second Droplet, Case 3; IS. Cij for the Third Droplet, Case 3; 18. C< for an Isolated Droplet; g 14.00 2.00

0.00 1.00 2.00 3.00 4.00 5.00 GAS HYDRODYNAMIC DIFFUSION TIME

99.63 93.65 87.68 81.70 75.72 69.74 INSTANTANEOUS REYNOLDS NUMBER

Figure 6: Time variBtion of droplet drag cocfndcnts und droplet spacings

for the cases of large initial droplet spacings (Cases 1.2 and 3).

725
Legends: Case 4: Iiiltlnl Rl =l, R2:=I, R3=l, D12=e, D23=6, Re=100

Case S: Initial Rlr^l, R2=l, R3=l, D12=4, D23=:8, Re=100

Case 8: Initial Rl=l, R2=I, R3=l, D12=6, D23=4, Re=100

I. D12, Case 4; 2. D23, Case 4; 3. DI2, Case S;

4. D23, Case 5; 5. D12, Case 6; 6. D23, Case 8;

7. Ci) Tor the Lead Droplet, Case 4; 8. Cj for the Second Droplet, Case 4;

9. C) for the Third Droplet, Case 4; 10. G> for the Lead Droplet, Case S; II. Ct> for the Second Droplet, Case 5; 12. Cj for the Third Droplet, Case 5; 13. C) for the Lead Droplet, Case 8; 14. Ci for the Second Droplet, Case 8;

IS. Ci, for the Third Droplet, Case 8; 18. C, for an Isolated Droplet;

9O 0.00 T r

0.00 1.00 2.00 3.00 4.00 5.00 GAS HYDRODYNAMIC DIFFUSION TIME

99.54 93.58 87.62 81.68 75.70 69.74 INSTANTANEOUS REYNOLDS NUMBER

Figure 7: Time variation of droplet drag coefficiena and droplet spacinga

for the aaes of snail initial droplet spacing! (Cases 4, S and 6).

100

U. 0.60

11] oQ 0.40

<0.20

ocO 0.00

1. Tha Flrat Droplat ol Two-DroplM Caleulallen, D12s4

2. Tlia Flnt OropM ol Tliraa-OropM Calculsllon, 012:4, 023=12

3. Tha FIret Droplat of Thra«.Oropl«t Calculation, 012=4, 023=6

4. Tba Sacond Droplat of Two-Droplat Calculalton, 012=4

S. Tha Second Oroplal of T)ira*-Oropl*l Calculation, 012=4, 023=12

6. Tha Sacond Droplat of Thraa-Droplat Calculation. D12=4, D23=6

1. Tha Fhvt Dreplat of TVm-Dioplat Calculallon, 012=12

2. Tha nrat Oraplal of Thma-DropM CaleutaHon, 012=12, 023=12

1. Tha FIrat Dioplat of Thraa-Droptal Calculation, 012=12, 023=4

4. Tha Swond OropM of Two-OropM Calculation, 012=12

R. Tha Saeond OrepM of Thraa-Oreplat CalcuUUon, 012=12, 023=12

e. Tha Saeond Droplat of Thraa-DropM CaleutaHon, 012=12, 023=4

2.00

89.69 91.66 93.63 95.60 97.57 99.54 INSTANTANEOUS REYNOLDS NUMBER

70J5 76.37 82.18 88.00 93.82 99.63 INSTANTANEOUS REYNOLDS NUMBER

Figure 8 Time variation of drag coefTicicntfl of the first two droplets

for D12 « 4 and variable D23s.

Figure 9 Time variation of drag coeffklents of the first two droplets

for DI2 - 4 and variable D2V

726
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 81

SOLUTIONS OF POLYDISPERSE SPRAY SECTIONAL-EQUATIONS VIA

A MULTIPLE-SCALE APPROACH: AN ANALYSIS OF A PREMIXED

POLYDISPERSE SPRAY FLAME

I. Silverman, J.B. Greenberg and Y. Tambour

Faculty of Aerospace Engineering

Technion - Israel Institute of Technology

Haifa, Israel

ABSTRACT

New solutions are presented for polydisperse spray sectional-equations. Simultaneous evaporation

and coalescence processes of the spray droplets are considered. The mathematical procedure takes

advantage of a multiple-scale approach in which different scales are applied to the different physical

processes (convection and evaporation). Two small spray parameters are identified and used for

expansion pmrposes. A special zonal density model is developed, thus enabling approximate analytic

solutions to be extracted from the governing equations for the entire system. The method is employed to analyze adiabatic premixed polydisperse spray flames in a one-dimensional flow field. Computed results are presented for flame location versus flame velocity, the influence of droplet evaporation rate (with

and without coalescence) on flame location, and the axial variation in temperature and chemical species

profiles. As to the fuel spray, the evolution in droplet size distributions, for the various size

sections, along the axial direction, is analyzed. Finally, all analytical results are compared with

numerical solutions for the governing equations which are obtained here using standard finite

difference methods. The good agreement between the analytical solutions and the numerical ones

indicates that a multiple-scale approach may be an important satisfactory tool for the analysis of

polydisperse sprays which undergo simultaneous evaporation and coalescence processes.

INTRODUCTION

The theoretical modelling of the behavior of a spray flame is of considerable importance in the

analysis, design, and improvement of combustion equipment. In most industrial systems, the fuel spray

is of a wide range of droplet sizes, i.e. it is polydisperse. The combustion processes are controlled

by a 'cloud' or 'group' behavior of the spray rather then single droplet behavior [1-15]. The analysis

of the evolution in local droplet size distributions, enables one to determine the amounts of fuel

vapor which are locally released via the evaporation process of the 'cloud' as a whole, whereas the

combustion chemical reactions are controlled by the local fuel vapor/oxygen concentration ratios and

the local temperature and pressure. Thus, the theoretical modelling involves solutions of polydisperse

intercoupled spray equations, which are coupled also to the flow field, energy, and chemical species

conservation equations.

A representation of the polydisperse spray by a set of sectional-conservation equations [16-21]

enables one to deal with a fairly small number of conservation equations. However, in relatively dense

sprays coalescence may play an important role [20-22]. The equations describing dense sprays are nonlinear due to the coalescence terms, and therefore difficult to solve. Thus, the purpose of the

present study is to present a mathematical procedure which takes advantage of a multiple-scale approach

in which different scales are applied to the different physical processes: convection and evaporation.

727
enabling approximate analytic solutions to be extracted from the governing equations for the entire

system.

The problem to be analyzed in the present study is a one-dimensional adiabatic laminar premixed

polydisperse spray flame. For this geometrically straightforward type of combustion wave, a wide range

of phenomena have already been investigated, including effects of transport properties, detailed

chemical kinetics, etc. on such parameters as flame speed, temperature, ignition and extinction limits,

to name but a few (see for example, the review paper by Clarke [23] and relevant material in

Dixon-Lewis [24], Williams [25], Buckmaster and Ludford [26]). However, in all these studies a gaseous

combustible premixture was assumed.

In fact, within the context of the 1-D premixed type flame a few attempts have been made to accommodate additional effects induced by the presence of fuel droplets [13, 27-30]. These analyses

have usually dealt with single evaporating droplets and use was made of an average droplet size. Quite

recently, Lin et al. [31] have proposed a theory of laminar flame propagation in dilute sprays. However, they have considered only monodisperse sprays, with no droplet-droplet interaction allowable.

A study of a premixed polydisperse spray flame in which droplet-droplet interactions withm the

spray were considered, was recently presented by the authors [32]. Using an asymptotic analysis, the

influence of the spray on the burning velocity and on the flame temperature was examined. In the

present study, however, a multiple-scale analysis will be exploited to provide tractable solutions for

the major variables of the problem, such as the axial variation in temperature, chemical species

profiles and the evolution in droplet size distributions for the various size sections along the axial

direction. The influence of droplet evaporation rate, with and without coalescence, on flame location

will also be examined.

CALCULATED RESULTS

The theory and the solution procedure have been presented elsewhere [37]. In order to provide some

measure of the predictive capability of the aforedescribed solution a few computed results are presented in Figs 2-10. In the figures the solid lines represent the results of the analytical solution

whilst the broken lines are for profiles obtained by solving the governing equations numerically using

standard finite-difference methods. As can be seen for the problem at hand the agreement is very good.

In Fig. 2, the variation of temperature, chemical species concentrations and liquid-fuel

integral-mass concentration along the axial direction are shown. (The evolution in droplet

concentrations m each size section due to vaporization is presented in Figs. 3a,b, and will be

discussed later.) As one approaches the flame front, the integral-mass concentration of the

fuel-droplets decreases due to vaporization (see Fig. 2), while the concentration of fuel vapors

increases in the beginning (defined as zone I in Fig, 1), and then decreases in zone 11 due to diffusion and consumption of ftiel vapors by the combustion process. At the flame front the oxidant

concentration approaches zero whereas the concentration of CO^ reaches its maximum value. Since the

flame is ftiel-rich, fuel droplets that pass the flame front continue to vaporize m the "post-reaction

zone" (zone III) and the concentration of fuel vapors starts to build-up again, since no more oxidant

is left to react with.

The results presented in Fig. 2, which were discussed above, were computed for a vaporizing spray

without coalescence, i.e. for G=0. The effects of coalescence are demonstrated in Figs. 4 and 5a,b and

are discussed next. Since coalescence hinders vaporization, the flame front is found to be further

downstream at x =10.1, in Fig. 4 (computed for 9=0.867), whereas for 6=0, i.e. without coalescence.

728
x^=7.8 (see Fig. 2).

The droplet concentration in each size section decreases with the distance x due to vaporization

of droplets, as shown in Figs. 3a and 3b, which were computed without coalescence (i.e. for 9=0). When

coalescence is considered (9=0.867) an initial increase in the concentrations of section VI, VII and

Vni is observed (see Fig. 5b). This is due to coalescence of droplets in lower sections resulting in

the production of larger droplets, i.e. causing a shift to the "right" of the size spectrum of

droplets. Such behavior is as anticipated, having been observed experimentally and verified

theoretically [20, 33]. This is also consistent with the steep drop in droplet concentrations in the

lower sections I, 11, in and IV (see Fig. 5a) whereas a rather gradual decrease in droplet

concentrations in these sections is noticeable when only vaporization is considered (Fig. 3a).

As to the comparison of: (i) the effects of vaporization only with (ii) the effects of combined

vaporization and coalescence, on flame location, these are shown in Figs. 6 and 7, respectively. The

larger the evaporation rate, i.e. the larger the value of e, the shorter the distance x^ which is

required to build-up conditions so that a flame front will be located at x=x^ For a given vaporization

rate of e=0.125, at low coalescence rates for which 9 is smaller than unity (9<1), coalescence causes a very slight shift of the flame downstream (see Fig. 7). This is due to the fact that coalescence

hinders vaporization. For 9»1 and for higher values of 9, the change in flame location is significantly

noticed.

In Fig. 8, we illustrate the way that the flame location, x^ behaves as a function of the flame

velocity. This curve is drawn for a fixed total equivalence ratio, i.e., the ratio of the initial fluxes of fuel Qiquid + vapor) and oxygen. However, the effective equivalence ratio, i.e., the ratio

of the fluxes of fuel vapors only and oxygen varies along x and governs the flame. It is interesting to note that unlike the case of the steady premixed gaseous flame in which the total equivalence ratio

uniquely determines the flame velocity and location, for the steady premixed spray flame this same

ratio leads to a band of permitted flame velocities to each of which corresponds a single flame

location. It is precisely this band of values that is shown in Fig. 8. For a steady flame, the flame

velocity is equal to the upstream velocity at x=0. Thus, for each permitted upstream velocity at x=0,

with a fixed total equivalence ratio, the downstream production of fuel vapors (for a given spray, it's size distribution, and e and 9) will build-up conditions in terms of the effective equivalence ratio,

at a single downstream distance x=x^ at which the flame will be located.

Next, the comparison between analytical solutions and numerical results is discussed. As far as evolution in droplet concentrations in each size section is concerned, the agreement is excellent for

the process of droplet evaporation (see Figs. 3a,b). For simultaneous vaporization and coalescence, the

deviation of the analytical solutions from the numerical ones grows as the rate of coalescence

increases, i.e. as 9 increases. These quantitative differences can be attributed to the order of

approximation assumed in the analytical calculation. For example, since we assumed 9«e«l, it is not

surprising that analytical results calculated for a "large" value of 9, such as 9=0.867 deviate ft-om

the numerical solutions (see Figs. 5a,b). Better agreement is obtained for smaller values of 9,

although still larger than e, such as 9=0.347; (compare the axial downstream evolution of pointwise

spray histograms for 9=0.347, Fig. 9 with those of 9=0.867, Fig. 10). Nevertheless, a comparison

between the numerical and analytical solutions demonstrates that even for relatively "large" values of

9, the analytical solutions remain qualitatively satisfactory (see for example Fig. 4).

729
CONCLUSIONS

To summarize, the multiple-scale analysis that has been presented here has proven to be a

satisfactory tool for the study of one-dimensional premixed polydisperse spray flames. All analytical

results have been compared with numerical solutions of the governing equations and show good agreement

with them. This was foimd for the variation with distance of many physical properties such as temperature, chemical species concentrations, and in the analysis of the effects of vaporization and

coalescence on flame location.

REFERENCES

1. Chigier, N.A., Energy Combustion and Environment, McGraw-Hill, 1981.

2. Chigier, NJE., "Group Combustion Models and Laser Diagnostic Methods in Sprays (A Review),

Combustion & Flame, 51, pp. 127-139 (1983).

3. Chigier, N.A., Ungut, A. and Yule, AJ.JParticle Size and Velocity Measurements in Flames by Laser

. Anemometer, Seventeenth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, pp. 315-324 (1979).

4. Chiu, H.H. and Liu, T.M.,"Group Combustion of Liquid Droplets", Combustion Science and Technology,

17, pp. 127-142 (1977).

5. Chiu, H.H., Kim, H.Y. and Croke, EJ.Jnternal Group Combustion of Liquid Droplets, Nineteenth

Symposium (International) on Combustion, The Combustion Institute, pp. 971-980 (1982).

6. Labowsky, M. and Rosner, DJE., "Group Combustion of Droplets in Fuel Clouds: Quasi-Steady

Predictions", Evaporation-Combustion of Fuels, American Chemical Society, (1978), pp. 63-79

(1978).

7. Labowsky, M., "Calculation of the Burning Rates of Interacting Fuel Droplets", Combustion Science

and Technology, 22, pp. 217-226 (1980).

8. Levebvre, A.H., Gas Turbine Combustion, McGraw-Hill (1983).

9. Twardus, E.M. and Brzustowski, T.A.,"An Experimental Study of Flame Spread and Burning in Arrays

of Monosize Hydrocarbon Droplets", Combustion Science and Technology, 17, pp. 215-225 (1978).

10. Presser, C, Gupta, A.K., Santoro, RJ. and Semerjian, H.G., Droplet Size Measurements in a

Swirling Kerosene Spray Flame by Laser Light Scattering, ICLASS-85 Meeting, 2 (1985).

11. Presser, C, Gupta, A.K., Santoro, RJ. and Semerjian, H.G., Velocity and Droplet Size

Measurements in a Fuel Spray, AIAA-86-0297, AIAA 24th Aerospace Sciences Meeting, January 6-9,

(1986), Reno, Nevada.

12. Beretta, F., Covaliere, A. and D'Alessio, A., "Drop Size and Concentration in a Spray by Sideward

Laser Light Scattering Measurements", Combustion Science and Technology, 36, pp. 19-37 (1984).

13. Aggarwal, S.K. and Sirignano, W.A., "Unsteady Spray Flame Propagation in a Closed Volume",

Combustion & Flame, 62, pp. 69-84 (1985).

14. Correa, S.M. and Sichel, M., The Group Combustion of a Spherical Cloud of Monodisperse Fuel

Droplets, Nineteenth Symposium (International) on Combustion, The Combustion Institute, pp.

981-991 (1982).

15. Choudhury, P.R. and Gerstein, M., Analysis of a Fuel Spray Subjected to Coupled Evaporation and

Decomposition, Nineteenth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, pp. 993-997 (1982).

16. Tambour, Y., "A Sectional Model for Evaporation and Combustion of Sprays of Liquid Fuels", Israel

J. of Technology, 18, pp. 47-56 (1980).

17. Tambour, Y., "Transient Mass and Heat Transfer From a Cloud of Vaporizing Droplets of Various Size

Distributions: A Sectional Approach", Chem. Eng. Commun. 44, pp. 183-196 (1986).

18. Tambour, Y., "Vaporization of Polydisperse Fuel Sprays in a Laminar Boundary Layer: A Sectional

Approach", Combustion & Flame, 58, pp. 103-114 (1984).

730
19. Tambour, Y., "A Langrangian Sectional Approach for Simulating Droplet Size Distribution of

Vaporizing Fuel Sprays in a Turbulent Jet", Combustion & Flame, 60, pp. 15-28 (1985).

20. Tambour, Y., "Coalescence of Vaporizing Kerosene Fuel Sprays in a Turbulent Jet", Atomisation and

Spray Technology, 1, pp. 125-146 (1985). See also: "Simulation of Coalescence and Vaporization of

Kerosene Fuel Sprays in a Turbulent Jet: A Sectional Approach", AIAA/ASME/ASEE 21st Joint

Propulsion Conference, Monterey, 1985. AIAA Paper No. 85-1315.

21. Greenberg, J.B. and Tambour, Y.,"Far-Field Coalescence Effects in Polydisperse Spray Jet Diffusion

Flames", Twenty-First Symposium (International) on Combustion, The Combustion Institute, pp.

655-663 (1986).

22. O'Rourke, PJ. and Bracco, F.V., Modelling of Drop Interaction in Thick Sprays and a Comparison

with Experiments, Conference on Stratified Charge Automotive Engines, The Institute of Mechanical

Engineers, London (1980).

23. Clarke, JJ., "Parameter Perturbations in Flame Theory", Prog. Aerospace Sci., 16, pp. 3-29

(1975).

24. Dixon-Lewis, G., Computer Modeling of Combustion Reactions in Flowing Systems with Transport,

Combustion Chemistry, ed. W.C. Gardiner, Jr., Springer-Verlag, New York, pp. 21-125 (1984).

25. Williams, F.A., Combustion Theory, 2nd Edition, The Benjamin/Cummings Publishing Co., Menlo Park,

California (1985).

26. Buckmaster, J. and Ludford, G.S.S., Theory of Laminar Flames, University Press, Cambridge (1982).

27. Williams, F.A., Monodisperse Spray Deflagration, Liquid Rockets and Propellants, ed. LE.

Bollinger, M. Goldsmith and A.W. Lemmon, Jr., Academic Press, London, pp. 229-263 (1960).

28. Seth, B., Aggarwal, S.K. and Sirignano, W.A., "Flame Propagation Through an Air-Fuel Spray Mixture

with Transient Droplet Vaporization", Combustion & Flame, 39 pp. 149-168 (1980).

29. Polymeropoulos, C.E., "Flame Propagation in a One-Dimensional Liquid Fuel Spray", Combustion

Science and Technology, 9, pp. 197-207 (1974).

30. Satil, P.B., Sichel, M. and NicoUs, J.A., "Analysis of Spray Combustion in a Research Gas Turbine

Combustor", Combustion Science and Technology, 18, pp. 21-31 (1978).

31. Lin, T.H., Law, C.K. and Chung, S.H., "Theory of Laminar Flame Propagation in Off-Stoichiometric

Dilute Sprays", International Journal of Heat and Mass Transfer, 31, pp. 1023-1034 (1988).

32. Silverman, I., Greenberg, J.B. and Tambour, Y., "Asymptotic analysis of a Premixed Polydisperse

Spray Flame", SIAM Journal on Applied Mathematics (1990), in press. 33. Yule, AJ., Ereat, P.R. and Ungut, A., "Droplet Sizes and Velocities in Vaporizing Sprays",

Combustion & Flame, 54, pp. 15-22 (1983).

34. Law, C.K. and Sirignano, W.A., "Unsteady Droplet Combustion with Droplet Heating - II: Conduction

Limit", Combustion & Flame, 28, pp. 175-186 (1977).

35. Law, C.K., "Adiabatic Spray Vaporization with Droplet Temperature Transient", Combustion Science

and Technology, 15, pp. 65-74 (1977).

36. Yule, AJ., Seng, C.Ah., Felton, P.G., Ungut, A. and Chigier, N.A., "A Study of Vaporizmg Fuel

Sprays by Laser Techniques", Combustion & Flame, 44, pp. 71-84 (1982).

37. Silverman, I., Greenberg, J.B. and Tambour, Y., "Solutions of Polydisperse Spray SectionalEquations via a Multiple-Scale Approach: An Analysis of a Premixed Polydisperse Spray Flame",

submitted to Atomization and Sprays.

o I > X X=X

J

Zone I Zone II Zone III

Evaporation Di f f usion/ Pos I reaction

Zone Heating Zone

Zone r

Homogeneous Flame

Mixture She e t

Fig. 1. Zonal subdomains of density model.

731
Species & Temperature Profiles

f 1500

t 2000

\ ISOO

I '000

soo

O.O 3.0 e.O P.O tM.O 15.

0

Parameters:

c = 0. tZS

\e\ = 0.000 U0= 0.ZB5

Legend: //umeKcai Solution _ _ _ _ Analytical Solution

Fig. 2. Variation of temperature,

chemical species concentrations,

and liquid-fuel integral-mass

concentration, along the axial

direction; no coalescence.

Profiles of Spray Sections

t

i.oio

0.00 3.00 6.00 9.00 12.00 15.00 X

Paramsters:

t " 0.1 zs

101 = 0.000 uo= o.zos

Legend:

Profiles of Spray Sections

. //umericat Solution Analytical Solution Section Wo. n

t.oio

.000

0.00 3.00 6.00 9.00 IZ.OO 15.00 X

Parameters:

t = 0. (25

|el = 0.000 uo= o.zes

LegeTuL

_ Wumericai Sotution „ „ vlTiati/ticat Solution n Section No. n

Fig. 3. Evolution in droplet concentrations in

each size section; no coalescence (a) Sections I

to IV, (b) Sections V to X.

Fig. 4. Variation of temperature,

chemical species concentrations and

liquid-fuel integral-mass concentration,

along the axial direction; including

coalescence.

Fig. 5. Evolution in droplet concentrations

in each size section; including coalescence

(a) Sections I to IV, (b) Sections V to X.

732
Co

o

o

o

10

I- >n

I

10

I

o 0

o o

Co Co

a

; • CD >o

d , V.

II

733
CO

c

HO

CO

Co

O O O O V>

E

o o-pi iJ to

o o

Co Co

o o

d

-d

d

N

Cod

10 <t5

CO

o

CO

P.

Co

(uSi/,)

hi

6

o ct o d

E

Cl ci o cf

o o

•M to

O O

Co Co

d d o o

03 Pi g d

"d

d

J-d

CO

II II

734
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 82

ON THE EVAPORATION OF BINARY-FUEL CLUSTERS OF DROPS

HAVING A LARGE VOLATILITY DIFFERENTIAL

K. Harstad and J. Bellan

Jet Propulsion Laboratory

California Institute of Technology

Pasadena, CA, U.S.A.

ABSTRACT

A formulation has been developed to describe the evaporation of binaryfuel, dense or dilute clusters of drops. The binary fuel is assumed to be made

of a solute and a solvent whose volatility is much lower than that of the solute.

Convective flow effects, inducing a circulatory motion inside the drops, are

taken into account, as well as turbulence external to the cluster volume.

Results obtained with this model show that, similar to the conclusions

obtained for single, isolated drops, the evaporation of the volatile is

controlled by liquid mass diffusion when the cluster is dilute. In contrast,

when the cluster is dense, the evaporation of the volatile is controlled by

surface layer stripping, that is by the regression rate of the drop which is in

fact controlled by the evaporation rate of the solvent. These conclusions are

in agreement with existing experimental observations. Parametric studies show

that the above conclusions remain valid with changes in ambient temperature,

initial slip velocity between drops and gas and initial cluster size.

I. INTRODUCTION

Most of the liquid fuels used to provide power through combustion are

mixtures of several pure fuels. For this reason there has always been interest

in the study of the combustion of multicomponent liquid fuels. This interest was

hightened by the realization that, due to fuel scarcity, the next generation of

fuels will have to contain more of the lower end of the distillation batch. The

behavior of these heavier and more viscous compounds mixed with small quantities

of more volatile compounds (left from the distillation process) became

immediately the focus of many studies.

Since in most situations these fuels are introduced into the combustion

chamber in the form of a spray, the focus of initial studies was on the

evaporation and combustion of single, isolated multicomponent fuel drops. For

example, the early study of Wood et al.[l] identified batch distillation, which

735
is a process relying entirely on volatility differentials, as the main

controlling parameter for changing the internal composition of the drops. Landis

and Mills [2] were among the first to identify the internal diffusional

resistance as another possible controlling process for multicomponent fuel drop

evaporation. Their niimerical results show that at short time scales, batch

distillation models fail to predict the rate of evaporation because they neglect

thermal resistance, and for the bulk of the drop lifetime the distillation model

fails because it neglects species diffusional resistance. Landis and Mills [2]

acknowledge that they do not take into account the effect of circulation inside

the drop, which would have the effect of bringing the results of the

conduction/diffusion model closer to the batch distillation model. Subsequent

investigators [ 3 , 4 , 5

] who took into account convective transport through internal

circulation found this assertion correct. More precisely, it was found[5] that

internal circulation reduces the characteristic length of diffusion by about onethird, which has the effect of reducing the diffusion characteristic time by

about an order of magnitude. The excellent review of Law[6] discusses in detail

various situations and the relative importance of liquid phase mass diffusion

with respect to batch distillation effects.

The importance of liquid phase mass diffusion was also investigated by

Randolph et al. [7] whose experimental results revealed that the effectiveness of

mass diffusion decreases with increasing volatility differentials of the mixture;

the conclusion was that the gasification mechanism is intermediate between batch

distillation and a steady state controlled by diffusion.

The results obtained from the studies of isolated, single, multicomponent

fuel drops have been used directly in presently- existing calculations [ 8 , 9 , 10]

dealing with multicomponent fuel sprays, which have thus all ignored effects due

to drop interactions. The problem with this procedure is that in all these

calculations the dense spray effects observed near the injector have been

ignored. Existing evidence based upon numerical results obtained from

calculations performed for single component fuel drops shows that these effects

are important and do not have merely a quantitative effect because the

controlling parameters might be totally different for dense and dilute

collections of drops [ 11 , 12 ]

,

The focus of the present study is on the behavior of dense and dilute

clusters of drops composed of a binary fuel which has one component that is

considerably less volatile than the other. This conceptual model depicts the

situation of the alternate fuels discussed above. Under the assumption of one

fuel being much more volatile than the other, it will be shown that for dilute

clusters of drops, similar to the results obtained for single drops, evaporation

of the volatile is controlled by mass diffusion. However, in contrast to the

dilute and single drop case, evaporation of the volatile is controlled by surface

layer stripping of the less volatile component when the cluster of drops is

dense

.

II. MATHEMATICAL MODEL

The model describes clusters which are spherical and are composed of

identical spherical drops, uniformly distributed within the cluster. The solvent

is assumed to have a relatively low volatility with respect to the solute;

practically the solute's volatility is infinite with respect to the solvent.

The formulation assumes that the drop is composed of a liquid core and a

thin boundary layer at the drop surface as shown in Figure 1. In the liquid core

736
the motion of the liquid is described by a Hill vortex solution

Uj. = AR2 (1-rVR^) cos^ (1)

ug = - AR2 (l-2rVR^) sin^

and the mass fraction of the high volatility component is uniform. The boundary

layer is considered to be quasi -steady because it is assumed that the circulation

time of the Hill vortex is much smaller than the regression time of the drop.

A parameter, Be, representing the ratio of the mass regression rate to a

characteristic volatile diffusion rate appears naturally in the analysis as

follows

:

Be = - [R/Dn,u, tiR/dt, u, = AR2 (2)

Thus, when Be«l, diffusion into the boundary layer governs the rate of

species transfer from the liquid core to the drop surface with subsequent

transfer from the drop surface to the gas phase (evaporation) governed by the

Langmuir-Knudsen evaporation law. Note that the overall rate of solute

evaporation is governed by the slowest rate between the diffusion rate and the

evaporation rate because these are sequential processes. When Be»l , the

transfer of the high volatility fuel from the liquid core to the gas phase is no

longer controlled by diffusion into the boundary layer, but instead by boundary

layer stripping, that is by the regression rate of the drop.

Note additionally that transfer of volatile to the drop surface is

controlled by the fastest between the diffusion rate and the drop regression rate

because these are competitive processes. It should also be pointed that the

quasi-steady boundary layer assumption is consistent with Be«l, when diffusion

is important, and inconsistent with Be»l, when diffusion is not important.

As mentioned above, it is further assumed in this model that the volatile kinetic

evaporation rate is much larger than the sum of the rates of mass diffusion

through the drop boundary layer and of loss through surface regression. This

assumption is checked continuously in the calculation using a relative activity

coefficient of unity, thereby providing a conservative estimate of this ratio.

It turns out that this is a very good assumption throughout the calculations.

This assumption means that a batch distillation- like evaporation is totally

excluded by this formulation.

In order to calculate the circulation velocity of the Hill vortex

represented by u, one needs to relate it to the incoming flow velocity, Ug. This

is done by relating u, and Ug through

K3 - 2u,/(3ug) (3)

and finding the functional form for Kg. In the present model

K3 = F, (Re, ReJ/(3Mp) (4)

where Re = Ip^ Ug R//ig, Re^ = ^, R | dR/dt | //ig , (Re, Re^) = 1/16 Re Cf /(Re^) , and Cp = 2Cd / [3(1 + 0.0045 Re°-^^)] . Cq is calculated using the numerical results

of Cliffe and Lever [13] for drag on a drop with blowing effects from the drop.

The function / is found from a curve fit and behaves approximately as exp (-0.1

Re^) . This insures that F^, goes to previously found limits [ 14 , 15 ] when Re 0

or Re^ -> 0. Thus the equations describing the evolution of drop are:

737
1) Conservation of mass

5i = - 47rR2 p^ dR/dt (5)

The evaporation rate for the drop is the sum of the evaporation rate of the

solvent according to the Langmuir-Knudsen law, the diffusion rate of the volatile

through the boundary layer and the rate of boundary layer stripping loss of the

volatile

.

2) The Langmuir-Knudsen kinetic evaporation law for the solvent

msoLV = ^^R^ «e [WsOLv/(2?rRg T^s) ] [X'sqlV Psat.s (Tds)-X«S0LV P«.] (6)

3) Conservation of species

^soLv = 1 - Yjjv (7)

4) The average rate of volatile diffusion through the boundary layer is

calculated by integrating the diffusion velocity over the surface of the drop as

follows

:

uj, = 0.5 /ud sin^ de = (D^u,/R)i/2 Yhv,c /u (Be, (8)

o In Eq. (8), Up is found from the boundary layer solution and Wj. is a parameter

much smaller than unity which describes the diffusion of volatile from the drop

core into the return region of the boundary layer, located near the drop

centerline. The dependence of upon Wj. is very weak and thus is further

ignored.

5) The global conservation equation describing diffusion of volatile mass

out of the core into the boundary layer

dYHv.c/dt = - 3(D,u,/R)1/VR fu (Be) Y^^, (9)

6) The rate of boundary layer stripping is found by estimating the

thickness of the boundary using a subsurface streamline flow time. By definition

The mass carried away ^

by the evaporating ^ - 47rR2 dR/dt = 47rR2 (Dn,Uj/R)^/2 Yg Be (10)

solvent

where

- l/2Ar? / dr? / Yhv sin^ d6 , r, ^ (u, R/DJ^f^ (1 - r/R) (11)

-^o -^o /• u

The boundary layer thickness is estimated by At? = Be (j) - dt = Be r, and

R _ numerical calculations for subsurface streamlines give t~ 5. Thus Yg can be

written as Y3 = Yhv,c /y (Be).

Note then that the mass loss of volatile from the drop is

47rR2 p, (D„ Uj/R)i/2 y^^^ (gg) + j-^ (g^) j ^12)

738
where functions and fy are shown in Fig. 2 for small Be.

Thus the rate of mass loss of the volatile from the drop is parameterized

by Be, The time dependent parameter Be which changes the mass loss rate is

calculated at each time step as it depends upon the solution of the global

conservation equations for the entire cluster. These global conservation

equations are similar to those previously used for single components

fuels [11, 12]

.

7) Conservation of energy inside the drop. The effect of circulation on

heat transfer was investigated in detail by Abramzon and Sirignano[15] who

performed detailed as well as approximate calculations. The results of that

study show that circulation effects can be ignored if the conduction model for

heat transfer inside the liquid uses an effective heat transfer coefficient.

This effective liquid heat transfer coefficient is obtained by multiplying the

conductivity by a factor depending upon the drop Peclet number. Numerically, it

was shown [15] that this factor increases from 1 to 2.72 as the circulation

strength increases. The advantage of this approximate model is that the global

effect of circulation obtained with the more detailed calculations is

retained [15] while the calculations are less cumbersome. This approach was

adopted here as well and coupled to the previously developed model[ll-12 , 16]

.

Ill . RESULTS

Baseline results were obtained for a two component fuel where the solvent

is No 2-GT (light diesel) and the solute is n-decane. The calculations were

performed at atmospheric pressure with the following parameters: T°ga = lOOOK,

T°gs = 350K, Y°Fva = 0, = 3cm, R° = 2xlO'3cm, u°d = 200cm/sec , Y°hv,c = 2xl0"2.

Both in these calculations and the parametric studies discussed below, the gas

is initially at rest and thus vP^ represents also the initial value of the slip

velocity between drops and gas. Table I shows the values of the thermophysical

properties used in the calculations for all the fuels considered here.

Figure 3 depicts the evaporation time versus the initial air/fuel mass

ratio for two turbulence models. In the first model, called the weak turbulence

model [11], turbulence is not initially present in the surroundings of the

cluster, but rather develops with time as the cluster becomes less "porous" to

the ambient air[ll] . In contrast, in the second model, called the strong

turbulence model [11] , turbulence is initially present in the surroundings of the

cluster

.

Similar to the results of previous work [11] it is found here that the

evaporation time is insensitive to turbulence in the surroundings of the cluster

in the dilute cluster regime, and very sensitive to turbulence in the

surroundings of the cluster in the dense cluster regime. As can be seen in Fig.

3, with weak turbulence it is possible to obtain saturation before complete

evaporation in the dense regime at initial air/fuel mass ratios for which

complete evaporation occurs if the turbulence is strong. Another important

feature of these plots is the fact that they show a considerable reduction in the

evaporation time in the dense regime due to turbulence effects. This is due to

the turbulent transport of hot air from the surroundings to inside the cluster

thus replenishing the source of heat for the liquid phase and thereby promoting

evaporation. When turbulence is weak, the liquid phase depletes the heat content

of the gas phase without the possibility of hot air from the surroundings

penetrating the cluster.

Most important, the plots of Be versus the residual drop radius presented

739
in Fig. 4 show that mass diffusion is a controlling parameter for dilute sprays

(initial drop number density of the order of 10^ cm"^ or less for drops having

R° = 2xlO"'' cm) , but it is not controlling for dense clusters of drops where

instead volatile loss is dominated by surface layer stripping. Basically, in the

dense regime the slip velocity between drops and gas relaxes very fast because

the cluster exposes a large area to the gas [12] and circulation in the drop

ceases when the residual radius is still very large. In the absence of

circulation, mass diffusion is too slow to be controlling evaporation, and

evaporation proceeds at a basically frozen drop composition. This finding is in

total qualitative agreement with the experimental observation of Law and Law[17]

who minimized internal circulatory motion by conducting experiments with

suspended drops in low pressure environments so as to minimize buoyancy effects.

The fact that the evaporation of the volatile is controlled by surface

layer stripping is also clearly seen in Fig. 5 where Ygy^c plotted versus R^.

The curves show that after an initial decrease in Yyv this quantity stays

constant during the entire life time of the drops proving that the drops

evaporate at constant composition.

Since control of the evaporation by surface layer stripping occur in the

dense regime due to the fact that the slip velocity relaxes very fast with

respect to the drop lifetime, parametric studies were performed to investigate

the possibility that these two characteristic times might be of comparable value

under realistic practical conditions. Thus, either T° g was increased or was

decreased to reduce the evaporation time, or u°d was increased to augment the

relaxation time of the slip velocity. The value of T°ga was increased in steps

of 200°K from 1000°K to 1600°K and the same trends were observed. When the value

of u°jj was increased from 200 cm/sec to 1000 cm/sec the results show that dense

clusters of drops cannot sustain large slip velocities for any substantial amount

of time and th^trends of the baseline case were recovered again. This was also

the case when R° was decreased from 3 cm to 1 cm.

IV. CONCLUSIONS

A model has been presented for use in the description of the evaporation

of dense and dilute clusters of two - component fuel drops. The model assumes that

the volatility of the solute is much higher than that of the solvent.

Results obtained with this model show that, similar to previously

established results, the evaporation of the volatile in a dilute cluster is

partially controlled by liquid mass diffusion. In contrast, results obtained for

dense clusters of drops show that the evaporation of the volatile is controlled

by surface layer stripping of the solvent; this is an entirely new result which

is not totally surprising when one considers that the slip velocity between

phases relaxes extremely fast in the case of dense clusters [ 12 ] , and that it is

the drop circulation induced by this slip velocity which enhances liquid mass

diffusion which is otherwise an extremely slow process. This is also consistent

with the experimental evidence of Randolph et al.[7] for fuels of increasing

volatility differentials. In the present case, internal circulation is not

effective both because there is negligible slip velocity and because one of the

components is much more volatile than the other.

Parametric studies performed by varying the ambient temperature, the

initial slip velocity, and the initial radius of the cluster show that surface

layer stripping controls volatile evaporation for dense clusters of drops. In

fact, it was also shown that dense clusters of drops cannot sustain finite slip

740
velocities for any substantial amount of time.

AKNOWLEDGEMENTS

The research described in this paper was performed by the Jet Propulsion

Laboratory, California Institute of Technology, and was supported by the U.S.

Department of Energy, Office of Energy Utilization Research, Energy Conversion

and Utilization Technologies Program, Mr. Marvin Gunn, Jr., Program Manager,

through an interagency agreement with the National Aeronautics and Space

Administration.

REFERENCES

1. Wood, B. J., Wise, H. and Inami, S. H. , "Heterogenous Combustion of

Multicomponent Fuels", Combustion and Flame . 4, 235-242, 1960.

2. Landis, R. B. and Mills, A. F. , "Effect of Internal Diffusional Resistance

on the Evaporation of Binary Droplets", Paper B7.9, The 5th International

Heat Transfer Conference, 4, 345-349, 1974.

3. Law, C. K. , Prakash, S. and Sirignano, W. A., "Theory of Convective,

Transient, Multicomponent Droplet Vaporization", 16th Symp. (Int.) on

Combustion . 605-617, 1977.

4. Sirignano, W. A. and Law, C. K. , "Transient Heating and Liquid-Phase Mass

Diffusion in Fuel Droplet Vaporization", Advances in Chemistry Series 166.

Zung, J. T. (Editor), 3-26, 1978.

5. Lara-Urbaneja, P., and Sirignano, W. A., "Theory of Transient

Multicomponent Droplet Vaporization in a Convective Field" , 18th Symp.

(Int.) on Combustion . 1365-1374, 1981.

6. Law, C. K. , "Recent Advances in Multicomponent and Propellant Droplet

Vaporization and Combustion", Paper 86-WA/HT-14 presented at the Winter

Annual Meeting, ASME, 1986.

7. Randolph, A. L. , Makino, A. and Law, C. K. , "Liquid-Phase Diffusional

Resistance in Multicomponent Droplet Gasification", 21st Symp. (Int.) on

Combustion . 601-608, 1986.

8. Continillo, G. , and Sirignano, W. A. , "A Numerical Study of Multicomponent

Fuel Spray Flame Propagation in a Spherical Closed Volume", 21st Symp.

(Int.) on Combustion . 1941-1949, 1988.

9. Aggarwal, S. K. , "Ignition Behavior of a Multicomponent Fuel Spray",

Combustion and Flame . 76, 5-15, 1989.

10. Raju, M. S. and Sirignano, W. A., "Multicomponent Spray Calculations in a

Modified Centerbody Combustor" , Paper AIAA-88-0638 presented at the 26th

Aerospace Sciences Meeting, January 11-14, Reno, Nevada, 1988.

11. Bellan, J., and Harstad, K. , "Turbulence Effects During Evaporation of

Drops in Clusters", Int. J. Heat Mass Transfer . 31, 8, 1655-1668, 1988.

12. Bellan, J. and Harstad, K. , "The Details of the Convective Evaporation of

Dense and Dilute Clusters of Drops", Int. J. Heat Mass Transfer . 30

.

6,

1083-1093, 1987.

13. Cliffe, K. A. and Lever, D. A., "Isothermal Flow Past a Blowing Sphere",

Int. J. Numer. Meth. Fluids . 5_,. 709-725, 1985.

14. Sadhal, S. S. and Ayyaswamy, P. S., "Flow Past a Liquid Drop with a Large

non-Uniform Radial Velocity, J. Fluid Mech. 133, 65-81, 1983.

15. Abramzon, B. and Sirignano, W. A., "Droplet Vaporization Model for Spray

Combustion Calculations" Paper AIAA-88-0636 presented at the 26th

741
Aerospace Sciences Meeting, January 11-16, Reno, Nevada, 1988.

16. Bellan, J. and Harstad, K. , "Evaporation, Ignition and Combustion of

Nondilute Clusters of Drops", Combustion and Flame . 19_, 272-286, 1990.

17. Law, C. K. and Law, H. K. , "A d^-Law for Multicomponent Droplet

Vaporization and Combustion", AIAA Journal . 20, 4, 522-527, 1982.

NOMENCLATURE GREEK SYMBOLS

A vortex strength M viscosity

Be defined by Eq. (2) <t> air/fuel mass

ratio

Cd drag coefficient P density

Cp friction coefficient ri boundary layer

heat capacity at constant variable

ir pressure for the gas T nondimens ional

C| heat capacity at constant

pressure for the liquid

e angular coordinate

within the drop

Dm coefficient of mass diffusion

in the liquid SUPERSCRIPTS

function of Re and Re^

k thermal conductivity o initial condition

defined by Eq. (4)

L latent heat of vaporization SUBSCRIPTS

•m evaporation rate

P pressure a ambient

Pr Prandtl number b back

r radial coordinate within the drop boil - boiling point

R

^

radius of the drop c core

radius of the cluster d drop

Re Reynolds number f forward stagnation

Re.^ evaporation Reynolds number point

universal gas constant g gas

Ri R/R° HV high volatility

T temperature i liquid

t time s drop surface

diffusion velocity sat saturation

Uj velocity of the vortex motion SOLV - solvent

inside the drop St stoichiometric

radial velocity 00 in the air field

velocity in 6 direction

Wr parameter much smaller than unity

W molecular weight

X mole fraction

Y mass fraction

742
TABVEI THgHMOPHYSICAL PROPERTIES OF FUELS

w (g/molB)

C. (cal/g K)

C, (cal/g h)

K. (cal/g)

T»,(K)

k,(cal/cm s K)

/>e (g/cm')

(i« (g/cm s)

D„ (cm'/s)

n-decan«

142

0.400

0.523

73.92

4477

3.25x10-

0.734

9.2x10^

4.4x10'(est)

14.97

14.4

N0.2-QT

(Ught Diesal)

HV.f Fig. 1 Pictorial of the circulatory

motion Inside half of the drop and of the

system of coordinates within the drop.

HV.b

743
10.00

800

600

•evapx 10 .sec

002

>

I

>

001

400

2.00

000

1.E 01

1 WEAK TURBULENCE

• STRONG TURBULENCE

O SATURATION BEFORE COMPLETE

EVAPORATION

-L

1 E+00 1 E+01 IE +02

Fig. 3 Variation of the evaporation time

with the initial air/fuel mass ratio; T°g,

lOOOK, T°^, - 350K, Vgv.c " 0 02, u°d - 200cm/sec # - 3cm, R° - 2xlO-3cm, YV.-O

Variation of Be with the residual

drop radius for several initial air/fuel

mass ratios for the strong turbulence case;

® 0250

0 0314

A 0400

0 785

1.18 O 1.57 > 3.14

<] 785

• 15 7 314

lOOOK,

200cm/sec, R*" - 3cm

350K, Y°,HV,c - 0.02,

R° - 2xlO-3cm, Y°f

u d

Fig. 5 Variation of the liquid mass fraction of the volatile with the residual

drop radius for several initial air/fuel

mass ratios for the stong turbulence case;

025

I 0.50

T°g, - lOOOK. - 200cm/sec,

350K, YV. 0.02. u°H 3cm, R° - 2xl0"'cm, Y°fv,-0

075 1.0

744
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 83

MODELLING OF ELECTRO-AEROSOL DEPOSITION IN

HUMAN AIRWAYS

W. Balachandran, C.N. Ahmad and S.A. Barton

Department of Electronic and Electrical Engineering

University of Surrey

Guilford, United Kingdom

ABSTRACT

A mathematical model has been developed for the deposition of electroaerosols in the human respiratory system. In this model an integrated approach

is used which takes into account inertial impaction, gravitational

sedimentation, Brownian diffusion, interception and electrostatic forces. The

model has been realised using a personal computer. This model is capable of

handling both homogeneous and inhomogeneous distributions of drop size and

charge .

INTRODUCTION

In aerosol therapy the primary aim is to deliver the drug directly to the

required site of action, and therefore the therapeutic effect is expected to

be more rapid than that of the oral dose. The devices used at present for drug

delivery into the human airways include nebulisers, metered-dose inhalers

(MDI) and dry particle inhalers (DPI) . The basic mechanisms involved in the

deposition at specific sites are the same in all three cases, except the

overall efficiency will depend on the nature of drug aerosol generated by the

respective devices. The deposition of inhaled drug aerosol in the respiratory

system is governed by inertial impaction, gravitational sedimentation,

Brownian diffusion, interception and electrostatic forces.

Yu (1) and Hashish (2) have made a significant contribution in terms of

theoretical studies on various lung deposition mechanisms of inhaled drug

aerosols. Both groups have shown that deposition is dominated by electrostatic

forces particularly in the alveolar region. Various simplifying assumptions

were made in developing these models. It is desirable to re-examine these

basic assumptions with a view to improving the model and achieving more

realistic and better estimates of deposition efficiencies. We have developed

an integrated deposition model for inhalation only, and tested it for

consistency with work done by other researchers. The model assumes the airway

structure to be tubular and does not include alveolar sacs. In this paper, the

theoretical background and the evaluation of the model for various particle

size ranges and charge levels is reported.

There are four main elements of the model - an anatomical model of the

lung, a model of the act of breathing, a description of the aerosol and

models of the mechanisms of deposition of aerosol particles. None of the

745
elements give a strictly true picture of the processes and structures

represented. Indeed a true picture would not be desirable as it would be

analytically incalculable. In each case an approximate model has been used,

and in each case there are known means by which to refine the model

should the extra complexity be justified by extra accuracy of descriptive or

predictive performance.

The anatomical model of the lung used is based on Weibel's symmetrical

model A (4) . The morphometric data for the human lung, cited by Weibel and Yu

has been modified in the light of more recent published data of Horsfield

et.al.(5). Each airway generation is represented by a cylinder whose total

length increases and diameter decreases with penetration into the lung. When

using the Weibel lung model it is a natural step to assume that during an

inspiration air flow divides symmetrically at branching points. We further

assume that the flow rate during inspiration is constant, as are the

conditions of the inspired aerosol. The properties of the aerosol used in this

model can be either homogeneous or inhomogeneous , in terms of drop size and

charge per drop, whereas the other parameters such as density and numerical

concentration remain the same.

DEPOSITION EFFICIENCIES

We define deposition efficiencies T| here in a simplified manner, valid as long

as the total deposition efficiency is small, as the number of droplets

deposited in a section of airway divided by the number entering the section.

The expressions for the efficiencies due to different deposition

mechanisms have been derived by various authors. For an airway distinguished

here by its generation j the deposition efficiency components are as follows :-

Inert lal Impaction; T|j

j

An empirical relation derived from the experimental data of Schlesinger et . al

(6) and given by Yu (1) yields:

T|,. = 0.768* (St )Q

(1)

where St is the Stanton number and 9j the assumed branching angle given by

where

S =

q)id

pdpU L _ and e = —

. 3 4d.

p= droplet mass density, dp = droplet diameter,

Lj = length of the airway (jth generation), Uj = air velocity in the

airway, dj = diameter of the airway, |J. = air viscosity

Gravitational Sedimentation ; Tjs j

Pich (7), for a parabolic flow, give the expression;

where

21 7 -]

2e - e' + sin ^ V 3 ij [^^ (2)

Ug = droplet settling velocity

746
The angle of inclination of the airway to the vertical, 0 is assumed to be

given by sin <j) = 7C/4 which corresponds to the case of a system of randomly

oriented airways, Heyder (8)

.

Diffusion . Two kinds of diffusion processes are considered in this study:

i) The expression for deposition efficiency due to diffusion in a

turbulent flow is given empirically by Landahl (9):

if i

\' U- 0.44A' +

ii) For .laminar flow , Ingham (10) gives as an empirical result

Tl = 4A'

'Dt3 (3)

T|jj^.= 1- 0.819exp(- 14.63A) - 0.0976exp(- 89.22A)

- 0.0325exp(- 228A) - 0.0509exp\,- 125. 9A'^

(4)

where in each case i and ii:

DL . A= '

2 d.U.

in which: D = The Brownian diffusion coefficient of an aerosol particle.

In the upper respiratory tract the Reynolds number is high. This together with

the non-straight nature of the URT suggests that turbulent flow will prevail

high up in the URT. However, as one penetrates deeper into the lung Reynolds

number falls very rapidly, so inertial effects in the air flow become

negligible and the flow will within any generation rapidly become established

as fully-developed laminar flow. For this paper we have used to calculate the

composite deposition efficiency a mean of the two values down to 13th

generation beyond which we have used the expression for laminar flow.

Electrostatic Depoaitlon ; 'Hspj ^Imj

Two effects of electrostatic forces are considered in this study. These are

the effects of forces on charged particles due to a space charge and those due

to the perceived image of a charge in the proximity of a conducting surface.

(i) The space charge force is due to the mutual repulsion of aerosol

particles of like charge. The deposition efficiency due to the space charge,

was derived by Yu (1977) and is given by:

CqR't. Bq^L.

Tl . = — where X . =

'Spj 3 j 3 H-C„R.X. 47tE„R.U. Odd 0 ] 3

in which C = concentration of particles at the entrance of the airway. o J , radius of the airway (jth generation), Xj = a dimensionless residence time.

|i. = particle mobility, q = charge per particle Eq = permitivity of air.

(ii) The image force is due to the interaction of charged particles with

the conducting respiratory walls. For a unipolar aerosol, under slug flow

747
conditions, Yu (11) derived analytically the following expression:

T1, .= 1-

radius vector' is related to T as :-

rr,\

R where rc , the 'critical

T= 4

C

\ r + 21n c R . R .

*^ 3 ) V (6)

The above set of equations is algebraically convoluted, to an extent not

necessarily justified by the purpose of the model or by the resolution of any

conceivable technique for testing or applying the model. A result of the

algebraic complexity is that it is impossible with the equations in that form

to have a feel of what the equations imply in a general manner. Any wellbehaved (smooth, continuous) mathematical function can be expressed as an

infinite polynomial, or approximated by a finite polynomial. Such

approximations to the above equations were derived, using series expansions

with just sufficient terms to leave in each equation a single (if composite)

variable term. We then checked these simplified equations for accuracy within

the range of deposition efficiencies suggested by the model using the full

equations

.

General Global Generation

constant. parameter. parameters.

11 ij = {l.5*10^}

4 N.d . 3 3 J -Impaction

,

sj = {3*10 }

I Q J Sedimentation.

ll,,j«{l.9*10"'} -Turbulent diffusion,

Tl^^.= {5.18*10

-- Laminar diffusion.

11 . « {l2*10 S -Tl .d^N .C„.l

L ] : D o:j Space charge.

n. .= {3.9*10 / Qd

pj

N L

j j

P J Image force.

number of airways in the jth generation ,Q = air flow rate

748
Logarithmic differentiation of the single term gives an instant indication of

the sensitivity of the solution either to variations of a controlled parameter

(such as the size of the aerosol particles or the inspiratory flow rate) or to

errors in assumed parameter values. Collecting global constants, general

constants and global variables together, the resulting simplified equations

were as listed above.

REALISATION OF THE MODEL AND RESULTS

The model was evaluated for both homogeneous and inhomogeneous distributions

of droplet size and charge. Main features of the results are that the

deposition efficiency due to impaction is extremely small throughout the lung

for the drop size (0.5-7.0 jXm) and drop charge levels (5.0 -300 charges per

particle) used in this study. The deposition efficiencies due to sedimentation

and diffusion, though generally higher than impaction, are still small

compared to those due to electrostatic forces which in fact dominate the

deposition throughout the lung. The total deposition, at this stage, is

considered to be a simple addition of the individual deposition

ef f iciencies . Fig . 1 indicates the effect of drop size on the amount of the

aerosol deposited in different parts of the lung. The area under the curve

between any two generations give the proportion of the original aerosol

deposited between those two generations. This figure also indicates that,

under the assumed parameters values, there is almost a constant deposition of

4 % throughout the lung. The figures within the square brackets represents

the proportion of the aerosol which is left undeposited after 23rd

generation.lt is assumed that this amount, because of their small droplet

size, high charge and laminar flow conditions, would be deposited in the

alveolar sacs. As the droplet size is increased to 3 }im the deposition in

Table 1

H o m o g (3 n e o u s Mod e 1

Curve a dp 1 1 1 1 1 1

q 100 100 100 100 100 100

I n •- h o m o g e n e o u s Mode 1

Proportion % 2 15 34 29 12 8

Curve b

dp 0.2 0.5 0.8 1.0 1.2 1.4

q 100 100 100 100 100 100

Proportion % 2 15 34 29 12 8

dp 0.2 0.5 0.8 1 . 0 1.2 1.4

Curve c

q 20 30 40 60 80 100

airways ( 0-12 generations) decreases due to the decreased charge

coarser droplets. This gives a peak deposition around 20th

and minimal amount (1.7 %) penetrates the alveolar sacs. With 7 ^m

the upper

effect on

generation

749
droplets , due to increased impaction and sedimentation, the deposition in the

upper airways is increased considerably The deposition peak also shifted from

20th to 14th generation with the result that nothing is left to be deposited

in the alveolar sacs. Fig. (2) shows the effect of droplet charge on

deposition. A small proportion of droplets is deposited up to 10th generation

when charge level of 5 electronic charges per droplet were considered. The

deposition increases slowly with the penetration into the lung attaining a

peak deposition at 23rd generation and about 50% of the aerosol enters the

alveolar sacs. As the charge is increased to 100 electronic charges per

droplet the deposition throughout the lung is increased resulting in a peak

deposition around 20th generation. When the charge level is further increased

to 3 00 charges per droplet the situation is entirely changed. The deposition

decreases exponentially from a maximum (13%) at zero generation to almost zero

at 23rd generation with the consequence that only 0.4% is left for deposition

in the alveolar sacs.

Fig 1. Drop size effect on deposition (homogeneous)

AMOUNT

DEPOSITED (%)

0 5 10 15 20 25

GENERATION NUMBER

All the results uptill now were based on the assumption that the aerosol

consists of collection of monosize droplets having equal amount of charge. In

practice this is hardly the case and the aerosol not only has a narrow size

distribution but also has a specific charge distribution. The results of the

model, when these distributions were considered, are shown in fig. 3. Curve

(a) indicates the deposition pattern in the lung when all the droplets of the

aerosol are considered to be of equal size and carrying same charge

(homogeneous case) as in fig.l. Curves (b) and (c) are the results of the

solution of the inhomogeneous model under an arbitrary set of parameters shown

in table 1

.

Curve (b) shows the effect of assumed narrow size distribution of inhaled

droplets, each carrying 100 electronic charges. Since the size distribution

taken for this model evaluation has a VMD of Ifim and the spread is narrow, the

effect is not significant. However when both inhomogeneous distribution of

size and corresponding charge distribution are used the model predicts a

significant deviation in comparison with the homogeneous model. In this case

the upper airway deposition decreases to about 1% and after 10th generation,

the deposition efficiency increases significantly attaining a peak value

750
around 22nd generation. Although the amount entering into the alveolar sacs

increases to 21%, it is believed that due to laminar flow conditions in the

alveolar sacs , the image charge force will be strong to deposit almost all of

these small droplets carrying high specific charge.

Fig 2. The effect of droplet charge on deposition

AMOUNT

DEPOSITED (%)

0 5 10 15 20 25

GENERATION NUMBER

CONCLUSIONS

.

A flexible integrated lung deposition model incorporating the effect of all

five different mechanisms has been developed and realised. The analysis of the

model indicates the approximate sensitivity of each deposition efficiency to

variations in each parameter. With control over appropriate parameters, it is

possible to predict or determine the primary target of an inhaled aerosol in

the bronchial tree.

751
ACKNOWLEDGEMENT

•The authors wish to acknowledge BESPAK Pic. for the financial support provided

for the study

.

REFERENCES :

1. Yu CP. and Diu C.K, "A comparative study of aerosol deposition different

lung models", Am. Ind. Hyg. Assoc. J., 43, pp. 54 - 65 (1982).

2. Hashish A.," The influence of electrostatic charge on the deposition of

therapeutic aerosols." PhD thesis University of Southampton UK (1988).

3. Melandri,C., Tarroni G., Prodi, V., De Zalacomo T., Formignani M., and

Lombardi C.C." Deposition of charged particles in the human airways",

J. Aerosol Sci., 14, pp. 657 - 669, (1983).

4. Weibel E.R., " Morphometry of the human lung ". Academic Press New York , pp.136 -140, (1963)

.

5. Horsfield, K. et . al .

, " Models of human bronchial tree". Journal of

applied physiology, vol.31 No. 2, pp. 207 - 217, (1971).

6. Schlesinger R.B., Bohning D.E., Chan T.L., and Lippmann M., "Particle

deposition in a hollow cast of the human tracheobronchial tree" J. Aerosol

Sci., 8 , pp. 429 - 445 (1977).

7. Pich J., " Theory of gravitational deposition of particles from laminar

flows in channels", J. Aerosol Sci., 3, pp. 351 - 361, (1972).

8 . Heyder J., " Gravitational deposition of aerosol particles within a

system of randomly oriented tubes", J. Aerosol Sci. 6, pp. 133 - 137,

(1975) .

9. Landahl H. D., "Particle removal by the respiratory system" Notes on the

removal of airborne particulates by the human respiratory tract with

reference to the role of diffusion'. Bull. Math . Biophys , 25 pp. 29 - 39.

10. Ingham D. B ., "Diffusion of aerosols from a stream flowing through a

cylindrical tube", J. Aerosol Sci. 6, pp. 125-132,(1975).

11. Yu CP.," Precipitation of unipolar charged particles in cylindrical &

spherical vessels", J. Aerosol Sci., 8 , pp. 237 - 241,(1977).

752
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 84

DROPLET INTERACTION IN SPRAY MODELLING

M.M. Elkotb*, O.M.F. Elbahar', T.I. Sabry^ and S.A. Wilson^

*Cairo University, Cairo, Egypt

^Menoufia University, Egypt

ABSTRACT

A mathematical model with more advanced concept about the spray droplets interaction with the

surrounding air is suggested in this work. Using the model, a parametric study has been carried out to analyse the effect of different operating conditions on the spray behaviour. A comparison of the

theoretical results with available experimental data shows a fair agreement.

INTRODUCTION

Reviewing the theoretical models in literature, one can classify them to three different categories.

The first category [ 1 - 5 ] is mainly based on correlations of experimental results to obtain the spray

penetration length, and spray tip velocity. Such technique cannot give any information about the spray

droplets and its behaviour and the interaction between spray droplets and the gas phase. The second is based on continuum mechanics [6,7,8]. These models enable the calculation of spray penetration , mean velocity and mean concentration distribution. The entrained air and the interaction between the air

field and the droplets are not considered. The lateral velocity, droplet size distribution and consequently

the local concentration cannot be calculated with this technique. The third category of spray modeling is based on discrete droplet model in which the spray droplets are divided into groups of certain size and

number [9-14]. The classical differential equations of motion and energy are applied on each group to define the velocity, temperature, size and the location of each one. Various modifications have been suggested such as variation of drag force and heat transfer with

Reynolds number, droplet size distribution and stochastic behaviour of droplets. However, complete

interaction between the two phases has not been satisfied. A trial has been performed in the present

work to inrtoduce several modificafions for the interaction between the two phases to improve the

accuracy of the discrete droplet model.

MATHEMATICAL MODEL

In the present work the conservation equations of momentum, continuity and energy in 3D form

have been used together with the two equations simulating the turbulence in the gas phase 'k-e' . The

constants of the turbulence model have been modified to fit the transient flow inside the cylindrical

combustion chamber of a Diesel engine. The conservation equations representing the gas phase flow are Eularian and solved using the iteratively SIMPLEC technique, while the differential equations of

droplets morion, which are in Lagrangian form, are solved for each group of droplet to obtain the

droplets trajectory, velocity, size and temperature.

An interaction mechanism between the gas phase and the spray droplets are proposed as follows:

- Momentum transfer from the droplets to the gas phase and mutual interaction of droplets and its effect

on drag force.

753
- Energy transfer to the gas phase due to fuel evaporation.

- Random motion of droplets and its effect on gas phase turbulence represented by the Pseudoturbulence equation.

Governing Equations And Interaction Scheme

The general form of the conservation equations and the two equations of turbulence can be written

with an arbitrary variable as foUows:

d/dz (p(!)) + a/ax [ pv(t> - (d^/dx)] + i/r a/ar [rpv(!) - r^j, r (a(j)/ar)] + i/fi a/ae [rpwcj) - di^/dQ)]

= + S(j) f (1)

Where Values of S(|) , ^ , and turbulence model constants are given in [20]. Shear stress,

generated from the turbulence motion of droplets, appears in the source term of the momentum equation

of gas phase as follows.

'^ij ps = [( ^i / ^^^j ) + ( 3uj / axi )] (2)

Where iXpg = 1.3 p^a R ( v/-- Ug ) (3)

The Lagrangian conservation equations of the droplets can be written as follows:

dm^/di =-CDA^PgV2rel/2 (4)

m^ Cp[ (dT^ /dT ) = H (Ta - T^ ) + L ( dm^ / dx ) (5)

- ( dm/-/dT ) = 4 7C R- (Sh Dc /2R) ( Ps- p^o ) = 4 ;c r2 p^ (dRf/dx ) (6)

diffusion coefficient,

Dc = Do(Ts/ro) (P/Pa)

Sherwood number,

Sh = 2 + 0.6 Rel^ Scl/3 [16]

And consequently the rate of change of droplet diameter becomes as follows.

dR /dT = - (Sh Dc /2Rpc) (Pg /RyTs) (7)

Where,

Ps = 2.39 X 109 e

- 6065 /Ts

Ts = 462 PaO.043 + 0.06 (Ta - 673) , 350<Ta<600 [14]

and the rate of temperature increase : dT /dx = (3 /R Cp^ [ h /p (Ta - T^) - (L Sh /2Rp^) (Pg /Ry Ts) (8)

Where h can be calculated from,

Nu = B 0.54 Re for 100 < Re < 1000 [15,17]

Nu = B (2+0.6 Re Pr ) for Re < 100

Where B is a correction factor to consider the effect of mass transfer;

6 = Z/Exp(z)-1 [18]

and,

Z = - [Cpv (dm^ /dx )] / 7C d K Nu

754
The modified drag coefficient used in the momentum equation is that proposed by ELK0TB[19] in

the following form.

Cpg(T^-Ts) + (QAF/AFst)

Cd = ( 28/ReO-79 ) [ 1/ p ( 2.5 / ReO-233 )] (9)

L

The spray is considered initially formed of droplets distributed in size groups according to the

following distribution function[14]:

1/nt [dn /d (D /D32 )] = 10 (D ID'^ip--^ e-2-4(D/D32) (10)

The Sauter mean diameter can be calculated from the following formula[14]:

D32 /dn =107 Re- 0.183 We - 0.442 c^- 0.422

(p^ /p^)- 0.05 (1 1)

Each group is treated as a single droplet according to the discrete droplet theory.

The traditional approach mentioned before, to modeling droplet dispersion in turbulent flow is based on that the diffusional mass flux is proportional to diffusion coefficient and concentration

gradient. Such approach is not sufficient to predict droplet motion in case unequal turbulent velocity of

air and droplets especially in a highly turbulent swirling air. In this case the air velocity fluctuation,

which has a random behaviour, affects the droplet velocity through a fluctuating drag force. Thus, the

air field transfers a random velocity component to the droplet motion, and a random droplet dispersion

is obtained. The instantaneous turbulent velocity of air field is represented with a random number

generator as follows:

u =

J

2 In (1/Rn) ^(2/3)K (12)

Where, Rn = random number ( 0.001 to 1

)

Solution Procedure

The two systems of equations ( the gas phase and liquid phase ) are solved together numerically in

iterative procedure method. The SIMPLEC algorithm is used to solve the finite difference form of the

gas phase governing equations. The combustion chamber is divided into four cells u,v,w and scalar

quantity cells covering the domain under consideration. The values of the dependent variables are

calculated at the pole of each cell. The grid used is represented in ref [1].

The diffusional equations representing the droplets behaviour are solved using RK-4 steps

technique taking into consideration the strong coupling between the two phases which is the interaction

mechanism.

RESULTS AND DISCUSSION

The theoretical model was used to determine the trajectory of the droplets and its velocity as well as penetration, fuel concentration and spray geometry. The following operating conditions were used.

- Fuel type Light diesel fuel

- Fuel density 850 Kg/m3

- Specific heat 2.299 Kj/KgK

- Cylinder bore 87.5 mm.

- Nozzle diameter 0.33 mm.

- Pump speed 750 rpm.

The instantaneous and rate of injected fuel during the injection period were calculated using the

following correlationsLl8].

755
Q* = 0.9 0* + 1.3 0*2-1.2 0*

dQ*/dx = 0.9 + 2.6 0* - 3.6 0*2

*3

(13)

(14)

The generation and development of spray droplets and air field with time, according to the spray

penetration had been calculated and the vector diagram of the gas phase and its variation with time are

illustrated in Fig(l&2). From these figures it is clear that, the generated air velocity takes its maximum

value along the spray center line. The higest value of the entrained air velocity in the spray zone lies near the injector and reaches its maximum value at the spray tip. This may be attributed to the small

momentum exchange. It is noticed also that the entrained air is directed to the lower spray zone and

outside in the upper part of the spray. Therefore, dispersion of droplet increases in this zone. Regarding

the injection pressure, the droplet velocity increases with its increase due to the increase of momentum

exchange. Thus the dispersion of droplets increases with the increasing of the injecfion pressure,

causing dilution of the fuel concentration along the spray zone especially at the spray tip.

Parametric study had been carried with different drag force correlations to validate the model

against the published experimental work[l]. The resulting turbulence from the spray droplets has been

included also to define its effect on the accuracy of results. Figure(3) illustrates the effect of drag force

represented by different correlations on the fuel concentration decay along the spray center line at stagnant medium condition. The results show that the concentration decay in the case of Ingebo

correlation is less in magnitude than that in the case of modified drag correlaUon. The reason may be

attributed to the small drag force value calculated by Ingebo drag correlation compared with that

obtained by modified drag correlation. Introducing the Pseudo-turbulence the decay of the fuel

concentration decreases. This may be attributed to the increase in gas phase turbulence level represented

by the turbulence kinetic energy K , which increases the fluctuation gas velocity component u . This

leads to the increase of the random droplet's motion and consequently the droplet dispersion and

decrease of the fuel concentration.

The same conclusion appears in Fig(4) which illustrates the effect of various drag correlafions on

the fuel concentration decay along the spray in the case of swirling flow.

The effect of various factors on the spray formation and penetration as well as the air field are represented in Figs(5&6). From these figures it is clear that a great deformation of the air field has been

resulted in the spray zone especialy behined the injector where the momentum transfer is maximum.

The effect of gas temperature on the spray behaviour is studied also. Figures(7&8) represents the

radial and axial concentration decay with increasing the gas temperature. This may be attributed to the

increase of mass transfer rate by increasing the gas temperature.

CONCLUSION

The proposed mathematical model in the present work has more advanced mathematical approach

to achieve better simulation for the spray droplets behaviour. The effect of Pseudo-turbulence generated

from the spray droplets mofion on the gas phase is considered. The results show that this factor

decreases the fuel concentration along the spray center line due to the resulting droplets dispersion.

Another important parameter which is the drag force was taken into consideration, so a modified drag

correlations [19] as well as other different correlations are tested to clear its effect on the spray droplets.

The results show that the use of modified drag correlation in the mathematical model gives a best

agreement with experimental results. This can be attributed to the realestic conditions used in deriving

this correlation. The effect of gas phase turbulence on the droplets motion is considered. This gives the

spray droplets their random motion. So statistical approach is used to simulate the turbulence nature

based on the turbulence kinetic energy of the gas phase, probability distribution function and random

number generator.

Parametric study has been carried out to clear the effect of spray behaviour in side the open

combustion chamber of Diesel engines . The spray inside a stagnant medium as well as the swirl air

were studied. The effect of fuel evaporation process was considered also.

756
REFERENCES

1 . SAAD M.A., "Spray Formation and t^ntrainment Characteristics in Liquid Fuel Sprays of Diesel

Engines. M.Sc. Thesis, Supervised by Prof. M.M. ELKOTB, Cairo University, (1985).

2. DENT, J.C.,"A Basis For The Comparison of Various Experimental Methods For

Studying Spray Penetration." SAE paper No: 716571 , (1971).

3. HAY, N. and JONES, P.L., "Comparison of Various Correlations For Spray Penetration" SAE

paper No: 720776 , (1972).

4. SHARED, S.M., CHIV, W.S. and LYN, W.T."A Transient Spray Mixing Model For Diesel

Combustion." SAE . Trans. Vol. 85 , pp. 502-512, (1976).

5. HIROYASU, H. and KADOTA, T., " Models For Combustion AndFormation Of Nitric

Oxide And Soot In Direct Injection Diesel Engines." SAE paper No:760129 , (1976).

6. ADLER, D. and LYN, W.T." The Steady Evaporation And Mixing Of A Spray In Gaseous

SwirL" Int. J. Heat Mass Transfer, Vol (14), pp. 793, (1971).

7. WAKURI, Y. FUJII, M. AMITANI, T. and TSUNEYA, R. " Studies On The Penetration Of

Fuel Spray In Diesel Engines." JSME. Bull., pp. 123-130 , (1960).

8. CATTON, I. , HILL, D.E. and McRAE, R.P., " Study Of Liquid Jet" AIAA Vol. 6, pp.

2084-2089, (1968).

9. WESTBROOK, C.K., "Three Dimensional Numerical Modeling Of Liquid Fuel Sprays."

California University, Livermore, Cal.94550.

10. CLIFFE, et. al,"A Finite Difference Calculation Of Spray Combustion In Turbulent Swirling

How." Int. Conf. on Numerical Methods. SWANSEA. U.K. (1979).

11. BRACCO, et. al.," Theoretical Analysis of Stratified Two-Phase Wankel Engines Combustion."

AIAA Vol. 16 , No. 10 , pp. 1053-1061, (1978).

12. ELKOTB, M.M. and RAFAT,N.M.,"Fuel Spray Trajectory In Diesel Engines." ASME, J. of

Engr. for power, No.77 DGP-1, Oct., (1977).

13. ELKOTB, M.M., ELSABILIGI, M. and DIAB, R.," Spray Behaviour Inside a Swirl

Chamber Of A Diesel Engine." Proc. of the 1st. Conf. of Mech. Power Eng., Cairo.,(1977).

14. ELKOTB,M.M., "Fuel Atomization For Spray Modeling." Progress in Energy and Combustion

Science, vol.8,(1982).

15. RANZ, W.E. and MARSHALL, W.R.,"Evaporation From Drops." Chem. Eng. Prog. 48,

141, Part I & II, (1952).

16. BIRDS, R.B. , STEWART, E.W. and LIGHTFOOT, E.N.

" Transport Phenomena. " John

Wiley, New York, (1960).

17. VIROPOFF, D.N., "Evaporation Of Fuel Drops." Bull, of the central Scientific Aviation

Institute, (1939).

18. ELWAKIL, M.M., UYEHARA, D.A. and MYERS, P.S. A, "Theoretical Investigation Of The

Heating Up Period Of Injected Fuel Droplets Vaporizing In Air." NACA T.N. 3179, (1954).

19. ELSALMAWY H.A. "Evaporation And Drag Characteristics Of Multicomponent Hydrocarbon

Droplets." M.Sc Thesis , Supervised by Prof. M.M.ELKOTB, Cairo University, (1986).

20. WILSON S.A. " Investigation of Spray Behaviour in Diesel Engiues". Ph. D. Thesis,

Supervised by Prof. M.M. ELKOTB, Menoufia University, Egypt, (1990).

757
NOMENCLATURE

A Cross section area, . AF Air to fiiel mass ratio.

AFg I Stochiometric air to fuel ratio.

C^ Coefficient of discharge.

Cp Specific heat at constant pressure, j / kg K.

D Droplet diameter, m. Dg 2 Sauter mean diameter, m. H Coefficient of heat transfer, w / m^ hr k'

L Latent heat of fuel, j / kg

Nu Nusselt number.

n Number of droplets in each group.

Total number of fuel droplets injected during certain period.

P Pressure, N/m^.

Pr Prandtl number.

Q Heat of combustion per unit mass of fuel reacted, j / kg.

Q* Nonnalized comulative quantity of injected fuel.

R Droplet radius, m.

Rn Random number ( 0 - 1). Rf Richardson number.

Ry Vapor gas constant.

Sc Schmidt number.

Sh Sherwood number.

Sr Swirl ratio, co/cag.

Source term of the conservation equations.

T Temperature, K.

U Velocity component in angular direction, m / s. V Velocity component in radial direction, m / s. W Velocity component in axial direction, m / s. We Weber number.

a Void fraction.

Effective diffusivity coefficient for variable ^ .

9 Augular distance coordinate, degree.

K Turbulence kinetic energy, m^/s^.

|J. Viscosity, kg / m s.

p

Density, kg/m^.

G Surface tension, N/m.

T Time, S.

Xy Shear stress, N/m^.

(j) Arbitrary dependant variable ((j) = U, V, W, h, K and e).

(j) Normalized crank angle degree.

SUBSCRIPT

a Air. e Effective.

[ Liquid.

Pg Pseudo turbulence.

758
759
Fig . 3 Effect of various values of drag force on the fuel

concentration along the spray centerHne, at P^j=120 bar

and P, =5 bar

Fig. 4 Effect of various values of drag force on the

spray concentration aloing centerHne at

w =749 rad/s and P. =40 bar

760
Velocity vector scale 30 m/s

T = 0 ms r=0.'/"/ ni X = 1.33 ms

Fig. 5 Effect of spray penteratLon on the air field,

to = 1135 rad/s and P. =40 bar ^3

Velocity vector scale 30 m/s

r : 0.0 m s T - 0.77 ms r - 1.33

Fig. 6 Effect of spray penetration on the air field,

(o = 749 rad/s and P. =40 bar ^3

761
Fig. 8 Effect of ambient temperature on the spray

concentration along the spray centerline,

P. .= 120 bar and P,= 5 bar

inj b

762
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper85

STUDY OF AN AIR BLAST COAXIAL ATOMIZER

EXPERIMENTS AND MODELISATION

I. Care and M. Ledoux

URA CNRS 230/CORIA - University of Rouen

Mont Saint Aignan, Fnuice

ABSTRACT

A coaxial air blast atomizer as used in rocket cryogenic engines has

been experimentally and theoretically studied. Structure of the spray, map

of spray sizes, evolution of air velocities and liquid have been

investigated. A model is given to predict drop sizes (SMD) , spray angles,

and skimming kinetics (atomisation) . This model is assessed by present

experiments and datas from previous Diesel sprays studies.

INTRODUCTION

Air blast atomization is widely used. In cryogenic rocket engines

(like HM7 engine of european rocket Ariane) atomization of a liquid (or

supposed so) jet of oxygen is achieved by a gaseous coflowing hydrogen

flow. Structure of such a pulverisation is studied here through

experiments and modelisation

.

In such pulverisation devices great relative velocities between gas

and liquid lead to a skimming of the surface : an attempt will be done

here to calculate both size and number (skimming kinetics) of drops

formed. The problem presents many similar features with Diesel injection.

Present results should then find an application in this field.

EXPERIMENTS

Apparatus

The atomization of a liquid cylindrical flow of initial radius a^ by

a surrounding annular air flow of lateral dimension h has been achieved

through a device described on figure 1. Particular attention has been paid

to axisymmetry (non swirling) of both flows. Three configurations have

been studied (A to C) . Table on fig.l give details.

Water and kerosene have been used as substitutes to O^ . Velocity of

liquids was 1 . 4m/s . Air was used as atomizing fluid. Velocities were

varied from 100 to 220m/s at the nozzle exit.

Photographic (short time -10|j,s- and long time exposure) investigation

763
were achieved as well as LDV (for air) , and Malvern spray sizing.

Atomizer a (mm) h (mm) (mm)

h 1.95 0.53 0.

B 1.95 0.53 2.5

C 1.95 0.93

Figure 1

The atomizer

Different procedures of enlightening of the spray may show the liquid

core or the fine (scattering) spray.

Dimensions of the experimental set-up are representative as well as

relative velocities of those encountered on real rocket engines.

Structure of the spray is shown on scheme (Fig.l)

Near the nozzle an atomization zone (fine spray) takes place. The

radius a (x) of the liquid is then decreasing downstream. The skimming of

this surface is not sufficient to atomize completely the liquid, then a second zone of coarse pulverisation is found. All present study will be

confined to first atomization zone.

Short time exposures show presence of large (order of a/2) structures

in the atomization zones. Owing to the very small structures of

instability leading to drop formation (A/2) long exposure time photographs

have been used for determination of a time-mean decrease of liquid core

Experimental results

Following parameters have been measured : -Radial profiles of drop sizes (given in terms of SMD values) at

different abcissas x. Malvern 2600 particle sizer was used. Due to

axisymetric structure an inversion process, similar to that of Gomi [2]

has been employed [1]

.

-Mean air velocity fields around a solid cone simulating the liquid

cone

.

-Photographic investigation of liquid cone shape (measurement of

a(x) . Injection parameters are U_ , a, h, L„ G R Some results are shown on figures 2,3 and 4.

Similar to Gomi's results [2] one can observe a radial increase of

SMD's from surface to external part of spray. This can be explained by

inertial effects of drops : greater droplets have greater response time

764
d3.2 (uffl)

Figure 2 : The radial evolution of

SMD at x=30mm

from the nozzle

Figure 3

Influence of air

velocity U on drop

G

sizes at different

downstream positions

d3.2 (um)

NOZZLE A

Dx - 10 aa Ox

Ax • 30 tuD

^x^zzLE A

d3.2 (um)

D eau/air O k^rosJne/air

NOZZLE A

Figure 4

Influence of

surface tension

765
T a ) and keep their radial velocity on a greater distance. Smaller

drops are immediatly entrained by the gas.

A smaller surface tension leads to smaller drops (reduction of the

stabilizing effect of a)

.

According to following atomization theory, increase of AU leads to

smaller droplets (increase of the destabilizing effect of AU)

.

Comparison between configuration A and B shows that an increase of

leads to higher drop sizes, (not shown here)

.

A complete set of results will be found in [1]

.

MODELISATION OF THE ATOMIZING ZONE

Drop size away from the surface (r > 5mm) can be influenced by

several factors : 1) Aerodynamical effects quoted above. 2) Secondary

atomization (breakup of drops by high velocity air stream) . So

calculations will be attempted here only for measurement made near the

surface

.

The authors developed previously a complete linear model of the

instability of a (swirling) liquid jet in a (swirling) coflowing air

current [1] . Due to high relative liquid-air velocities this model is in

fact asymptotic to the classical Kelvin-Helmholtz theory for capillary

waves at a simple interface. The dominant wavelength A of maximum growth fl is then found [1]

.

A

SiTcr

(3.1)

2Pg AU^

'L J

1 / 2

(3.2)

and the propagation velocity of the wave c is

c = (p^^U^^ + p^Uj )/ ~ according to p^/pj^=0(10 ^) (3.3)

From A and Cl one can deduce different parameters, of the atomizing

zone : SMD, (initial) angle of spray cone, skimming kinetics. Here we have

to deal with two aspects of pulverisation : non linearity at breakup and

polydispersion . This can be achieved in a simplified way.

Drop diameters : determination of SMD

The wave is sinusoidal only in its initial part. Then the

perturbation has the form : ti = Tj^ exp Qt exp i (Kx - Oi^t) where (0^, = kc

.

The velocity of the crest is drj/dt = Of] where d/dt is a Lagrangian

derivative. One assumes that, at breakup, the wave has the form of a cylindrical ligament of diameter A/2 (non linear theory of Nayfeh [3]

gives an insight of this) and amplitude AA where A is an unknown constant.

In reality atomization results in a spectrum of drops of diameters

being the number of drops of diameter d^ formed, one can write the

conservation of mass for a ligament

d,-N,

N. 3A AV8 (3.4)

One can also conserve surface tension energy Z Nj^d^= aA^ (3.5)

766

I
SMD is defined as : d3 2= ^

^i*^?/ S (3.6) Unknown constant A is

then eliminated, and some effect of polydispersion is included in

definition of d3 2

• <^3 2^ 3/4A (3.7)

Mean number of drops formed by ligament is then of order A : N=(8/9)/A (3.8)

Skimming kinetics

An important parameter, not yet found at our knowledge in the

litterature, is the mass flow rate of drops teared off by unit surface.

This is needed to predict the evolution of the diameter of the liquid

cone

.

The number dN^ of drops formed on a length dx of liquid cone during

dt can be written :

, 64 IT A a d^N^ = C, dx dt (3.9)

9 t^

where a (x) is the local diameter of the cone and t^. the lifetime of a ligament. C^^ is an unknown constant (or assumed so) . Many authors write [8] : t^=C2/Q (which is equivalent to : ^rupture/^o= ^2 )

•

64 '^a Then d^N^ = Q dr dt (3.10)

" 9 ^2

From a mass balance in the liquid core, assuming that the liquid

velocity is constant :

da IT , , p, TT n 2a — dx dt = p, — d^ , d^N^ (3.11)

So :da/dx = -C f^^/U^ (3.12)

AC^

Here constant C, found equal to 0.78 is introduced to summarize

all unknowns A, C^^ , C2 . . . ) . In case of total atomization, a length cone could be calculated :

L^= ^2\fl a^Uj^/9TrcAu)^ {p^/p^)^^^ (3.13)

This expression can be compared to expressions obtained by

Cheroudi [6] and Aral [7] , in the case of Diesel sprays : L/d =c' [p^/pg]^^^ (3.14)

Initial spray angle

Following an idea proposed by Reitz [4] and Chigie^ [5] the initial

spray angle 0 can be deduced from the velocity vector V of drops teared

off from the surface. The axial component of V is U^^ and the radial

component can be identified with dr]/dt.

767
Using relations (3.1) and (3.2) one finds

e 2ttaAU

tg — =

2 \fT" n

1 / 2

= A — (3.15)

Some comparisons can be made with studies of Diesel sprays found in

the litterature.

Reitz [9] found a relation of the form :tg6/2 = 0.188 A fJ/U^^ leading

to A = 0.188.

e

Reitz and Diwakar [10] found experimentally : tg — = 0,7

1 / 2

Also can be quoted similar expression by O'Rourke and Amsden [11] and

Ranz [12] . Taking A = 0.188 (Reitz) one founds : 0.68 instead of 0.7 four our

atomizer. In the case of Diesel sprays Au ~ and the result is consistent

with the constant 0.7 of Reitz and Diwakar.

ASSESSMENT OF THE MODEL

Determination of SMD

Calculating SMD was possible only for the vicinity of the surface

(owing to secondary atomization and aerodynamic effects at r > 5mm) . A

comparison is shown on fig. 6 between experimental and calculated values

for water and kerosene. A pretty good agreement is found.

D32 Calculated

"Water

A-Kero

10 20 30 FIG. 5 Comparison between measured and calculated SMO

at X = 10mm and r = 5mm

40 50 D32 Measured

m/s

1 23

1 47

1 70

1 95

2 20

da

dx

0.35

0.27

0.41

0.38

0.44

0.022

0.014

0.018

0.015

0.015

da

dx

0.53

0.50

0.36

0.50

0.50

0.029

0.023

0.014

0.017

0.015

TABLE 1

Determination of C

From (3.12) C can be deduced from measurement of da/dx. As the liquid

core is not too far from a cone (Uj^ , Q and A varying too fast) da/dx can

be deduced from photographs. Results are shown on table 1 -for water and

kerosene jets-. More extended experiments should be needed to assess C as

a constant, but if assumed so (which is not too unrealistic) C is found

here equal to 1.8.10"^ with a standard deviation of 0.5. 10"^.

Let's remark that with A = 0.188 and C2 of order 10 (as assumed by

768
many authors [8] is of order 0.5 - 1.

We did not succeded to determine accurately initial angle from

photographs (fine spray limits are difficult to appreciate) but an order

of 0.15 for A could be deduced from some views.

CONCLUSION

The experimental study of a coaxial air blast atomizer showed a complex structure. Particular attention was paid to atomization zone (fine

spray) . A model of atomization in this zone presented here allows the

determination of the particle sizes (SMD) near the surfaces, the length of

the liquid cone (or the evolution of radius of this zone when atomization

is not completed), the skimming kinetics and the initial spray angle.

A more complete determination of the spray should involve : 1) Aerodynamic in the fine spray zone as well as the secondary atomization.

2) The dynamical behaviour of the non-atomized liquid core and its

break-up, which in practical cases, may lead to coarse pulverisation and a bad combustion quality.

Nomenclature

a radius of the liquid jet

C skimming kinetics constant

k wave-number = 2tr/A r radial coordinate

velocity of gas (air)

velocity of liquid

x axial coordinate

Ti instability

CT surface tension

A dominant instability wavelength

dominant instability growth coefficient

BIBLIOGRAPHY

[1] CARE I. - Etude d'un injecteur coaxial assiste. Ph.D Thesis Rouen.

Dec. 1990

[2] GOMI H. - Pneumatic atomization with coaxial injectors. Ph.D Thesis

Sheffield. 1984.

[3] NAYFEH A.H. - Phys . Fluids 1970, 1970, vol.13 (4), p. 841-847.

[4] REITZ R.D., BRACCO F.V. - Phys. Fluids 1982, vol.25 (10),

p. 1730-1742

.

[5] RUIZ F., CHIGIER N. - The mechanics of high speed atomization Iclass

1985 proceedings.

[6] CHEROUDI B., CHEN S.H., BRACCO F.V., ONUMA Y. - SAE Paper 850 12 6.

1985.

[7] ARAI M., TABATA M., HIROYASU H., SHIMIZU M. - SAE Paper 840 275.

1984 . [8] DOMBROWSKI N., HOOPER P.C. - Chemical engineering Science, Vol.17,

1962, p. 291-305.

769
[9] REITZ R.D. - Atomization and spray technology 1987, Vol.3,

p. 309-337.

[10] REITZ R.D., DIWAKAR R.- SAE Paper 870 598, 1987.

[11] O'ROURKE P.J., AMSDON A. A. - SAE Paper 872 089, 1987.

[12] RANZ W.E. - Canadian J. of Chem. Engineering, 1958, p. 175-181.

770
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 86

CHARACTERIZATION OF A LOW FLOW RATE/HIGH CHARGE

DENSITY ELECTROSPRAY

A. Gomez and K. Tang

Department of Mechanical Engineering

Yale University

New Haven, CT, U.S.A.

ABSTRACT

Detailed measurements of droplet size and velocity distributions by

phase Doppler anemometry were performed in a low-flow-rate/high-charge-density

electrospray of methanol, producing micron-size droplets. Complementary

information was also obtained by pulsed shadowgraph imaging of the spray

break-up. When the spray was operated in the so-called rainbow mode,

characterized by the emission of high order Tyndall spectra under white light

illumination, small droplets were generated with average diameter of the order

of 2 |lm and narrow size and velocity distributions. Both velocity and

electric field monotonically decrease as function of the distance from the

thin jet formed at the liquid meniscus, where the field is most intense.

Electric fields as high as 10^ V/m were achieved in this spray. Velocitydiameter correlations showed that the charge to volume ratio varies

approximately with the inverse of the square of the droplet diameter. Coulomb

explosion is unlikely to play any role in the formation of this spray.

INTRODUCTION

One of the simplest implementation of an electrospray consists of a

small metal tube maintained at several kilovolts relative to a ground

electrode few centimeters away. If a liquid of sufficient electric

conductivity is fed into the tube, its meniscus at the outlet of the capillary

takes a conical shape under the action of the electric field, with a thin jet

emerging from the cone tip. This jet breaks up a short distance downstream

into a fine spray of charged droplets, which, under certain conditions, may

have a possibly narrow droplet size distribution. In fact, almost forty years

ago some observations of High Order Tyndall Spectra were reported in the

scattering of white light by "smokes" produced in electric dispersion of

liquids (1,2) . This optical phenomenon is indicative of droplet formation of

narrow size distribution in the size range 0.1-2 |J,m (3). Whereas studies,

both experimental and theoretical, of the electrohydrodynamic phenomena

leading to the formation of the liquid conical meniscus have been reported (4-

9) , no quantitative measurements of droplets sizes in this regime have been

reported to date, with the exception of some recent work of Fernandez de la

Mora et al . (10) on the generation of submicron monodisperse aerosols in

771
electrosprays of glycerol. Consequently, it is in the area of the spray

formation that we have concentrated the initial efforts discussed here.

The selection of the spray was motivated by the following consideration:

electrosprays of water/methanol solutions operated in the same conditions as

in the present study are used to ionize very effectively large macromolecules

of biological interest, which are subsequently introduced in a massspectrometer. This approach is bringing about a revolution in the field of

mass-spectrometry , as documented by the pioneering work of Fenn ' s research

group at Yale (11), and has resulted in probably the only practical

application of this type of electrospray to date. A preliminary account of

part of this work was presented elsewhere (12).

EXPERIMENTAL METHODS

Methanol at a flow rate of 0.003 cc/min was fed into a stainless steel

metal capillary (I.D.=0.20 mm) by a syringe pump. The tube was maintained at a

voltage of about 3 kilovolts relative to a ground electrode positioned 3 cm

away. In a typical experiment, while methanol was pumped through the system,

the voltage applied to the capillary was progressively increased until a

diffuse spray of very fine droplets was formed, which under white light

illumination exhibited the characteristic Tyndall Spectra. At these operating

conditions, hereafter referred to as rainbow mode, detailed single point

measurements were taken throughout the spray by a commercial Phase Doppler

Anemometer (Dantec Electronics), which yielded simultaneous inference of

droplet size and axial velocity distribution, as well as size-velocity

correlation. Some pulsed shadowgraph imaging of the spray break-up was also

attempted. The total magnification of the system (= 800X) did not permit a

quantitative use of the images; nevertheless, some qualitative features of the

break-up were observed. Further details on the imaging technique are given by

Gomez and Tang (13) . The system electrical behavior was monitored by

connecting the virtual ground plate to an oscilloscope, thereby recording the

current collected by the electrode. All experiments were performed at ambient

temperature and atmospheric pressure.

RESULTS AND DISCUSSION

Characteristic size and velocity distributions of the methanol spray

operated in the rainbow mode are shown in Fig. 1. The size distribution appear

roughly symmetric, with a remarkably small ratio of standard deviat ionover

mean diameter of 0.13. This result already shows the electrospray capability

to generate narrowly disperse droplets with a size two orders of magnitude

smaller than the diameter of the spray orifice. Clogging problems that

characteristically prevent monodisperse aerosol generators from operating in

this size range can be circumvented by following this approach. The velocity

distribution is even narrower than the size distribution, suggesting that at a

given location the velocity is essentially independent of particle size; this

finding can be exploited to infer information on droplet charge density, as

shown below.

The spray evolution was probed in both axial and radial scans. As a

validation of the Phase Doppler Anemometer measurements, total mass

conservation was verified within about 14% by integrating droplet average size

and velocity over the spray radius, at a fixed axial coordinate, and

comparing the result with the metered liquid flow rate.

772
2.71 [sec] Samples: 11639

4.5 [kHz] flee. /Sph. 94.7/100.0

Data File: MERl Elapsed Tine:

Date/TiHe: 900718/15:14:07 Data Rate:

3,\

0

0

\

• 0z

VI

c

i0

(J

,52E>04

,26E+04

,00

,00 14. 28.

Dianeter [nicronsl

,28E+04

,14E>04

,00

,00 12. 23.

U Velocitu In/sl

Concentration

Uol. fraction

Uol. flUK

4.84E+04 /CO

6.66E-0S '/. 3.60E-04 cc/oMi/s

Mean Dianeter (DIO) 2.93

Area Mean (D20> 2.96

UoluMe Mean (D30> 2.99

Sauter Mean <D32> 3.03

Mean Velocity

RMS Velocity

5.37

.426

Microns

Microns

Microns

Microns

m/s

m/s

Fig. Representative size and velocity distribution

Jet break up

Qualitative aspects of the spray break-up were

investigated by pulsed shadowgraph. It was

observed that under the rainbow mode of

operation: i) the liquid exits the cone formed at

the outlet of the charged capillary as a thin,

stable thread that persists for a short distance;

ii) this ligament then breaks up into an array of

droplets; iii) droplets move downstream and are

displaced radially because of coulombic repulsion

from neighboring droplets, thereby leaving a wavy-

"signature" under pulsed illumination; iv) once

displaced radially, droplets tend to follow

diverging lines of field because of both

relatively small inertia and high charge to mass

ratio, as shown below. They eventually form a

fan which coincides with the visible

electrospray . Figure 2 is a digitized picture

showing some details of the break up, as

ii) and iii) . Because of

the resolution of both imaging

system and PDA no further probing was possible to

see if satellite droplets were generated at

break-up as observed in sprays of larger droplets

(13) . However, a rough comparison is still

possible, at least on the basis of average

quantities. In particular, it is interesting to

address the question of Coulomb explosion which

occurs when the so-called Rayleigh limit is

attained and electrostatic repulsive forces

overcome surface tension forces.

discussed in

limitations in

* HI

*Hi

1^

Fig. 2

773
This process inevitably takes place at some point during the droplet

evaporation, since there is experimental evidence that charge remains attached

to the droplet while it is evaporating (14). The question is if this mechanism

determines or, at least, contributes to the droplet dispersion and the

resulting spray formation. In a related study on electrostatic sprays of

heptane quite similar to the one examined here, except for the fact that they

produced larger droplets (13), we showed that there is no evidence of Rayleigh

disruption in this type of electrospray if the average charge to mass ratio is

below about 60% of the Rayleigh limit. This methanol electrospray has a high

charge density of 6.2 10"^ C/cc, which however corresponds to only about 35%

of the Rayleigh limit based on an average diameter. This observation and the

similarity of the break up patterns observed in the two studies suggest that,

in analogy with the heptane sprays. Coulomb explosion is not occurring in the

probed region of the spray.

Velocitv and Electric Field Evolution

Figure 3 shows the average droplet diameter and average axial velocity

component measured along the axis of the jet, with the origin of the

coordinate system taken at the capillary outlet. The spray shows a diameter

monotonically decreasing as function of the axial coordinate, as a result of

evaporation effects. Monotonically decreasing diameters are also observed

when scanning the spray in the radial direction. These observations are

consistent with the visual observation of Tyndall spectra, which are typically

attributed to monodisperse aerosols in the (0,1-2 flm size range) (3). The

velocity also monotonically decreases along the axis, from around 26 m/s to

about 4 m/s in the scanned region. Thus, the liquid is first rapidly

accelerated in the conical meniscus into a thin jet, that breaks up into a

stream of droplets, which in turn rapidly decelerate to the measured

velocities

.

n

' 1

• 1'

1

' 1'

1

'—

I

• D

Axial Position (mm)

Fig. 3: Average diameter and axial velocity component versus axial

coordinate

774
To understand this behavior, let us consider the equation of motion of an

isolated droplet under the action of electrostatic, gravitational and drag

forces

:

Pd^^=^(Pd-Pg)g + "eeE - ci^ (1+0. ISSRe^/^) 37C|id(V-Vg)

(1)

where at the left hand side of the equation is the inertia term with a

lagrangian derivative, and the right hand side terms are the gravitational

term, the electrostatic force and the droplet drag, respectively. Here, V is

the velocity, d, the droplet diameter, ne denotes the number of charges per

droplets, e is the elementary charge, E is the electric field. Re is the

Reynolds number, Cd is the drag coefficient and V-Vg is the relative velocity

between the droplet and the surrounding gas. Under steady state condition,

the time derivative equals the convective derivative, which, evaluated on the

centerline of the spray, is given by Vji^Vz/dz. An order of magnitude analysis

yields negligibility of the gravitational term. Since the Reynolds number is

0(1), the drag force can be approximated to the Stokes formula by setting

CD=24/Re. We will also assume negligible gas velocity, i.e. VgSO , which

amounts to neglecting the gaseous entrainment in the otherwise stagnant

atmosphere were the spray discharges. Thus, the axial component of the

equation above reduces to

ay Pd^Vz^ = neeEz- (l+O.lSSRe^^) STC^idVj

6 dz (2)

since size and velocity are known, from eq. 2 we can try to infer the local

electric field, which is very important for the determination of limits of

operation of the electrospray , for spray modelling and can also help future

analyses of the complex electrohydrodynamics in the liquid meniscus. To this

end, a value for ne can be obtained from

Q 6 (3)

where I is the total measured current, Q is the volume flow rate and dg is the

volume average droplet diameter as measured at the very beginning of the

spray. Substituting eq. (3) into eq. (2) and solving for E^, we obtain

Indl

Pd^Vz^ + (l+0.158Re2/3) 3n[idW,

6 dz

(4)

The resulting electric field is plotted in Fig. 3. Electric fields of the

order of 10^ V/m are generated, with the strongest field near the breakup

region of the spray, as expected since the liquid ligament can be regarded as

a "sharp point", which, as is well known, is surrounded by very intense fields

when charged at a high potential. The velocity profile is also superimposed

on the same plot. Notice that when the inertia term in eq. 4 can be

neglected, which holds true everywhere in the spray except in the initial

775
2 4 6

Axial Position Z (mm)

Fig. 4: Average droplet velocity and electric field along the spray axis.

region where the deceleration is most rapid and the velocity gradient high^,

eq. 2 reduces to a balance between electrostatic force and drag.

Consequently, the reason why droplets rapidly decelerate is that once

generated at the jet break-up they relax within a short distance to the local

terminal velocity which roughly decreases as the electrostatic field. The

field calculated by this procedure is the total field acting on the droplet

and accounts for both the external field applied between capillary and ground

electrode as well as the "internal" field due to the coulomb interaction

between charged droplets.

Charge Density

From local simultaneous measurements of size and velocity distribution, we can

also gather information about charge density. Figure 5 shows the

size/velocity correlation as obtained with the PDA at a typical location along

the spray axis, where both the inertia term and the Reynolds number correction

to the drag term in eq. 2 are negligible. The velocity is essentially uniform,

regardless of particle size. Eq. 2 reduces her to

n^eEz = 37i|XdVz

^Estimates of the Stokes number, defined as the ratio of a droplet relaxation

time over characteristic flow time (15), yielded values between 10"^ and 10"*,

indicating the negligibility of inertia over most of the spray region.

776
12 3 4

Diaireter (micron)

Fig. 5: Typical velocity-size correlation.

At a given location, the electric field is constant and the velocity is

essentially independent of droplet size, as per Fig. 5. Consequently, n© <^ d

and the charge to volume ratio should scale as the inverse of the square of

the diameter.^

CONCLUSIONS

Detailed measurements of droplet size and velocity distributions were

performed along the axis of a low flow rate/ high charge density electrospray

of methanol, producing micron-size droplets. The following conclusions can be

drawn from these measurements:

1) When the spray is operated in the rainbow mode, corresponding to the

emission of Tyndall spectra under white light illumination, small droplets are

generated with average diameter of the order of 2 )j.m and narrow size and

velocity distributions;

2) Both velocity and electric field monotonically decrease as function of

the distance from the thin jet formed at the liquid meniscus, where the field

is most intense. Electric fields as high as 10^ V/m were determined;

3) Velocity-size correlation showed that charge to mass ratio varies

approximately with the inverse of the square of the droplet diameter.

4) Coulomb explosion is unlikely to play any role in the formation of this

spray

.

-^This dependence was independently confirmed by deconvolving the total current

measurement, as well as size and velocity measurements obtained in a radial

scan. This approach also confirmed that the charge to mass ratios are

substantially below the Rayleigh limit in the probed region.

777
ACKNOWLEDGMENTS

We are indebted to Profs. J. B. Fenn and J. Fernandez de la Mora, for manystimulating discussions and helpful comments. Thanks are also due to Dantec

Electronics for allowing us to use their particle measurement systems in

extensive laboratory demonstrations. This research was in part supported by an

NSF Equipment Grant #CBT-8806608 and a NASA Grant #NAG3-1259.

REFERENCES

1. Vonnegut, B. and Neubauer, R.L., J. Colloid Sci . , 7, p. 616 (1952).

2. Vonnegut, B. and Neubauer, R.L., J. Colloid. Sci., 8, p. 551 (1953).

3. van de Hulst, Light Scattering by Small Particles, Wiley, 1957.

4. Zeleny, J., Phys . Rev., 10, p.l (1917).

5. Taylor, G.I., Proc. Roy Soc. Lond. A, 313, p. 453 (1969).

6. Hayati, I., Bailey, A.I. and Tadros, Th. F., J. Colloid. Interface.

Sci., 117, p. 205 (1987a).

7. Hayati, I., Bailey, A.I. and Tadros, Th . F., J. Colloid. Interface.

Sci . , 117, p. 222 (1987b)

.

8. Smith, P.H., IEEE Trans. Industr. Appl . , IA-22, p. 527 (1986).

9. Turnbull, R.J., IEEE Trans. Industr. Appl., IA-25, p. 699 (1989).

10. Fernandez de la Mora, J., Navascues, J., Fernandez, F. and RosellLlompart, J., J. Aerosol Sci., 21, Suppl . 1, p. 673 (1990).

11. Fenn, J.B., Mann, M., Meng C.K., Wong, S.F. and Whitehouse, CM.,

Science, 246, p. 64 (1989).

12. Gomez, A. and Tang, K .," Characterization of Liquid Fuel Electrostatic

Sprays", Paper presented at the Western States Section/The Combustion

Institute 1990 Fall Meeting, La Jolla CA.

13. Gomez, A. and Tang, K., "Atomization and Dispersion of Quasimonodisperse Electrostatic Sprays of Heptane" , Proceedings of the Fifth

International Conference on Liquid Atomization and Spray Systems,

ICLASS-91, Gaithersburg, MD, U.S.A., 1991.

14. Abbas, M.A. and Latham, J., J. Fluid Mech. , 30, p. 663 (1967).

15. Fuchs, N.A., The Mechanics of Aerosols, Dover, 19 54.

778
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 87

CHARACTERIZATION OF THE SPRAY STRUCTURE OF

A LARGE SCALE H.F.O. ATOMIZER

G. De Michele, M. Graziadio, F. Morelli and G. Novelli

ENEL-C.R.T.N., Pisa, Italy

ABSTRACT

The spray produced by a large scale h.f.o. Y-jet nozzle has been

characterized in cold conditions using simulating fluids and two different laser

instruments (diffraction and phase-doppler techniques) . The study has been

carried out to obtain a detailed spray description to be utilized in the

prevision of combustion performances.

It has been found that some important variations of the spray structure

cannot be completely described using only mean spray parameters, such as mean

drop diameter and velocity. For this reason, in addition with the mentioned

values, the analysis has also regarded the spray angle and the position of the

largest drops in the cross section.

All variations of spray structure have been interpreted with reference to

different fluidynamic regimes inside the nozzle. General relationships have been

obtained which allow the spray quality to be predicted, as well as the spatial

position of larger drops in the spray cone.

INTRODUCTION

Y-jet atomizers are widely employed in power stations due to the

versatility of operation and reliability. Notwithstanding their extensive use,

the number of published works on this class of atomizers are relatively few

[1-4].

An extensive work on burner development, at the Nuclear and Thermal

Research Center (C. R.T.N.) of ENEL (the Italian Electricity Generating Board),

has included research carried out to characterize a Y-jet nozzle in many

different operating conditions. The work has allowed the definition of

correlations between input parameters and spray characteristics.

EXPERIMENTAL APPARATUS

Test rig and atomizer

The atomization rig "I.S.A." employed during the tests has been described

elsewhere [4] . The rig is operated by using a mineral oil which at 30°C has the

same physical and rheological properties as h.f.o. in industrial burners.

Compressed air is used to simulate the atomizing steam.

779
AIR / SIBVI

OIL

Fig. 1 Geometry of the atomizer lY-45-55-49,

measurement volume

light guide

WLVERN RECEINGi

2 Relative position of the laser instruments and the atomizer.

3a Interpolation maps of DvO . 5 Fig. 3b Interpolation maps of the

on the spray cross-section. percentage of drops larger

than 225 urn.

780
The trials have been carried out with a single port Y-jet atomizer labelled

lY-45-55-49. The Y-duct geometry, shown in Fig. 1, is the same characterizing an

industrial multi-hole nozzle used with the following conditions (for each Yduct)

:

Wo = 0.050 Kg/s Po = 1.4 MPa Ps = 1.15 MPa

Measurement technique

The spray quality is measured using a Malvern 2600 D particle sizer, while

the drop velocity is obtained with an Aerome tries PDPA. The optical arrangement

of the instruments is shown in Fig. 2. To reduce the obscuration during the

trials with the Malvern analyzer, two light guides are employed.

A reference X-Y-Z coordinate system is defined with the Z axis orthogonal

to the laser beam and coincident with the axis of the outlet port. The oil port

lies in the Z-Y plane on the negative Y side. All trials are carried out with Z = 250 mm, by moving the nozzle while the laser instruments are fixed.

EXPERIMENTAL RESULTS

Evaluation of mean spray characteristics

The evaluation of mean spray characteristics on a fixed cross -section

requires the measure of local diameters and velocities followed by an integration

based on droplet concentrations. This procedure is time consuming, except when

the spray is axisjrmmetric [5].

To verify this possibility, the present investigation has started with the

complete spray mapping at nominal conditions Wo=0.050 Kg/s and Pa = 1.15 MPa.

Granulometric measurements have been obtained with the Malvern on six equispaced

diameters on the cross-section (Fig. 2) and, using these data, drop size maps

have been calculated.

Figure 3a shows the map of mass median diameters. The dropsize isocontours

are slightly asymmetric around 0 (Z axis) , This result is confirmed by the map

of the percentage of drops larger than 225 /xm, shown in Fig. 3b.

The asjnnmetric structure of the spray is in agreement with other results

obtained by ENEL with Y-jet nozzles of different geometry [4] . Due to the lack

of spray axisymmetry, the study of Y-jet nozzles at different operating

conditions is prohibitive if the above-mentioned mapping is used.

Figures 3a and 3b however suggest that a simplified calculation of mean

spray quantities can be obtained by evaluating the means only on A-A' . The

highest values of drop size distribution, in fact, are often found near diameter

A-A' as a consequence of its position in the plane Y-Z of the oil and outlet

ports [4] . For this reason A-A' can be utilized to detect the main spray

asymmetries and to evaluate rapidly mean diameters which never underestimate the

spray quality all over the cross-section. This simplified approach has been

adopted in the spray characterization described in the following section.

Spray characterization

The spray characterization has been carried out with Wo = 0.050, 0.047,

0.041, 0.036, 0.030, 0.025, 0.019 Kg/s; at each flowrate the nozzle has been

operated with Pa = 1.15, 1.0, 0.9 and 0.8 MPa. In this way, 28 different

operating conditions have been studied by measuring the spray quality over A-A'

The DvO . 5 profiles at different oil flowrates and the same air pressure are

shown in Fig. 4. The variation of volume concentration C with Y has the same

trend. Both profiles demonstrate that, at a fixed air pressure, the liquid

781
200

150 -

E

a. w 100

6

>a

50 -

Pa = 0.8 MPa

Wo , Kg/s

0.050

0.041

0.030 —e—

0.019

-60 -40

I I I -20 0 20

Y , mm

200

40 60

150 -

E

a.

in" 100

o

>o

50 -

Wo = 0.050 Kg/s

1 r

-60 -40 -20 0 20

Y , mm

40 60

Fig. 4 Dropsize profiles along Y-direction

at different oil flowrates.

Fig. 5 Dropsize profiles along Y-direction

at different air pressures.

0,06

0,05 -

0,04 -

^. 0,03 o

0,02 -

0,01 -

Pa , MPa

0.6

Poor combustion

0,8 1 1,2 Po , MPa

1,4 1,6

Fig. 6 Optimum operating range of atomizer lY-45-55-49.

782
flowrate strongly influences the spatial position of the largest drops in the

cross section and consequently the liquid mass spatial distribution.

Using the Aerome tries PDPA, it has been found that the velocity profile

varies with Wo, but there are no pronounced asymmetries as those existing in the

case of drop diameters.

When the liquid flowrate is fixed, the atomizing air pressure influences

the spray structure, as shown in Fig, 5. In this case the velocity profiles are

slightly asymmetric.

ANALYSIS OF THE EXPERIMENTAL RESULTS

Definition of the nozzle operating range

The atomizer operating range has been examined in order to define a zone

of good performance in real combustion trials.

For this scope a mean spray diameter DvO.5 M has been calculated in each

condition and compared to that of nominal operation (Wo = 0.050 Kg/s , Pa = 1.15

MPa) . By the exclusion of the conditions corresponding to diameters higher than

the reference, an optimum combustion zone has been defined. Fig. 6.

However, the mean drop diameter is not the only factor determining good

combustion performance, and other spray parameters (such as liquid fluxes and

spray momentum) must be considered.

Figure 7 shows, for instance, the shapes of opposing sprays in the case of

a multi-port atomizer with several Y-ducts similar to the one studied. Each spray

angle has been drawn in accordance with the drops ize profiles found for the lY45-55-49 nozzle, assuming no interaction between adjacent sprays and no

deflection induced by combustion air. Two operating conditions are shown in

Fig. 7, each corresponding to an acceptable mean spray diameter. Notwithstanding

that, the inner and outer spray angles and the location of drops larger than 225

fixa are strongly influenced by input conditions, thus probably giving different

combustion performance for the two cases.

Prevision of spray characteristics

An attempt has been made to correlate input data with the mean spray

quality and the spatial location of the largest drops in the cross section.

The values of DvO.5 M have been related with flow conditions using the

relationship proposed by Lefebvre [6] :

DV0.5M/D = A-We"°-^ +B-Zo

The coefficients A and B of (1) can be evaluated only if the atomizing air

mass flowrate Ga and the mean pressure Pm in the Y-duct are known, thus allowing

the calculation of We.

In the present case Ga has been measured while Pm has been evaluated using

the relationship:

_Qa In

K

£i

P2

K

K+1

_Pi

Vi

(K*1)

2

G. = 0

(2)

which has successfully proved to describe the pressure losses in the exit port

of Y-jet nozzles [7]. Using this approach the following relationship has been

obtained:

783
HIGH CONCENTRATION OF

DROPS LARGER THAN 225

— 92*

Kg/s.hole

MPa

m/s

Wo = 0.019 Kg/s.hole

Pa = 1.15 MPa

DvO.5 M = 62 pm

V M = •'?.8 m/s

Fig. 7 Variation of spray structure with oil flowrate

0,08

0,06

at

o

0,04 -

0,02 -

30

0,6 0,8 1 1,2 Po , MPa

1,4 1,6

15

200

Fig. 8 Relationship between the flow Fig. 9

regime in the nozzle and input

conditions

.

Wo , Kg/s

0.050

0.047

400 600

Ga , Kg/m2.s

800

Liquid jet lateral momentxim,

normalized with respect to the

gas axial momentum, plotted

versus the gas mass velocity.

784
Dvo.5 M/D=0.75-We'°"^+a34-Zo (3)

The values of coefficients (A = 0.75 and B =- 0.34) are slightly different

from those calculated in [7] . The differences could derive from scaling effects

and/or from inadequacies in the evaluation of Pm.

The spatial position of the largest drops in the cross section mainly

depends on the presence of liquid films in the atomizer exit port [4] . Figure 8 shows that this approach allows, in the test case, the determination of

three different operating zones.

An interpretation of the spray regimes can be done by considering the ratio

between the liquid momentum and the atomizing air axial momentum [4]

:

2 2

a, = -%-S2- -A. sm^ W

Gl Da a

The dimensionless parameter has been evaluated by estimating Pm as shown

before. Figure 9 shows as a function of Ga for all input conditions. It can be

noticed that, when Ga is constant, varies with Go due to the different lateral

momentum of the liquid and the different of air.

If each operating point is related to the corresponding liquid film regime,

as reported in Fig. 8, it appears that the parameter can be usefully employed

to predict both the liquid film position and the spatial drop distribution.

CONCLUSIONS

The following conclusions can be made:

The atomizer lY-45-55-49 can be operated at oil flowrates that are

different from the nominal ones, provided that the operating conditions

remain within the range of Fig. 6. This optimum zone has been defined

solely by the comparison of dropsize data with those of nominal operation.

The input variables influence not only the mean spray quality but also the

spray angle and the location of the largest drops in the spray cross

section. As these parameters can be important for combustion performance,

the influence of input data on the total spray structure must always be

considered in order to predict the operating region of full scale nozzles.

Predictive correlations have been obtained for the drop diameter and the

location of largest drops in the spray. These correlations will be

systematically tested in future developments using different Y-jet

geometries and conditions

.

785
NOMENCLATURE

A B Coefficients

c Volumetric concentration

D Diameter of the mixing duct

Do Diameter of the oil port

DvO. 5 Mass median diameter

F Frictional loss coefficient

G Mass velocity

k Coefficient of polytropic expansion

P Pressure

V Velocity

V Specific volume

W Mass flowrate

We Weber number

X,Y,Z Cartesian coordinates

Zo Ohnesoree number

Angle between fuel and mixing duct

yP Density

Q> Ratio between liquid and gas momentum

a Air

o Oil

s Steam

m Mean

M Mean on A-A' , Fig. 2

1,2 Inlet and outlet sections

REFERENCES

1. Mullinger, P.J., Chigier, N. , Jour. Inst. Fuel, Vol. 47, pp. 251-261

(1974)

.

2. Bryce, W.B. , Cox, N.W. , Joyce, W.I., I CLASS '78. paper 10-4, Tokyo

(1978)

.

3. Prasad, K.S.L., ICLASS '82. paper 4-3, Madison (1982).

4. De Michele, G., Graziadio, M. , Novelii, C. Giacomelli, A., Andreussi, P.,

9th IFRF Members Conf .

.

Noordwijkerhout (1989).

5. Andreussi, P. Tognotti, L. Graziadio, M. De Michele, G. Aerosol Sci.

and Techn. , Vol. 13, pp. 35-46 (1990).

6. Lefebvre, A.H. , Prog. En. Comb. Sci. Vol. 6, pp. 233-261 (1980).

7. Andreussi, P., Graziadio, M. , Novelii, G., Tognotti, L. , ICIASS Europe '90

Meet, pp. 115-124 (1990).

786
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 88

CHARACTERIZATION AND PERFORMANCE OF SPRAY NOZZLES

FOR WATER DISTRIBUTION IN COOLING TOWERS

L. Tognotti*, S. Zanelli*, B. Bellagamba^ F. Mattachini^ and G. Lotti^

Dipartimento Ingegneria Chimica, Universita' di Pisa, Pisa, Italy

+ENEL-CRTN, Pisa, Italy

ABSTRACT

An experimental study has been carried out on the water distribution

system of a pilot cooling tower of 160 m^/hr . The performances of different

industrial water distributors have been evaluated by changing the operative

conditions of the pilot tower. In particular, the efficiency and the

uniformity of the water distribution have been investigated and compared with

the results obtained in a small-scale loop, in which the single nozzles were

tested. Measurements in both systems, pilot tower and small scale loop,

included the geometric characteristics of the jet umbrella by ensemble

photography, the wetted zone by measuring the specific flowrate, the drop-size

distribution and liquid concentration by high-speed photography. The results

show that correlations exist between the nozzle behaviour in single and pilot

tower configuration. The uniformity of water distribution in the pilot tower

is strongly related to the nozzle installation pattern and to the operative

conditions. Coalescence plays an important role on the drop size distribution

in the pilot-tower. Comments upon the influence of these parameters on tower

behaviour are also included.

INTRODUCTION

The distribution of water to the top of counterflow fill of cooling

towers is a key aspect for the performance of the whole cooling system. It is

a function of nozzle design and installation, height of the spray zone and the

structural cleanliness of the spray chamber. The impact of water distribution

on performance is a combination of uniformity of water distribution, air-side

pressure drops, and heat transfer in the spray zone (1).

As the matter of fact, simulation of cooling tower behaviour by means of

a two-dimensional code (2) showed that drop size distribution in spray zone

plays an important role in tower performances.

In spite of its practical importance, the water distribution process has

not been studied enough. In this work different nozzles were studied in a

pilot tower of 160 m^/hr and in a single-nozzle small-scale loop in order to

find out the parameters to be measured in evaluating the water distribution

performances. On the basis of the results, some performance parameters have

been defined and evaluated, which can be related to heat transfer and pressure

drops in the spray zone of a cooling tower.

EXPERIMENTAL

Two different experimental loops were used in this work. A small scale

circuit was employed in order to study the behaviour of each single nozzle,

while the effect of nozzle installation pattern on tower performance has been

787
(3). In the small-scale

liquid specific flowrate

tested in a pilot-tower of 160 m^/hr potentiality

loop, a single nozzle was studied by measuring the

and the drop size distribution (4,5).

In the small scale circuit, a single nozzle was mounted under a reservoir with

controlled and adjustable head, and measurements were carried out on the jets

or sheets produced by the single nozzle. Measurements included the geometric

characteristics of the jet umbrella by photographing the whole jet, the wetted

zone by measuring the liquid specific flowrate, the drop-size distribution and

liquid concentration by high-speed photography (fig.l)

The pilot tower has been described in (2). The liquid flowrates tested

were: 220, 160 and 100 m^/hr, maximum, nominal and minimum loading,

respectively. The air mean velocities through the tower were 1, 2.5 and 3.5

m/s, in order to consider the typical conditions of natural and forced draught

towers

DISTRIBUTOR j'

AOJUSTABLC

HEAD

COLLECTING PANS

NOZZLE

Fig.l: Small scale

experimental

loop

local mass flowrate in zone 1

( nozzle A. 160 m3/hr, 2.5 m/s)

ta4i«Tt( i«iiui3i«i9i«iriti« numerical and volumetric drop size

distributions

( nozzle B, PI .100 m3/hr.

air velocity 0 m/s)

Fig. 2: Measurement stations in pilot-tower and

examples of local flowrate and drop size

results

Figure 2 shows the configuration of the local flowrate measurement

stations on the top of the packing. They have been chosen in order to

completely characterize three different configurations of water distribution:

central zone, wall zone and corner zone (1,2 and 3 in Fig. 2). The liquid

flowrate has been measured collecting the water in small vessels placed as

described in Fig. 2. The spray characterization has been performed, by means of

high-speed photography, in three points (Pi, P2 and P3 in Fig. 2) inside the

zone 1, at the centre of the packing. A 50 mm macro lens has been used, which

gave a magnification of 0.475 and a distance of 23.5 cm from the lens to the

best focus zone, with a depth of field of about 5 cm. About 500 drops have

been classified and counted for each flow configuration, in order to evaluate

788
statistically representative parameters of drop size distribution.

Four commercial nozzles were tested. Table I reports the main

characteristics of the nozzles. In small-scale experiments the adjustable head

(Fig.l) was chosen in order to provide the pilot-tower nominal loading for

each nozzle (1.77 Kg/s) Nozzle D was tested, in the single configuration, with

an half of the nominal flowrate, since it needed a hydraulic head not feasible

in the small-scale loop.

TABLE I

Single nozzle characteristics

! NOZZLE CHARACTERISTICS

! NOZZLE !

I DIAMETER!

! (cm) 1

STANDARD

HEAD ! (cm)

SPECIFIC MASS!

FLOWRATE ! (kg/s) !

1 A plate cone with

carves

! 3 ! 50 I 1.77 !

! B three cones Y T. 2 1 "'50

1 iVe ^

! c plate with

edges

1 2.6 I 100 ! 1.72 !

! D elicoydal

structure

! 4.8 ! 100 1 0.75 !

RESULTS AND DISCUSSION

Small scale experiments

Visual and photographic observation of the jet, as well as the

measurement of the specific mass flow rate, allow the complete

characterization of the jet from a macroscopic point of view. It is possible

to evaluate:

- size of the umbrella produced by the nozzle;

- the wetted zone under the nozzle at a certain distance from the nozzle (in

this case 1.8 m);

- the zones of the jet formed by coherent liquid (as ligaments or sheets), the

breakup zone and the spray zone.

Figure 3 shows some ensemble photographs of the jets produced, while Fig.

4

shows the results of the macroscopic characterization for nozzles A and B.

Nozzle A produces single cylindrical jets with different initial angle respect

to the horizontal. These jets are produced by carves of different geometry in

the nozzle plate. The jet breakup takes place in the initial zone, due to the

initial perturbations produced by nozzle vibrations. The specific flowrate

changes with 0 (see Fig. 4).

Nozzle B produces three paraboloidal umbrellas of different size. The size of

each sheet decreases as the umbrella enlarges, forming ligaments and then

drops, of size of the same order of magnitude of the ligament size (6), as

shown in Fig. 5. Figure 4 shows the geometry of the jet and the specific

flowrate q at 1,3 m as a function of the distance from the nozzle axis

Nozzle C produces single jets, but narrower than those produced by nozzle A.

Nozzle D forms a swirling paraboloidal sheet, breaking up immediately outside

the nozzle ( fig. 3)

.

As already pointed out, two different breakup mechanisms are recognizable for

the jets investigated. Nozzles A and C present cylindrical jets which produces

789
Fig. 4: Results of the macroscopic characterization for nozzles A and B

790
I—I 50 mm I—I 5 mm

Fig. 5: Example of breakup of a paraboloidal sheet (nozzle B): the sheet;

ligament produced by the sheet breakup; drops produced by the ligament breakup

drops of range dependent on the breakup regime (7). For nozzle B and D breakup

takes place by means of the growing disturbances on the sheet, coupled with

the reduction of sheet size.

From the photographic analysis of the spray zone it is possible to

evaluate the drop size distribution and drop concentration for the spray zone

produced by each nozzle. Table II reports, together with other parameters that

will be defined later on, the diameter of the wetted zone at a distance 1.8 m

under the nozzle and two different mean drop sizes Dpq (8) for the four

nozzles.

It is worthwhile to note that this characterization is important either to

verify the goodness of nozzle installation pattern suggested by nozzle

manufacturers for cooling towers, or to measure quantities needed to evaluate

some nozzle performance parameters as described in the following section.

Definition and evaluation of performance parameters for water distribution

From the characterization on single nozzles, it is interesting to define

and to evaluate some parameters related to the behavior of water distribution

system in cooling towers. In particular, these parameters should provide a

first estimate of the heat transfer and pressure drops in the spray zone, as

well as on the uniformity of water distribution on the tower packing.

As "performance" parameters, one might define:

1) the ratio Sb between the surface of the drops and their volume, which

is related to the heat transfer in the spray zone.

Sb = 6 ^ N.j(d.i)2 = 6/ D32 * [1]

XN.i(d.i)3

2) the ratio between the wetted surface at z=z* and the total surface

under the nozzle.

3) the uniformity of liquid distribution cr„, defined as

aw = [ I(qi - q)2/ (n-l)]i/2 [2]

where qi is the local flowrate and q^ is the mean liquid flowrate.

Parameters 2) and 3) give a first idea on the efficiency and uniformity of the

791
water distribution produced by each nozzle. However, many nozzles operate in

practical systems so that these parameters may not be directly extrapolated to

battery configuration, but they suggest the optimum configuration to be used

in practical systems.

4) the mean liquid hold-up in the spray zone. This parameter gives indication

on the reduction of the net section for air flow due to the presence of the

drops.

Hl = Nv (D3o)3 / 6 [3]

where Nv is the number of drops per unit volume and D30 the volume median

diameter

5) the velocity heads lost per unit length of spray zone. This parameter

is related to the air pressure drops in the spray zone and can be evaluated by

means of the relation (9):

N/Ls = (2g/vo) (qL/q«) / Vo (vr-Vo) [4]

where N is the number of velocity heads lost for a length Ls , ql and qR the

specific flow rates of the liquid and gas, respectively, Vo is the mean air

velocity and Vr the terminal velocity of the drops. Assuming Vo=3m/s as

reference air-side velocity, and evaluating Vr for drops of size D32, the

values of N/Ls reported in Table II are obtained. In the calculation it has

also been assumed that qi is equal to the mean specific flowrate of the liquid

phase and q« is given by the product of Vo and the air density.

Table II gives an interesting set of results on which it is possible to

make comparison among the behavior of the four nozzle in single configuration.

TABLE II

Single nozzle behaviour: performance parameters

! NOZZLE

DIAMETER OF

WETTED ZONE

at H = 1.8 B

the!

(m)!

1 1

D32 1 (mm) 1

D30

1 n

! Vy,

nr

q

. Kg/sm2

! Hl i !xl04!

Sb

m2 /m^

Vr

m/s

N/Ls 1 m-i !

1 A 4 6.5 1 6.0 10.35 1.5 ; 0.14 !"6-"7

! 923 13 0.012!

! B 1.5 4.2 ! 2.5 10.7 0.7 ! 1.00 !4-5 ! 1428 11.2 0.11!

! c 5 4 ! 2.7 !0.9 0.09! 0.093 ! .2 ! 1500 11 0.01!

1 D 3 4.1 ! 2.8 10.85 0.2 ! 0.25 !2-6 ! 1463 11.1 0.028!

Nozzle C and D present a good wetting uniformity and efficiency even in single

configuration, suggesting a slight dependence of these parameters on the

nozzle installation pattern in practical application. Nozzle A gives the worst

behavior, since it presents the largest mean drop size (that is to say the

lowest Sb ) , together with the lowest wetting efficiency coupled with a high 0^

As regards the air flow in the spray zone, the values of Hl are so low, as

expected, that the reduction of the net section for air flow is negligible. It

is more important to consider that the presence of a coherent paraboloidal

sheet (for nozzles B and D) produces a reduction ot the net section for air

flow up to a value of 7% (nozzle C). Finally,

the values of N/Ls indicates that the lowest air-side pressure drops are to be

expected for the nozzle producing the largest umbrella, being the mean

792
specific liquid flowrate the minimum and the D32 quite similar for the four

nozzle in the single configuration.

Pilot-tower experiments

Figure 6 shows an example of the results on local flowrates measurements

for the four nozzles in the central zone, in a three-dimensional

representation. These figures, obtained by interpolation (spline smoothing) of

the measured values, give an idea on the uniformity of wdter distribution by

the different nozzles. Nozzle A presents the worst behaviour in these

conditions, since the local

flowrate is strongly

oscillating inside the

examined area. In order to

compare the behaviour in

single and battery

configurations, a

pilot-tower "ensemble"

standard deviation for the

liquid flowrate can be

evaluated by Eq.[2] by

using different weights for

the mean values obtained in

the three zones of Fig. 2.

Figure 7 reports the values

of as a function of

pilot-tower operating

conditions. The standard

deviations are generally

greater than those for the

single configuration, but

it should be noted that the

same nozzle spacing was

employed for the nozzles

(Fig. 2) and a different

geometrical arrangement for

each nozzle type should

probably result in a

smaller Oy,, However,

information about the

be obtained. The standard

100 nJyU - 3.5 Hz-S 100 nJ/H - 3.5 rvs

100 n3>fi" - 3.5 H/'S 100 nj'fi" - 3.5 rvs

Fig. 6: Three-dimensional representation of local

flowrate for the four nozzles ( liquid

flowrate 100 m^/hr, air velocity 3,5 m/s).

still

air velocity. The effect of liquid

influence of operating condition can

deviation does not seem to be affected by

flowrate is generally to enhance Oy, . Table III shows the values of the Sauter mean diameter of the water drops

as a function of measurement position and air velocity. The position Pi

generally gives the greatest value of D32 and the smallest drop concentration

for nozzles A and B, because of the big drops produced by the dripping

immediately under the nozzle and the poor cooperative effect of the

overlapping ,iets for these nozzle configurations. The "mean" D32 on the whole

section has been calculated by weighing the measured values on the basis of

suitable geometric coefficients. The values are generally greater than those

measured in the single nozzle configuration. The nozzle D "single" D32 was

measured with a liquid flowrate about an half than the pilot tower case and it

is not directly comparable. This increase of mean drop diameter is the result

of the coalescence of small drops occurring when two or more .iets, produced by

793
contiguous nozzles, impact during the fall. This fact is confirmed by the

mass flowrate 160 mc/hr

standard deviation

air velocity 2.5 m/s

standard deviation

o

o

H

NOZZLE

1

,

0 > • 9C

0

-0 i —_

liquid flow/rate 160 mc/lir

SMD (mm)

air velocity (m/s)

120 leO 200 240

liquid flowrate (mc/hr)

1 2 3

air velocity (m/s)

Fig. 7: Effect of operating conditions on the

uniformity parameter Ov

Fig. 8: Effect of air

velocity on SMD

increase of Da 2 values, shown in Fig. 8, for the nozzle for which is greater

the deformation of the umbrella with air velocity (nozzle B). In other words,

nozzle B produces a paraboloidal coherent sheet (Figs. 3,4) that is wider as

air velocity increases; thus the overlapping of two contiguous sheets is more

pronounced and the probability of drop collisions increases as well. The mean

drop size is also important for the amount of drifting.

These results suggest some concluding remarks about the behaviour of these

nozzles. The best performing nozzle should give the higher wetting efficiency

with the lowest standard deviation, and a relatively low Sauter mean diameter

in order to improve the heat transfer efficiency in the spray zone (3), and

these characteristics should be relatively insensitive to changes in operating

conditions. Among the four nozzles employed in this work, nozzle D seems to

present the most satisfactory behaviour.

As regards the air-side pressure drops, the calculated values of N/Ls are very

similar (-0.23), since the mean liquid specific flowrate and terminal velocity

of the drops are the same for the four nozzle ensembles.

TABLE III

Sauter mean diameter (mm) of the drops produced by the nozzles:

effect of position and operative conditions

WL(m3/h! 160 160 160

Va(m/s)! 0 1 2.5

1 NOZZLE D3 2 D32 1D32 D32 D32 ! D3 2 ! D3 2 1 D3 2 D3 2 D3 2 !D3 2 D32 ; Pi 1 P2 ! P3 mean Pi ! P2 ! P3 1

1 1 mean Pi P2 ! P3 mean!

! A . 8.9 6.9! 5.3 6.93 9.1 7.2! 7.1j 7.43 9.2 6.61 6.7 7.01!

! B 7.9 3.31 6.4 5.19 8.0 5.3! 7.01 6.38 7.2 7.01 7.1 7.06!

! c 4.7 4.0! 6.0 4.89 4.6 5 • 3 1 5 • 9

1

5.43 4.3, 3.91 5.9 4.75!

! D 3.7 2.7! 3.9 3.31 , 4.0 3 • 2 1 3 • 3

1

3.35 4.7 3.713.75 3.86!

794
CONCLUSIONS

Careful experimental characterization of water distribution nozzles is

needed in order to evaluate the performances of water distribution systems in

cooling towers. On the basis of the results, some performance parameters have

been defined and evaluated, which can be related to heat transfer and pressure

drops in the spray zone of a cooling tower.

Measurements in two configurations, pilot tower and small scale loop,

included the geometric characteristics of the jet umbrella by ensemble

photography, the wetted zone by measuring the specific flowrate, the drop-size

distribution and liquid concentration by high-speed photography. The results

show that correlations exist between the nozzle behaviour in single and pilot

tower configuration. The uniformity of water distribution in the pilot tower

is strongly related to the nozzle installation pattern and to the operative

conditions. Coalescence plays an important role on the drop size distribution

in the spray zone of the pilot-tower.

REFERENCES

1. Berman, L.D. , "Evaporative Cooling of Circulating Water", Pergamon Press

1961.

2. B.Bellagamba, G.Dinelli, F.Macchi, M.Ceccanti, C.Scali, S.Zanelli "A

study on the performance of evaporative cooling towers with natural

draft", 5*** International Association on Hydraulic Research Cooling

Tower Workshop, Monterey, Cal, (1986).

3. B.Bellagamba, G.Dinelli, F.Macchi "Physical and engineering aspects of a

cooling tower performance test programme" 5*** lAHR Coolin. Tower.

Workshop, Monterey (1986).

4. B.Bellagamba, G.Dinelli, L.Tognotti, S.Zanelli "Water distribution in

cooling towers: characterization of industrial spray nozzles" 6**> lAHR

Cooling Tower Workshop, Pisa, Italy, (1988).

5. F.Macchi, B.Bellagamba, S.Baldacci, L.Tognotti, 42"^ Associazione

Termotecnica Italiana National Congress, Genova , Edited by CLEUP,

Padova, Italy (1987).

6. Dombrowski,H. , Hopper, P. C. ; Chem.Eng.Sc. , 17, 1962,291-305

7. Reitz, R.D., Bracco,F.V.; Mechanism of atomization of a liquid jet,

Phys. Fluids, Vol.25, 1730-1742, (1982).

8. Mugele, R.A., Evans, H.D.; Ind.Eng.Chem. , 43 (1951)

9. Lowe,H.S., Christie, D.J. : "Heat Transfer and Pressure Drop Data on

Cooling Tower Packings, and Model Studies of the Resistence of National

Draught Towers to Air Flow" Conference on Heat Transfer, (1961)

795

ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 89

SUPPRESSION OF SMOKE PARTICLES USING

ELECTROSTATIC SPRAYING

C.N. Ahmad and W. Balachandran

Department of Electronic and Electrical Engineering

University of Surrey

Guilford, United Kingdom

ABSTRACT

Experimental investigation has been undertaken to study the efficiency of

precipitating smoke particles using electrically charged water spray . A

laser attenuation measurement system was employed to study the

precipitation rate of cigarette smoke particles using charged and uncharged

water spray . The negatively charged water spray cleared the test chamber

three times faster than the uncharged spray . Based on the test results , a

realistic time for complete clearance of the test chamber independent of

source was estimated to be 60 seconds .

INTRODUCTION

A spray of fine water droplets is a well known means of airborne dust

removal . Many of the commercially available fog devices for fugitive

emission control are not very efficient . One method of improving the

effectiveness of water spray is by applying electrical charge to the spray

that is opposite in polarity to the charge on the dust to be suppressed . It has been found that most industrial pollutants and naturally occurring

fugitive dust acquire an electrostatic charge , well above the Boltzman

equilibrium , as they are dispersed in the air . If this charge of airborne

material is exposed to an oppositely charged water spray , contact between

the particulate matter and the water droplets is enhanced . After contact

is made , the wetted particulate matter agglomerates rapidly and falls out

of the atmosphere . The use of technology to suppress smoke particulates in

controlling indoor air quality has not been studied extensively.

There are mainly four methods that are commonly used to control the

effects of indoor smoke contamination . They are ventilation , filtration,

electrostatic precipitation and odour masking . In our study electrostatic

precipitation technique is the only method that was considered

Suppression of smoke particles by electrostatic precipitation involves

drawing smoke and air into an ionisation chamber in which the particles

acquire an electric charge . The mixture of air and smoke particles passes

through two or more plates between which electrical fields are maintained . The charged particles accelerate in the field and are deposited onto the

797
plates . Such devices can be even more effective by incorporating an

adsorptive filter , such as activated charcoal . A simpler device than an

electrostatic precipitator is an air ioniser , which generates an abundance

of small ions , which due to its high mobility , bombards onto the smoke

particles and accelerates their precipitation . Although there is no doubt

that these devices are efficient in removing particulate matter , several

studies have failed to show that the devices do alter the concentration of

the gas phase contciminants and therefore have limited value in removing the

odour . There is at present no alternative technology available for

eliminating smoke or controlling it to acceptable limits , other than the

ventilation which must be adjusted to cope with any practical situation . In many instances the volumes of ventilation required to control odour and

irritant effects are unrealistically high , it is therefore necessary to

investigate alternative methods of suppressing both smoke particulate

matter as well as the odour . In the work reported in the literature

regarding the use of charge and water sprays for the suppression of smoke

particles , there have been some limitations in achieving efficient

precipitation due to low specific charge levels and lack of understanding

of the optimxim droplet size required for such applications . It is believed

that the effectiveness of these charged sprays can be improved by

controlling the spray drop size distribution and specific charge . In this paper a brief theoretical background and some results of the

preliminary empirical studies carried out will be presented .

THEORETICAL BACKGROUND.

The collection of an aerosol particle by charged droplet is the result of a

number of simultaneous mechanisms of interaction between them such as

inertial impaction , direct interception , Brownian diffusion and

electrostatic dif fusiophoretic and thermophoretic forces . The particle

collection efficiency of uncharged water sprays where inertial impaction is

the major collection mechanism is given by Cheng (1) as

3 L E=1-EXI1--*— (1)

where E is the overall collection efficiency , T| single droplet collection

efficiency , volumetric flowrate of water, volumetric flowrate of

air, L characteristic length of the total capture process , d mean droplet

diameter

.

When the droplets and particles are electrically charged, T| is given

by Prem et . al . (4)

Cqpq^

n -2——2— (2)

en e^r^r^^U^

where qp is the charge on the particle., q^^ the charge on the droplet.

C the Cunningham slip correction factor, Eq permitivity of free space,

H viscosity of the liquid, droplet free stream velocity

For a given particle size distribution , electrostatic forces are

proportional to the magnitude of the charge and inversely proportional to

the droplet ' s size and its free stream velocity. When charged droplets are

sprayed into the air they will evaporate unless the air is saturated with

water vapour. The droplet lifetime determines the effective contact time

798
between the droplet and particles and thus strongly influences the overall

particle control efficiency of the system . The lifetime of water droplets

depends upon the temperature and relative humidity of the medium into which

it is introduced . To obtain the best collection efficiency , the droplet

must be small enough to provide both an adequate spray rate per volume of

gas treated and sufficient contact time yet large enough so as not to

evaporate too quickly . Equations (1) and (2) show that the addition of electrical charge onto

the droplets yields values of E which are 5 to 10 times higher in certain

ranges for uncharged spray , the overall collecting efficiency of an

operating system may not be that high. Pilat (3) observed that an increase

in collection efficiencies of about 15% for 1 (Xm particles to over 45% for

0.3 [MCI particles . Hoenig (2) reported values of 10 - 40% with charged

water droplets under controlled experimental conditions. He also

demonstrated that the greatest effectiveness is obtained when water

droplets are of a size similar to that of the particulates .

EXPERIMENTAL ARRANGEMENT

.

The experimental setup used in this study consisted of the three parts as

(a) the chamber (b) smoke generator (c) spray nozzle.

The Chambar

A cubical chamber (0.5 * 0.5 * 0.5 m ) was constructed of 5 mm thick

perspex (as shown in fig 1) . Two optical windows are incorporated on the

opposite sides of the chamber . A smoke generator and the spray nozzle are

mounted on top of the chamber . A recirculator fan mounted on the side near

the top provides a link between the top and bottom of the chamber to

achieve uniform density of smoke within the chamber . The chamber is fixed

on an optical bench along with a He-Ne (2 milliwatt) laser and a polariser

on one side and a pin photodiode detector on the other side of the chamber

A drain plug is also provided at the bottom of the chamber to empty the

contents after the experiment . The floor of the chamber is covered with

fine wire wool to eliminate the spray bounce back which tends to interfere

with the laser beam and give erroneous measurements during the experiment . Two cases were considered in the series of experiments to be described here

One with insulating walls and the other with conducting walls to the

chamber . The latter was achieved by sticking aluminium foil on all sides

inside the chamber. This case is more representative of the real situation

since in a room the walls and the floor are always at or near ground

potential compared to the charged spray and ions . The photodiode detector

output was connected to a Keithley (610C) electrometer . The analogue

output of the electrometer was fed to a chart recorder via a 10 K resister.

After initial tests the charged spray in the chamber fouled the optical

windows inspite of using anti-fog spray, thus making measurements impossible

The problem was overcome by mounting two circular metallic ring electrodes

around the windows . These rings when maintained at potential of the same

polarity as the spray, repelled the charged spray away from the window.

The Smoke Generator . The smoke generator consists of a bell shaped glass chamber with one inlet

and two outlets. A cigarette is held on a flat circular glass plate with a

hole in the centre . The main glass tube sits on this plate and a good seal

is obtained at the bottom . The air blown through the inlet is forced

through the cigarette, generating smoke into the test chamber by

controlling the outlets with an on-off solenoid to simulate human 2 -second

duration puffs.

799
Svrav Nozzle

The spray nozzle used in these experiments is the standard spraying systems

air atomising (1/4 JAU) nozzle . The metallic air cap has been replaced

by a PVC air cap incorporating an embedded ring electrode to charge the

spray by induction . When the nozzle is in operation a sheet of high

velocity air flows between the liquid film and the electrode so that the

gap between the electrode could be kept very small (1mm) without any

electrical breakdown . Moreover since this gap is small , hence high

charging efficiency is obtained with relatively low applied potential

(<1KV) . Both air supplies to the nozzle as well as to the piston were

controlled by two on - off solenoid valves so that the duration of the

spraying could be controlled precisely. The spray nozzle is characterised

in terms of the droplet size distribution and the charging efficiency . In

these tests the nozzle was operated at a liquid flow rate of 75ml/min. and

air pressure of 20 and 40 psi to the piston and to the nozzle respectively

giving a typical droplet size VMD (volume median diameter) of 20 JJm . The

spray size distribution was measured using a Malvern Particle Size Analyser

and typical results are shown in fig 2 . The corresponding charge to mass

ratio was found to be 2mC/Kg . Fig 3 shows the specific charge values (q/m)

obtained for different operating conditions of the nozzle .

Fig. 2 Droplet size distribution

11.34

3C

V

a

1u

n

c

5.77

6 • 66

3.6 44.8 54.3

BiaMeter uM

.8

Lcnsth

Surface

Volune

Span

18 '/. 98 •/

K««[ian

Hcdian

Mrdian

Median

Point

Point

9.8

12.

e

17.1

20.1

1.4

8.5

35.8

uM

vM

uH

UH

800
Fig. 3

3

0> M

1-

Charge to mass ratio Measurements

for 1/4 JAU nozzle.

Liquid flowrate 75 ml/min.

Cylinder pressure 20 psi

T T

0 250 500

APPLIED POTENTIAL (volts)

1000

20

25

30

35

40

45

psi

psi

psi

psi

psi

psi

LASER MEASUREMENTS

Before each set of measurements the chamber , laser and the detection

system was properly aligned to get a maximum signal output . The chamber

was then filled with smoke through the smoke generator . In these

experiments only 1/4 of the Marlboro cigarette was used each time, to

simulate a realistic situation . The recirculating fan was kept on until a

uniform density of smoke was achieved throughout the chamber and then it

was switched off . The smoke was then exposed to an appropriate treatment

i.e. exposed to the ioniser, (positive or negative) charged spray or

uncharged spray . The treatment duration was fixed throughout these

experiments , to 3 0 seconds only and the signal output was recorded for

upto six minutes . Once the measurements were completed the chamber was

cleared of the residual smoke by starting the recirculating fan and by

disconnecting the bottom end of the link.

RESULTS AND DISCUSSIONS

From the experimental recordings , the signal was monitored upto 2 minutes

from the start of the treatment . In each case the signal was normalised to

the level of signal at the start of the treatment and then plotted against

time to get the rate of smoke clearance from the chamber . These results

are shown in Figures 4 and 5 for conducting and non - conducting walls of

the chamber respectively . From these results , the times for complete

clearance of the chamber could be easily projected and are given in table 1

The laser measurement system used in these experiments relies on the level

of attenuation of the beam intensity due to light scattering . When the

laser beam propagates through a cloud of smoke particles , light is

scattered by the particles , and therefore only the transmitted beam will

reach the detector . The intensity I of a plane electromagnetic wave of

wavelength I and intensity 1^ after traversing a distance L through the

smoke particles is given by Bouguer ' s ( Lambert -Beer ' s ) Law , namely

l(A,) = Io(X)ecp[-T(A.)] (3)

where the optical depth xiX) is given by '^^^^ = 4 Pext^^^^^

801
= L'\Q (x,m) 7cr^n(r)drdl (4)

where ^^^^

(cm""'') is the volume extinction coefficient, Qext (x,m) the

extinction efficiency, ^~ 27Cr/A, ^-^e size parameter, m = m' - i m" the

aerosol complex refractive index, r-j^ and r2 are the upper and lower

limits of the aerosol radii.

If we assume that n ( r ) remains constant along the path of the laser

beam, then

r x(X) = nhl 'q (x,m) r^n(r)dr. (5)

•'r ext

1 Hence the transmission coefficient can be written as :

T =

I, a)

= exp - 4-

f 'r^n (r)Q ^ (x,m)

'r ext

f 3

J r n(r)dr

(6)

where C is the volume concentration of the aerosol.

In practice , due to the detector's finite field of view , some

forward scattered light invariably enters into the detector's view cone

along with the direct beam . This effect can be minimised using a pin hole

appropriately placed in front of the detector and further correction

factors can be added to equation (6) . However , for this study , such

effects are assumed to be negligible .For a mono dispersed aerosol of mean

radius r:

T= exp - V Qext(^"^)

LU (7)

The above equation indicates that the transmission coefficient is a

function of volume concentration and the size distribution of the aerosol

cloud. In our experiments , as indicated initially the attenuation of the

laser beam was due to the smoke particles . When the spray was introduced , the attenuation further increased as expected from equation (7) . This

lasted as long as the spray was in the field of view and therefore , when

the spraying ceased , it was possible to then relate the rate of clearance

of the chamber to the increase in detected signal . This way it was

possible to quantify and compare the operation of the system under

different conditions . These tests were repeated several times and the

repeatability was good .With ionisers , additional attenuation of the laser

light was not observed, because the small ions are submicron in size and

hence do not have any influence on the measurement system . The results shown in table 1 clearly indicate that the suppression of

smoke particles with the conducting chamber , in general is better than the

non - conducting chamber . This is not surprising , as the conducting walls

are at ground potential , therefore the precipitating charge agglomerates

are strongly attracted to the ground surfaces . In the case of non - conducting chamber , based on the level and nature of surface charge on the

surface of the walls , the precipitation rate varied . Although the net

charge on smoke particles appears to be neutral (equal amount of positive

and negative ) , the projected results for positive and negative ioniser

are significantly different . The mechanisms involved in suppressing smoke

particles appears to be complex and has not been studied in depth in this

802
preliminary investigation . However , it is clear in both cases that the

presence of positive ions clears the chamber much faster than either

negative ions or negatively charged water spray . There is no doubt that

the negatively charged water spray clears the chamber three times fast^er

than the uncharged spray . With the conducting chamber , the clearance rate

of smoke particles with negatively charged spray is 1.5 times faster than

the presence of negative ions . These results clearly indicate that

negatively charged water spray significantly increases the suppression of

smoke particles . Another observation that was made was that the use of

positive spray did not significantly improve the precipitation rate

compared to the negative spray . This obviously contradicts the

observations made with positive ions , and therefore needs further

investigation .

Table 1 Projected Time for Complete Clearance of the Chzunber.

Source Non-conducting

Chamber

Conducting

Ch«unber

Negative Ioniser

Positive Ioniser

Uncharged Spray

Charged Spray (-ve)

6 . 5 minutes

1.4 minutes

6.4 minutes

2 . 1 minutes

2.2 minutes

0.61 minutes

4.5 minutes

1.5 minutes

803
Fig. 5 Rate of smoke clearance In

non-conducting cheunber.

i.n

0 12 3

TIME (MINUTES)

CONCLUSION

Electrically charged water spray can be used successfully to suppress smoke

particles . Typical specific charge necessary for efficient precipitation

is about 2mC/Kg . The rate of precipitation of smoke particles by negative

ions is about 50% less than charged water spray of VMD 20|Jm. The realistic

time of complete clearance of an enclosed room is estimated to be three

times faster than an uncharged water spray.

Acknowledgement

.

The authors wish to acknowledge S.C Johnson & son Inc. for the financial

support provided for this study.

REFERENCES

1. CHENG L ."Collection of airborne Dust by water sprays " Ind. Eng.

Chem. Process Des . Development , 12 pp 221-225 (1973).

2 HOENIG S A , "Control of Industrial fugitive dust with

electrostatically charged fog devices "

, A Status report , Conf

.

Record , IEEE - IAS , pp 982- 986 (1982).

3. PILAT M J , "Collection of Aerosol particles by electrostatic droplet

spray scrubbers "JAPCA ,25 pp 176-178.(1975).

4 PREM A and PILAT M J, " Calculated particle collection efficiencies

by single droplets considering inertial impaction, Brownian diffusion

and electrostatics " 12 , pp 1981-1990 (1978).

804
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 90

ATOMIZATION AND DISPERSION OF QUASI-MONODISPERSE

ELECTROSTATIC SPRAYS OF HEPTANE

A. Gomez and K. Tang

Department of Mechanical Engineering

Yale University

New Haven, CT, U.S.A.

ABSTRACT

Detailed measurements by digital imaging techniques and phase Doppler

anemometry were performed to study the atomization and evolution of liquid

fuel electrostatic sprays. The electrospray was chosen to assess its potential

as a research tool for laboratory scale combustion experiments under

conditions much better defined than in practical sprays. Experiments were

carried out on heptane doped with an antistatic additive to enhance its

electric conductivity. Stable electrosprays with droplets characterized by a

remarkably quasi-monodisperse size distribution were generated over a wide

range of experimental conditions. Average droplet diameters anywhere from 1 to

over a 100 |Xm were obtained depending on liquid flow rate and electric field

intensity. A phenomenological picture of the atomization and dispersion

processes leading to these results is formulated. Images of the jet break-up

showed that, when narrow size distributions are generated, the spray is formed

from a thin, stable jet that breaks up into a stream of droplets. The

resulting size distribution is typically bimodal with two peaks corresponding

to primary droplets and satellite droplets, respectively. The smaller

droplets at the breakup are electrostatically/inertially separated from the

larger ones. Therefore, farther downstream, the size distributions is locally

quasi-monodisperse. Spray dispersion is determined by both coulombic repulsion

between droplets and diverging lines of field. No Coulomb explosion of the

droplets was observed in the region where the electrospray is formed.

INTRODUCTION

Fuel atomization processes typically involve a two-phase flow of a

polydisperse distribution of droplets in a turbulent gaseous environment : a

formidable challenge to both experimentalists and theoreticians. It would then

be extremely useful in sorting out the roles played by different aspects of

the process to use an experimental arrangement that allow a systematic study

of spray evolution and burning in configurations of gradually increasing

levels of complexity, in both laminar and fully turbulent flows. An

electrostatic spray of charged droplets seems to lend itself admirably to

combustion experiments under well-defined conditions. In its simplest

implementation, which is also the one we adopted in this study, a liquid is

fed into a small metal tube maintained at several kilovolts relative to a

ground electrode few centimeters away. If the liquid has sufficient electric

805
conductivity, its meniscus at the outlet of the capillary takes a conical

shape under the action of the electric field, with a thin jet emerging from

the cone tip. This jet breaks up farther downstream into a fine spray of

charged droplets. Distinctive advantages of the electrospray over alternative

atomization techniques are: the self -dispersion property of the spray due to

coulombic repulsion; the absence of droplet coalescence; and the opportunity

of manipulating the trajectories of charged droplets by suitable disposition

of electrostatic fields. Furthermore, in contrast to aerodynamic atomization

processes, the electrospray relies on charging to atomize the liquid fuel, so

that atomization and gas flow processes are relatively uncoupled. This last

feature provides an additional degree of freedom and may in principle allow a

systematic study of the spray evolution and burning starting from laminar

sprays to fully turbulent ones.

A well-defined spray combustion experiment requires control on droplet

size distribution which should ideally be monodisperse to simplify the

experiment interpretation. The electrospray appears also promising in this

respect, on the basis of fragmentary data reported in the literature that

indicated the generation of narrow droplet size distribution under disparate

experimental conditions. Almost forty years ago some observations of High

Order Tyndall Spectra were reported in the scattering of white light by

"smokes" produced in electric dispersion of liquids (1,2). This optical

phenomenon is indicative of droplet formation of narrow size distribution in

the size range 0.1-2 (im (3) . At the other end of the size spectrum. Thong and

Weinberg (4) showed that a kerosene electrospray, in a very narrow range of

operating conditions, could generate monodisperse droplets varying in size

from 50 to 120 |im. These authors also demonstrated the feasibility of burning

the electrostatically atomized fuel.

The application of liquid fuel electrostatic spraying to more practical

situations, involving high flow rates, has been studied by Kelly (5) at

atmospheric pressure and, more recently, by Kwack et al . (6) at Diesel engine

pressures. Spraying involves a charge injection technique based on the use of

a patented "triode" (7). The distinctive advantages of handling large flow

rates is unfortunately partially offset by the polydispersion of the droplet

size distribution. We chose the simpler configuration described above,

which, even though limited to relatively small flow rates, appears more

promising for the controllability of spray combustion experiments.

Detailed studies, both experimental and theoretical, of the

electrohydrodynamic phenomena leading to the formation of the liquid conical

meniscus have been reported (8-12) . With regard to the details of the jet and

its break up, the most relevant contributions, from our perspective, are the

work of Thong and Weinberg (4), briefly discussed above, and the study of

Mutoh et al. (13) who investigated the break up of charged liquid jets of

xylene. Observations on the spray formation, including both jet break up and

droplet dispersion, are more difficult and information available is rather

scanty; it is therefore in this area that we have concentrated the initial

efforts discussed here. This experimental study focuses primarily on the

characterization of the atomization process and subsequent spray evolution in

heptane electrosprays , with the ultimate goal of identifying the mechanism

governing the generation of narrowly disperse particles over a wide size

range. A preliminary account of part of this work was presented elsewhere

(14) .

EXPERIMENTAL

Heptane, whose flow rate was controlled and metered by a syringe pump,

was fed into a stainless steel metal capillary (I.D.sO.20 mm) maintained at a

fixed voltage within the 3-5 kV range relative to a ground electrode

positioned 3 cm away. The spray break-up was observed by using an optical

806
system consisting of: a (25 ns) pulsed flashlamp, as illumination source,

which was focused on the electrospray by a quartz lens; a stereozoom

microscope, positioned in a shadowgraph configuration; and a digital camera

mounted on the microscope. The camera signal was recorded by a VCR and

digitized, off-line, by a frame grabber installed on a personal computer. The

quantitative use of this optical arrangement was restricted to experimental

conditions yielding droplet diameters larger than 10 |im because of limitations

in the resolution of the microscope. The overall magnification on the

computer monitor was =800 X. Detailed single point measurements taken

throughout the spray by commercial Phase Doppler Anemometers (PDA) allowed

simultaneous inference of droplet size and axial velocity distribution, as

well as size-velocity correlation. In a typical experiment, for a given flow

rate, the voltage was operated in a range corresponding to the generation of a

narrow size distribution, as monitored on line by the PDA along the spray

axis. Heptane was doped with an antistatic additive (Stadus 450, Dupont; 0.3%

by weight) to enhance its electric conductivity. The total current collected

by the ground electrode was measured by an oscilloscope. All experiments were

performed at ambient temperature and atmospheric pressure.

RESULTS AND DISCUSSION

Stable jets with droplets characterized by a locally quasi -monodisperse

size distribution were generated over a broad range of experimental

conditions. Average droplet diameters anywhere from 1 to over a 10 0 |lm were

obtained depending on: liquid flow rate; shape and intensity of the electric

field; and, although not investigated in the present study, liquid physical

properties. Four sample distributions are shown in Fig. 1, corresponding to

the following flow rates: 0.0067 cc/min (Fig. la), 0.017 cc/min (Fig. lb),

0.17 cc/min (Fig. Ic) and 0.34 cc/min (Fig. Id), respectively. Size

distributions appear roughly symmetric, with a remarkably small ratio of

standard deviation over mean diameter of 0.10, 0.07, 0.04 and 0.06,

respectively. It should be particularly emphasized that narrowly dispersed

micron-size droplets where generated from a relatively large tube, thereby

circumventing all clogging problems that characteristically plague

alternative, monodisperse aerosol generators.

Figure 2 shows the dependence of droplet diameter on applied voltage,

for selected flow rates; all measurements were taken in conditions yielding

quasi-monodisperse size distributions. We observe that: i) droplet size is

primarily dependent on liquid flow rate and increases monotonically with it;

ii) at fixed flow rate, an increase in voltage, on the other hand, causes a

decrease in size anywhere between 10 and 30%; and iii) for each flow rate,

there is a fairly wide (1, 1.5 kV) range of voltages over which quasimonodispersity prevails.

In an effort to investigate the mechanism determining the quasimonodispersity of the generated sprays, details of the dynamics of the jet

break-up were examined in imaging experiments. Typically, a sequence of events

shown in Figure 3, in this particular case obtained for a 0.17 cc/min

heptane spray, was observed:

1) the liquid exits the cone formed at the outlet of the charged capillary as

a thin, stable, thread that persists for a short distance, here of roughly 2

mm (Fig . 3 a)

;

2) this ligament then breaks up into droplets, which typically, as in the

present case, exhibit a bimodal distribution (Fig . 3 b)

;

3) the smaller (satellite) droplets at the break up are rapidly displaced

radially with respect to the larger, primary particles generated (Fig. 3 c)

;

4) the primary droplets left behind are quasi-monodisperse. They move

downstream, are gradually displaced radially and leave a "wavy" signature

under instantaneous illumination (Fig. 3 d)

.

807
3.

. S2E>g3-

.00 .00

Concentration

Uol. fraction

Vol. flux

2.49E+04 /CO

3.33E-05 X 6.37E-05 CC/CM2/S

Mean Diane tep (D10> 2.76 Microns

Area Mean (D20) 2. 85 Microns

Uolune Mean (D30> 2.93 Microns

Sauter Mean <D32> 3.10 Microns

a)

. 10E+04-

3>

0

0\ .51E+03

.00

Concentration

Uol. fraction

Uol. flux

3.08E+04 /cc

1.88E-03 •/ 1.O5E-02 CC/CH2/S

Mean Diane ter <D10> 10.6 Microns

Area Mean (D20> 10.6 Microns UoIuMe Mean <D30> 10.6 Microns

Sauter Mean <D32) 10.7 Microns b)

.62Et03-^

a.\

0

u\ .31E+03

.00

Concentration

Uol. fraction

Uol. flux

1.72E+04 /cc

4.32E-02 •/. ,34 cc/ch2/s

Mean Dianeter <D10) 36.3 Microns Area Mean (D20) 36.3 Microns

UoluMe Mean <D30) 36.4 Microns

Sauter Mean (D32) 36.4 Microns

C)

.16E+03

\ 79.

.00

20. 52. 85. Dianeter tniorons]

Concentration

Uol. fraction

Uol. flux

3.61E+03 /cc

5.15E-02 •/ .27 cc/cM*/s

Mean Dianeter <D10) 64.8 Microns Area Mean <D20) 64.9 Microns

Uolune Mean <D30) 65.0 Microns

Sauter Mean (D32) 65.2 Microns

d)

Fig. 1: Droplet size distribution for sprays with different mean diameters.

It can be clearly seen from Fig. 3b that, before break-up occurs, varicose

waves propagate along the ligament. The ratio of primary droplet to ligament

diameter was measured at about 1.9, in good agreement with the classic theory

of Rayleigh on the stability of capillary jets (15). Extension of Rayleigh

theory to account for electrification effects, worked out by Schneider (16)

and corrected to second order terms by Neukermans (17), showed that, even at

relatively large charge levels, the predictions of Rayleigh theory, at least

in so far as dominant propagating wavelength is concerned, are not

significantly altered. Our findings are consistent with the theoretical

predictions. Experimental confirmation was also provided by a study on the

break up of charged xylene jets (13). We conclude that electrification is

controlling droplet size through electrohydrodynamic processes in the cone

determining the formation and diameter of the jet.

The satellite droplets generated at break up are first displaced

radially by a small disturbance and the Coulomb interaction with neighboring

droplets; subsequently, they tend to follow the diverging lines of field

808
2.5 3.0 3.5 4.0 4.5

Voltage (KV)

Fig. 2: Average droplet diameter versus applied voltage at different flow

rates

.

because of both relatively small

inertia and, possibly, higher charge

to mass (18). Thus, they are removed

from the core of the spray by

e lectros tat ic/ inert ial separat ion

.

The establishment of the sinuous

pattern in the array of primary

droplets in Fig. 3d can also be

attributed to the coulombic

interaction among droplets: it can be

readily shown that if a droplet in a

linear array is displaced radially by

a small disturbance, the vector

summation of the resulting

electrostatic repulsive forces

inevitably leads to the observed wavy

pattern. Notice that this is not to be

confused with a droplet trajectory.

Figure 4 shows the variation

of the droplet average diameter as a

function of radial coordinate, at

fixed axial position, for the same

0.17 cc/min heptane spray.

Superimposed on the plot are

histograms of the size distribution at

Fig. 3: Jet breakup images obtained

at increasing distance from the

capillary (a—>d) .

809
0 12 3 4

Radicil Position (rrm)

Fig. 4: Average droplet diameter versus radial position at a fixed axial

coordinate

.

three radial positions. It is evident that the average droplet size

monotonically decreases as a function of the radial position. This is not a

consequence of different evaporation rates, which have negligible effects

under the present experimental conditions, but it is consistent with the

interpretation of Fig. 3c in terms of electrostatic/inertial separation. At

intermediate radial positions, we observe the existence of a bimodal size

distribution in correspondence with the region between the spray core,

characterized by the primary droplets, and the outer shroud of satellite

droplets formed at the jet break up.

We also looked for evidence of Coulomb explosion which occurs near the

so-called Rayleigh limit, when electrostatic repulsive forces overcome

surface tension forces and the droplet disrupt into smaller droplets. This

process inevitably occurs at some point during the droplet evaporation, since

there is experimental evidence that charge remains attached to the droplet

while it is evaporating (19) . The question is if it determines or, at least,

contributes to, the droplet dispersion and therefore the spray formation. A

detailed examination of the shape and evolution of the size distribution along

and around the axis of the 0.17 cc/min spray provided no evidence of the

occurrence of droplet disruption in the probed region.

We can now attempt to generalize these observations to other flow rates

or experimental conditions. At all flow rates, the break-up region showed a

stable ligament, as in Fig. 3a, which would disrupt into a stream of droplets

leaving the wavy "signature" as in Fig. 3d. However, for the smallest flow

rates, and correspondingly smallest average droplet diameter, no observation

of satellite droplets, nor bimodal size distribution, was possible, possibly

because the limits of resolution of both imaging system and PDA were reached.

Figure 5 shows the average charge to mass ratio (C/M) versus the applied

voltage for different flow rates for the same conditions as in Fig. 2; the

C/M's are normalized with respect to the Rayleigh limit calculated for the

810
2.5 3.0 3.5 4.0 4.5

V ( KV )

Fig. 5: Average charge to mass ratio, normalized with respect to the

Rayleigh limit, versus applied voltage.

average droplet size of each spray. Since the normalized C/M's of the other

two sprays tested are lower than the values of the 0.17 cc/min spray for which

Coulomb explosion was ruled out, we expect that no droplet disruption is

occurring in any of the sprays examined.

Our observations indicate that only when a stable, thin jet is formed

at the outlet of the metal capillary quasi-monodisperse size distribution are

generated. For a given flow rate, if the voltage was lowered below a lower

threshold the jet would be unsteady and pulsating; if, on the other hand, it

was raised above an upper threshold, whip-like lateral instabilities would

ensue, as also theoretically predicted in similar geometries and at high

charging levels (20) and experimentally verified by others (9,13). In this

case the break-up mechanism appears qualitatively unchanged in the imaging

experiments; however, the whipping of the ligament causes unsteadiness in the

local electric field and mixing of the droplets generated. Consequently, the

droplets are no longer spatially separated by electrostatic-inertial effects

and locally the size distribution is broadened to encompass the whole size

spectrum generated at break-up. Under this condition, monodispersity can

still prevail if utmost care is taken towards the elimination of satellite

droplets at break-up, as Thong and Weinberg (4) reported.

CONCLUSIONS

1) Stable electrosprays with droplets characterized by a locally quasimonodisperse size distribution can be generated over a wide range of

experimental conditions. Average droplet diameters anywhere from 1 to over a

100 lim were produced depending on liquid flow rate, shape and intensity of the

electric field. Size distribution is typically symmetric, with a ratio of

standard deviation over average diameter of approximately 0.1;

2) Liquid flow rate is the primary variable influencing droplet size, with

the average diameter increasing monotonically with it; at fixed flow rate, an

increase in voltage, on the other hand, causes a decrease in size anywhere

between 10 and 30%. For each flow rate, there is a fairly wide (1, 1.5 kV)

voltage range over which quasi-monodispersity prevails;

811
3) Images of the jet break-up, showed that, when narrow size distributions

are generated, the spray is formed from a thin jet that breaks up into

droplets with a bimodal distribution. After break-up, electrostatic/inertia

separation of the smaller droplets from the larger ones yields locally

quasi-monodisperse droplet size distributions;

4) No Coulomb explosion was observed in the region where the electrospray

is formed. Coulombic repulsion between droplets and diverging line of fields

cause the spray dispersion.

ACKNOWLEDGMENTS

We are indebted to Profs. J. B. Fenn, J. Fernandez de la Mora and D.E. Rosner

for many stimulating discussions and helpful comments. Thanks are also due

to both Dantec Electronics and Aerometrics for allowing us to use their

particle measurement systems in extensive laboratory demonstrations, during

which some of the data of this work was gathered. This research was in part

supported by an NSF Equipment Grant #CBT-8806608 and a NASA Grant #NAG3-1259.

REFERENCES

1. Vonnegut, B. and Neubauer, R.L., J. Colloid Sci . , 7, p. 616 (1952).

2. Vonnegut, B. and Neubauer, R.L., J. Colloid. Sci., 8, p. 551 (1953).

3 van de Hulst, Light Scattering by Small Particles, Wiley, 1957.

4. Thong, K.C. and Weinberg, F.J., Proc. Roy. Soc. Lond. A, 324, p. 201

(1971) . 5. Kelly, A.J., The Electrostatic Atimization of Hydrocarbons, Second

International Conference on Liquid Atomization and Spray Systems, ICLASS82, Madison, Wisconsin (1982).

6. Kwack, E. Y., Back, L.H. and Bankston, CP., Electrostatic Dispersion of

Diesel Fuel Jets at High Back Pressure, Fall Technical Meeting, Western

States Section, The Combustion Institute, Dana Point, Ca, 1988.

7. Kelly, A.J., Electrostatic Atomization Device, U.S. Patent No, 4, 255,

777, March 10 1981.

8. Zeleny, J., Phys . Rev., 10, p.l (1917).

9. Taylor, G.I., Proc. Roy Soc. Lond. A, 313, p. 453 (1969).

10. Hayati, I., Bailey, A.I. and Tadros, Th. F., J. Colloid. Interface. Sci.,

117, p. 205 (1987a).

11. Hayati, I., Bailey, A.I. and Tadros, Th. F., J. Colloid. Interface. Sci.,

117, p. 222 (1987b).

12. Turnbull, R.J., IEEE Trans. Industr. Appl . , IA-25, p. 699 (1989).

13. Mutoh, M., Kaieda, S., Kamimura, K., J. Appl. Phys., 50, p. 3174 (1979).

14. Gomez, A. and Tang, K. ," Characterization of Liquid Fuel Electrostatic

Sprays", Paper presented at the Western States Section/The Combustion

Institute 1990 Fall Meeting, La Jolla CA.

15. Rayleigh, Lord, Proc. Roy. Soc, 29, p. 71 (1879).

16. Schneider, J.M., Lindblad, N.R., Hendricks, CD. and Crowley, J.M., J.

Appl. Phys., 38, p. 2599 (1967).

17. Neukermans A., J. Appl. Phys., 44, p. 47 69 (1973).

18. Gomez, A. and Tang, K., "Characterization of a High Charge Density

Electrospray of Methanol"", Proceedings of the Fifth International

Conference on Liquid Atomization and Spray Systems, ICLASS-91,

Gaithersburg, MD, U.S.A., 1991.

19. Abbas, M.A. and Latham, J., J. Fluid Mech. , 30, p. 663 (1967).

20. Melcher J., Field Coupled Surface Waves, Chap. 6, M.I.T. Press, Cambridge,

Mass. (1963) .

812
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 91

EVALUATING THE INFLUENCE OF FUEL PREPARATION

ON THE PERFORMANCE OF A SPARK-IGNITED ENGINE

C.H. Daniels, L.W. Evers and Z. Han

Michigan Technological University

Houghton, MI, U.S.A.

ABSTRACT

An experimental technique has been developed for evaluating

the influence of mixture preparation on the performance of a

spark-ignited engine. The preparation components studied were

fuel vapor, droplets, and liquid streams. These components were

created and physically measured in a specially designed mixture

preparation channel. Their effects on in-cylinder pressure

performance and exhaust gas concentrations were investigated for a

1.9 L Ford engine operating at a speed of 1500 rpm and an absolute

manifold pressure of 49.1 kPa (14.5 in Hg)

.

Results from this investigation showed that the performance

of the engine was, in general, diminished by increasing the amount

of fuel in liquid stream form. In addition, an "equivalent"

mixture resulting from the engine's conventional port injection

system was identified in relation to engine performance.

INTRODUCTION

The degree of fuel atomization and vaporization in the intake

port of a spark-ignited engine is generally thought to influence

engine performance. Many investigators have performed tests with

a variety of engines in the hopes of determining the effects of

air/fuel mixture preparation on engine performance [1-4] . In

almost all cases, however, no physical measurements of the degree

of fuel vaporization and atomization were made. Because of this,

the results of these studies tended to be qualitative in nature.

In general, the results of these studies showed an extension of

the lean limit, a reduction in hydrocarbon and carbon monoxide

emissions, an increase in nitrogen oxide emissions, and greater

stability with increased mixture homogeneity.

The purpose of our work is first to develop a testing

procedure for evaluating quantitatively the influence of mixture

preparation in the intake port on engine performance. Second, to

show how the results obtained can be used to evaluate the

operation of a conventional port injector.

813
The fuel in the intake port appears in three distinct

physical forms. These are vapor, droplets, and streams. Figure 1

shows these fuel forms produced from the spray of a port injector.

Fuel Injector Intake Valve

Figure 1 : Mixture Components in the Intake Port

Droplets are formed from the atomization of the fuel spray.

Some of the droplets, especially the larger ones, impinge

on the passage walls and form liquid streams. Vaporization of the

droplets and streams occurs throughout the intake process and is

limited by the fuel's vapor pressure, time, and surface

temperatures. It is these components and their effects on engine

performance that is investigated in this study.

The engine performance parameters are obtained from

in-cylinder pressure and exhaust gas measurements. The

in-cylinder pressure parameters studied are the gross indicated

mean effective pressure (IMEP) , the coefficient of variation in

the IMEP (GOV), and the 10-90% burn time (BT) . The exhaust gas

parameters studied are the concentrations of carbon monoxide (CO)

,

carbon dioxide (CO2) , oxygen (O2) , hydrocarbons (HC) , nitric oxide

(NO) , and nitrogen oxides (NOx)

.

TEST EQUIPMENT

Engine

The engine used is a Ford 1.9 L EFI HO. The EGR system of

the engine is removed. In addition, the intake manifold is

modified such that one cylinder (the test cylinder) of the engine

receives air and fuel via a special mixture preparation channel.

Fuel Preparation Channel

The mixture preparation channel, shown in Figure 2, contains

devices for the creation and measurement of the different mixture

preparation components: fuel vapor, droplets, and liquid streams.

814
Figure 2: Mixture Preparation Channel

Atomizer

.

A pneumatic atomizer introduces fuel into the

preparation channel. The atomizer produces droplets small enough

so that some of droplets can traverse the channel length and avoid

impingment on the channel walls.

Psychrometer

.

A psychrometer is used in the preparation

channel to determine mixture equilibrium. Mixture equilibrium

insures that the vaporization is completed. The "wet-bulb" probe

consists of a thermocouple with its junction wrapped in a gauze

wick. This probe is wetted by the droplets of the mixture. The

"dry-bulb" probe consists of a thermocouple with its junction

coated in a non-wetting, teflon wax. A temperature difference

indicated between these probes is the result of evaporative

cooling. Hence, equilibrium is demonstrated by the same

temperature being indicated by both probes.

Stream Separator

.

The stream separator is a device used to

measure the liquid streams. When the separator is in the "open"

position, the streams are diverted to a buret for measurement.

Droplet Smasher. The droplet smasher is a device consisting

of fine-grade steel wool. When it is placed in the channel, the

fuel droplets collect on it and are converted to liquid streams.

Droplet Deflector

.

The droplet deflector is another device

consisting of fine-grade steel wool. When it is placed in the

channel (of which it occupies part of the cross-section) , some of

the droplets are converted to liquid streams.

Laser Droplet- si zer

A Malvern 2600 Particle-sizer is used to measure droplet

sizes and distributions in the preparation channel. This

instrument uses the principle of light diffraction in its

determination of droplet sizes and distributions.

815
In-cylinder Pressure Measurement

The pressure in the test cylinder is measured through a

Kistler pressure transducer (type 6123) located in the head of the

engine. The signal from this transducer is recorded and analyzed

using Ford Motor Company's Combustion Pressure Analysis System

(COMPAS) . IMEP, COV, and BT are obtained directly from the COMPAS

program

.

Exhaust Gas Measurement

The exhaust gas is sampled from the exhaust pipe of the test

cylinder. CO and CO2 are measured by Beckman 315A infrared

analyzers. O2 is measured by a Beckman 715 O2 monitor. HC is

measured by the "flame ionization detection" method using a

Beckman 402 analyzer. NO and NOx are measured by the

"chemiluminescence" method using a Beckman 955 analyzer.

EXPERIMENTAL PROCEDURE

The engine is allowed sufficient time to reach a steady-state

operation. By changing the source air temperature, the

equilibrium mixture temperature is adjusted until it reaches a

prescribed value. The performance measurements are then taken.

After this, the streams are measured via the stream separator.

When the above is completed, the droplet deflector is

inserted into the preparation channel to convert some of the

droplets to liquid streams, and the above testing is repeated.

Next the droplet smasher is inserted into the channel to convert

all of the droplets to liquid streams, and the above testing is

again repeated. In this manner, various percentages of vapor,

droplets, and streams can be evaluated.

The air-to-fuel ratio (A/F) is determined from the measured

exhaust gas concentrations and from an oxygen senser located in

the test cylinder exhaust. Variations in A/F are kept within 3%

of the stoichiometric ratio. Although these variations may seem

small, the performance variables are quite sensitive to the

overall A/F. For this reason, the variation of the overall A/F

ratio is included in the analysis of results.

DROPLET SIZE

From experimentation, the droplet size at the outlet of the

preparation channel is found to be nearly constant under the

equilibrium temperature range studied. The Sauter mean diameter

was measured between 4 and 5 /jm with 10% of the droplets under

3.5 jirni and 90% under 8.0 jum. This is probably the result of the

smaller droplets being able to traverse the channel length, while

the larger ones impinge on the channel walls.

REGRESSION MODELING OF DATA

Variable Analysis

The performance values (IMEP, GOV, BT, CO, CO2, O2, HC, NO,

and NOx) obtained from the above procedure are treated as

functions of the percentage of fuel in two forms (vapor and

816
streams) and the equivalence ratio ($) . Since the sum of the

percentages of fuel in vapor, droplet, and stream form must equal

100%, they are not independent of each other, and only two of the

three are required. The equivalence ratio is defined as the

actual A/F divided by the stoichiometric A/F.

Regression Analysis

The data obtained is fitted to a linear regression model

presented below. "Y" corresponds to the dependent variables (IMEP,

Y = A + B (vapor %) + C (streams %) + D ($)

GOV, ect.). "A", "B", "C", and "D" are constants determined

through a least-squares regression fit. With the regression

relations obtained, "performance maps" can be constructed for a

range of independent variables. This is done later on.

ENGINE PERFORMANGE WITH THE MIXTURE PREPARATION GHANNEL

The 1.9 L engine was operated at a speed of 1500 rpm and a

manifold absolute pressure (MAP) of 49.1 kPa (14.5 in Hg) . The

test fuel was Amoco Silver (93 octane, summer grade) . Table 1

presents regression results obtained using this procedure.

IMEP GOV BT ^°2 CO °2 HC NO NOx

psi kPa •/. deg 7. 7. •/. ppm ppm ppm

A 80. O 551 ^.92 -29.^ 19.3 -16.3 5.50 -255'tO -2'f2 -1 1 .0

B O.i^b 3.2 -0.021 0.'f63 -O . 037 -0 . 008 0.0^6 232 3.ao 5.28

C -0.13 -0.9 0.0't9 0.503 -0.052 -0 . 037 0.0'f3 285 -27.

0

-21.8

D -60.9 -^20 -1 . IB 13.5 -2.28 17.6 -9.03 5699 968 578

87. !•/, 93 . 9V. 77 . 27. 21 .77. 60.07. 88 . 07. 92 . 07, 76 . 37. 6'f . 57.

a 0.570 3.93 0.0<f7 0.^6 O. I'f 0.13 O.O'fa 127 39.6 «»5.9

Table 1: Results from Linear Regression Model for the 1.9 L

Engine Operating at 1500 rpm and 49.1 kPa (14.5 in Hg) (MAP)

Figures 3 and 4 show contour plots of the best correlated

performance parameters. The contour plots are presented on

triangular graphs with the investigated region outlined. The

results generally show that performance is significantly affected

by the mixture preparation.

By increasing the percentage of fuel in stream form; GO, NO,

and NOx are shown to be decreased, as indicated by the slopes in

Table 1. GOV, BT, O2, and HG are shown to be increased. These

trends are generally indicative of less desirable combustion.

This is probably the result of a less homogeneous mixture present

in the cylinder. NO, NOx, GO, and GOV seem more strongly

817
Figure 3: Performance Contours of GOV and O2

3400 3000 2600

Figure 4: Performance Contours of HC

818
influenced by streams, as indicated by their greater slopes.

By increasing the percentage of fuel in vapor form; IMEP, NO,

and NOx are shown to be increased. GOV is shown to be reduced.

These trends are generally indicative of more desirable

combustion. This is probably the result of a more homogeneous

mixture present in the cylinder. IMEP seems to be more strongly

influenced by vapor.

By increasing the percentage of fuel in droplet form; BT, O2,

and HC are shown to be reduced. CO is shown to be increased.

These trends are generally indicative of desirable combustion.

BT, O2, and HC seem to be more strongly influenced by droplets.

EQUIVALENT PORT INJECTOR OPERATION

An equivalent fuel preparation (percentage of fuel in vapor,

droplet, and stream form) produced by the port injector can be

deduced from the above regression relations. In theory, all nine

regression relations may be used. Ideally, only three of the nine

relations are required for a solution. But because of the

existence of variation in data measurements, the solutions from

different combinations may not be the same. The accuracy of the

deduction depends upon the fit of the regression relations and

upon the magnitude of error in the data. The fit of the data with

the regression relations is indicated by the coefficients of

correlation (R ) . Large R values are indicative of a "good" fit.

Therefore when deducing an "equivalent" fuel form, regression

relations with relatively high R values should be used.

Figure 5 shows the "equivalent" fuel form produced from the

port injector. This deducted fuel form may not represent the

actual fuel form but the "equivalent" fuel form with respect to

engine performance. In other words, the fuel form produced by the

injector burns in the same manner as an "equivalent" fuel form

prepared in the channel. Figure 5 also shows the gradients

corresponding to the various performance parameters . The arrows

show the direction of maximum increase in performance. The length

of the arrows is proportional to the degree of performance change.

From Figure 5, the direction of improvement, in relation to

mixture preparation, in the existing port injection system can be

readily seen.

CONCLUSIONS

(1) The testing procedure can be used to evaluate the

relative effects of mixture preparation on engine performance.

(2) The testing procedure offers a means of evaluating an

equivalent mixture preparation introduced by a port fuel injector

with regards to engine performance.

The following conclusions are valid for the Ford 1.9 L EFI HO

engine operating at a speed of 1500 rpm and an absolute manifold

pressure of 49.1 kPa (14.5 in Hg)

:

(3) By increasing the amount of fuel in stream form; CO, NO,

and NOx are decreased. COV, BT, O2, and HC are increased. NO,

NOx, CO, and COV seem more strongly influenced by streams.

819
BT I R| = 0.505

HC I R| = 0.037

NOh I Rl = 0.002

NO I R| = 0.003

GOV R| = 0.053

Figure 5: "Equivalent" Fuel Form and Performance Gradients

(4) By increasing the amount of fuel in vapor form; IMEP,

O2, NO, and NOx are increased. GOV is decreased. IMEP seem

more strongly influenced by vapor.

(5) By increasing the amount of fuel in droplet form; BT,

O2, and HC are decreased. CO is increased. BT, O2, and HC

seem more strongly influenced by droplets.

(6) In general, engine performance is diminished by

increases in the amount of fuel in liquid stream form.

REFERENCES

1. Robison, J. A. and Brehob, W. M., "The Influence of

Improved Mixture Quality on Engine Exhaust Emissions and

Performance," Journal of the Air Pollution Control Association,

Vol. 17, No. 7 (July 1967).

2. Matthes, W. R. and McGill, R. N., "Effects of the Degree

of Fuel Atomization on Single-Cylinder Engine Performance,"

Society of Automotive Engineers Transactions, No. 760117 (1976)

.

3. Nogi, T., Ohyama, Y., Yamauchi, T., and Kuroiwa H.,

"Mixture Formation of Fuel Injection Systems in Gasoline Engines,"

Society of Automotive Engineers Transactions, No. 880558 (1988)

.

4. Namiyama, K., Nakamura, H., Kokubo, K., and Hosogal, D.,

"Development of Ultrasonic Atomizer and Its Application to S. I.

Engines," Society of Automotive Engineers Transactions,

No. 890430 (1989)

.

820
ICLASS-91 Gaithersburg. MD. U.S.A. July 1991 Paper 92

EFFECT OF FLOW RATE ON THE SPRAY CHARACTERISTICS

OF IMPINGING WATER JETS

P. Vassallo and N. Ashgriz

Department of Mechanical and Aerospace Engineering

State University of New York

Buffalo, NY, U.S.A.

ABSTRACT

The effect of flow rate and injector type on the spray characteristics of

impinging, turbulent, water jets for jet Reynold's number in the range of 3,000

< Re < 7,500 is investigated. These characteristics are obtained using a Phase

Doppler Particle Analyzer, and include diameter, velocity, number density and

volume flux. For two impinging jets at a 45° angle the shattering improves (i.e.

drop sizes reduce) as the jet velocities increase. For four impinging jets (two

45° jets with two straight jets in the middle) there exists a critical straight

jet velocity for any given 45° jet velocities. The shattering at the edges of

the spray improves above this critical condition and worsens below it. The

shattering in the center of the spray constantly worsens as the straight jets

are added.

1 . INTRODUCTION

The performance of liquid propellant rocket engines is intimately related

to the spray characteristics generated by the fuel injectors. In rocket engines

the fuel injectors are typically of impinging jet type. Two or more jets of

liquid fuel and oxidizer are impinged on each other at high velocities, the

outcome of which is shattering of the liquid into small drops and mixing of the

fuel and oxidizer. The spray characteristic of these injectors were extensively

studied in late 1950's and early 1960's [1], and some information their overall

behavior was obtained. The basic characters of these injectors are discussed

next.

When two liquid jets impinge, they form a liquid sheet (also called a fan)

perpendicular to the plane of the two jets. The liquid sheet disintegrates

intermittently to form groups of drops, which appear to propagate as waves from

the point of impingement. Heidmann and Humphrey [2] measured the frequency of

this wave formation using a photoelectric technique over a finite time interval

under constant operating conditions. Later, Heidmann, et al. [3] extended their

previous results to cover the effects of liquid viscosity and surface tension.

They discovered that the wavelike pattern was more pronounced with higher

viscosity liquids. Also, they identified four regimes of spray pattern as a

function of jet velocity: Closed rim. Periodic drop, Open rim, and Fully

developed impinging spray. Each of the aforementioned regimes is encountered in

the order shown as the velocity is increased. Continuing their investigations,

Heidmann and Foster [4] studied the effect of the impingement angle on drop size

distributions for a range of velocities. They obtained drop counts from

shadowgraphs using a particle analyzer, and concluded that all distributions

showed bimodal characteristics. Overall volume-number mean and mass median drop

821
diameters were observed to increase with a decrease in velocity and a decrease

in impingement angle, with impingement angle having the most pronounced effect.

In all of the above studies, simple photographic techniques were used to

obtain drop sizes and liquid sampling technique was the main measuring method

for the distribution of the volumetric flow rate. Due to the recent advances in

the spray diagnostic systems it is now possible to obtain detailed information

throughout the whole spray. In this paper, the effect of flow rate on the spray

characteristics of a two- jet and four- jet rocket injector is described.

2. EXPERIMENTAL SET-UP

An actual rocket injector of a 22N thruster manufactured by Atlantic

Research Corporation is used in these studies. The injector contains four

orifices so that liquid fuel and liquid oxidizer may impinge on each other to

provide for the mixing necessary for the combustion process. The diameter of the

two fuel orifices are 0.3 mm, and are set to impinge at a 90° angle. The oxidizer

streams are 0.356 mm in diameter, 1.27 mm apart from each other, and emanate

straight down into the fuel streams, resulting in a 45° impact angle between each

fuel -oxidizer pair. Figure 1 shows a schematic diagram of the impingement.

Although water is used in place of fuel and oxidizer for testing purposes , we

will still refer to the liquid jets using the terms fuel and oxidizer to

distinguish between the two.

A Phase Doppler Particle Analyzer (PDPA) , designed and manufactured by

Aerometries Inc. [5-7] is used for simultaneous measurements of diameter,

velocity, number density, and volume flux throughout the complete spray field.

A 3-D traversing system is designed and built in order to map the spray field.

The injector is mounted on the traversing system, and the PDPA is kept

stationary. The traversing system is mounted on the top of a 90 cm x 90 cm x

152 cm stand. The traversing system moves by a leadscrew- stepping motor

arrangement on linear motion ball bushing bearings. The range of motion is

approximately 46 cm in each direction and the stepping motors provide accurate

steps for motion as small as 1 mm. A conical liquid catcher collects the spray

and channels it into a drain. Besides the traversing stand, another stand was

built to hold two 30 liter water storage tanks (one for the fuel line and the

other for the oxidizer line) , and to house the control valves and pressure gages

that were used to monitor the flow. The flow rate is measured by an orifice flow

meter.

3. VARIATION OF SPRAY PROPERTIES WITH FLOW RATE

For this particular injector the drops are not spherical up to 5 cm

downstream of the injector. At this location a 4 cm x 4 cm zone encompasses the

whole spray region. Therefore, detailed measurements of the spray character are

confined to this zone . The spray distribution throughout a 4 cm x 4 cm region

at an axial distance from the face of the injector of 5 cm is obtained for

several different flow rates and jet orientation. Figure 1 shows the reference

axes used throughout this report. For example, y = 0 corresponds to a centerline

perpendicular to the plane of the impinging jets, and x = 0 corresponds to a

centerline parallel to the plane of impinging jets. The z coordinate is kept

constant at 5 cm. In order to see the effect of the flow rate on the spray character several

critical points in the spray are examined to determine how each spray parameter

changes with changing flow rate. These points are (0,0), (±2,0), and (0,±2). It

is observed that the results at (2,0) corresponded well with those at (-2,0) and

822
the results at (0,-2) corresponded well with those at (0,2), indicating fair

symmetry. The results for (0,0), (0,-2) and (2,0) are shown in Figs. 2-4.

Figure 2 shows the variation in parameters as a function of flow rate at

the center of the spray (0,0). This figure is divided into two regions. In the

first region, only the fuel jets are flowing and the spray character of two

impinging jets are investigated. In the second region all four jets are flowing.

However, the flow rate of two fuel jets is kept constant at 3.18 ml/s, while the

oxidizer jet flow rate is changed. It can be seen that for fuel only injection,

the diameter decreases, the velocity increases, the number density increases,

and the volume flux decreases with increasing flow rate. The diameter decrease

and number density increase follow from increased shattering, and the velocity

increase follows from increased flow rate . The volume flux decreases because at

higher impingement velocities (higher flow rates) , flow deflects more to the

sides. As the oxidizer is added at a constant fuel flow rate of 3.18 ml/s, the

diameter and volume flux constantly increase. This is due to the added mass in

the center of the spray. At first the velocity feels little change, then it too

increases as the oxidizer is increased. Note that the oxidizer flow rate where

the velocity begins to increase is 2.72 ml/s. This implies below 2.72 ml/s, the

added oxidizer stream does not significantly effect the drop velocity in the

center, but after 2.72 ml/s it does. The number density is seen to constantly

decrease. The increase in drop size and decrease in number density indicates that

the efficiency of shattering at the center of the spray constantly reduces with

adding the oxidizer jets.

Figure 3 shows how diameter, velocity, number density and volume flux vary

as a function of flow rate at (0,-2). Again the diameter decreases, velocity

increases, number density increases as fuel flow rate increases. However, the

volume flux increases in this section of the spray with increasing flow rate

where in the center it decreases. This is due to the fact that the number density

increases at a faster rate at (0,-2) than it does at (0,0), since more flow is

deflected toward the sides. As the oxidizer streams are added to the nominal fuel

streams, an interesting pattern is observed. For oxidizer flow rates less than

2.72 ml/s, increasing oxidizer flow increases diameter and velocity while for

oxidizer flow rates greater than 2.72 ml/s, increasing oxidizer flow rate

decreases diameter and velocity. This implies a critical oxidizer flow rate

exists for a particular fuel flow rate. The reason such a critical value exists

is that at relatively low oxidizer flow rates the added oxidizer stream reduces

the effective impact velocity of the fuel jets, and also causes the impingement

angle to decrease. This results in an increase in diameter and a decrease in

number density due to less efficient shattering. This effect continues up to the

critical flow rate. As the oxidizer flow rate is increased further than critical,

the impingement angle decreases even further such that a smaller fraction of the

spray reaches (0, ±2). This small fraction contains small drops. The decrease

in impingement angle may be observed by comparing Fig. 5a, Fig. 6a, and Fig.

7a. Note that the point of impingement is pushed further downstream with higher

oxidizer flow, thereby decreasing the mass of fluid that travels to (0,-2). It

is also observed that both number density and volume flux constantly decrease.

Figure 4 shows variation in spray parameters at (2,0) as a function of flow

rate. As before, for fuel only injection, the diameter decreases, and velocity

and number density increases as fuel flow rate increases. However, the volume

flux is nearly constant indicating that the spray becomes more dispersed in the

vicinity of this point as the flow rate increases. As the oxidizer is added,

the critical peak at 2.72 ml/s is again observed. Here again, diameter and

velocity increase, with increase in oxidizer flow rate up to 2.72 ml/s. This

effect in the direction of y = 0 is due to the splitting of the relatively low

823
speed oxidizer stream impacting the high velocity fuel jets. Figure 6b shows a

photograph of this case. As the oxidizer impinges on the fuel jet, it splits in

the X direction at an angle dependant on its flow rate. As the flow rate

increases, the split angle goes down. The split flow passes through point (2,0)

at the critical flow condition. Below the critical flow rate the split flow is

outside of point (2,0), therefore smaller mass of liquid reaches that point.

After the critical flow condition, the oxidizer -fuel jet impingement results in

a better shattering, reducing the splitting effect. This explains the observed

increase of the volume flux up to the critical point, and then its decrease,

along with diameter and velocity, after the critical value. Since the number

density increases in the primary shattering direction (fan direction) for fully

developed impingement it increases after the flow rate of 2.72 ml/s.

4. FULL SPRAY CHARACTERIZATION

Although the variation in spray parameters as a function of flow rate has

been discussed for several important points in the spray, it is difficult to gain

a global perspective without looking at entire spray cross section. In the next

section, the complete contours of the spray parameters for a four jet injector

is given. The flow rate of the fuel jets is set at 3.18 ml/s and the flow rate

of the oxidizer jets is set at 4.08 ml/s.

4.1 Four- Jet Injector

The spray contours of diameter, velocity, number density and volume flux

for the four -jet case are shown in Figs. 8-11, respectively. The diameter contour

of Fig. 8 shows that the largest diameters are bunched about the center of the

spray. This is a result of the addition of the heavy flow of oxidizer stream

straight down into the fuel jets. If the oxidizer jets were not flowing, the

diameters at the center of the spray would be among the smallest present, since

the fuel jets would be shattering with full impact at that location. Also, for

two -jet impingement, the magnitudes of the diameters in the fan direction (xdirection) would be relatively the same, but for the four- jet injector, this is

no longer the case. In fact, the pattern that exists is for the diameter to

decrease in every direction away from the center.

The number density contour is shown in Fig. 9. As this figure shows, the

largest number density occurs in the area of (±2,0), a medium number density

occurs in the center, and the smallest nvunber density occurs in the area of

(0,±2). The reason the number density is highest at (±2,0) is that this is the

primary shattering direction for the fan which is formed by each pair of fueloxidizer streams. The number density is lowest at (0,±2) because the oxidizer

stream issues straight down and tends to inhibit any spray from reaching the

edges in the y direction. There is no simple relationship between diameter and

number density, as there is for the two -jet case. There the number density is

seen to be inversely proportional to the diameter raised to a power, but here

the largest drops in the center have a higher number density than the smaller

drops in the region of (0,±2).

The volume flux is shown in Fig. 10. It is observed that a symmetric heavy

band of flux occurs about y = 0, and that the flux falls off sharply with

increasing y. The largest flux occurs in the center, and the flux decreases

slightly as the x position is increased. The heavy band of volume flux coincides

with the fan, which can be observed by referring to the pictures in Fig. 7, which

show perpendicular and parallel views of the four- jet case.

The velocity contour of Fig. 11 has a similar pattern to the diameter

824
contour. This is a result of the general size -velocity correlation present in

sprays where the largest diameters also have the largest velocities. The reason

this correlation exists is that in the presence of air resistance, small drops

initially moving at the same velocity as large drops will lose their velocity

faster than the large drops. Thus, by the time the drops reach the z = 5

measurement plane, the smaller drops are moving slower than the larger drops.

It is seen in Figs. 8 and 11 that both diameter and velocity plots lose a portion

of their symmetry in the region near (-1,-1), and the magnitudes are both

slightly larger than would be expected for a perfectly symmetric case.

5. CONCLUSIONS

Spray characterization of two and four impinging water jets is undertaken

to investigate the effect of flow rate on different spray parameters. For 45°

impinging water jets (referred to as fuel jets) the flow rates in the range of

1.81 ml/s to 3.63 ml/s are examined. The x and y directions are defined as the

directions perpendicular and parallel to the impinging plane. In all spatial

positions, increasing flow rate decreases diameter, increases number density

and increases velocity. The volume flux decreases in the center, increases along

y, and remains relatively the same along x for increasing flow rate. A four

impinging jet injector composed of two jets at 45° with two straight jets in the

middle (referred to as oxidizer jets) is also investigated. The flow rates of

the straight jets are varied in the range of 1.81 ml/s to 4.08 ml/s, while the

45° jets are kept constant at 3.18 ml/s. For this type of injector, for any given

fuel flow rate, a critical oxidizer flow rate exists for which the shattering

behavior changes. Drop diameters and velocities increase as the oxidizer flow

rate increases up to the critical condition, after which they reduce (except at

the spray center) . This effect is shown to be a result of the splitting of the

fluid in the x direction, and the suppression of impact angle in the y direction.

At high flow rates for both the fuel (3.18 ml/s) and the oxidizer (4.08 ml/s)

jets it is obtained that the largest diameters are in the center, and that in

all directions from the center the diameter decreases. The number density is

largest at the edges of the x axis, medium in the center, and smallest at the

edges of the y axis. The volume flux is heavy about y = 0, decreasing

substantially as y increases. The velocity is largest in the center, and is

similar in pattern to the diameter contour.

6 . REFERENCES

1. Harrje, D.T., "Liquid Fropellant Rocket Combustion Instability", NASA SP194 (1972).

2. Heidmann, M.F. , and Humphrey, J.C. , "Fluctuations in a Spray Formed by Two

Impinging Jets", NACA TN 2349 (1951).

3. Heidmann, M.F., Friera, R.J., and Humphrey, J.C, "A study of Sprays Formed

by Two Impinging Jets", NACA TN 3835 (1957).

4. Heidmann, M.F., and Foster, H.H. , "Effect of Impingement Angle on DropSize Distribution and Spray Pattern of Two Impinging Water Jets", NASA TN

D-872 (1961).

5. Bachalo, W.D. , and Houser, M.J., "Development of the Phase Doppler Spray

Analyzer for Liquid Drop Size and Velocity Characterizations", 20th Joint

825
Propulsion Conference, AIAA-84-1199 (1984).

Bachalo, W.D. , and Houser M.J., "Phase/Doppler Spray Analyzer for

Simultaneous Measurements of Drop Size and Velocity Distributions", Opt.

Engr., Vol. 23, No. 5, pp. 583 (1984).

Bachalo, W.D. , Rudoff, R.C., and Brena de la Rosa, A., "Mass Flux

Measurements of a High Number Density Spray System using the Phase Doppler

Particle Analyzer", AIAA 26th Aerospace Sciences Meeting, AIAA-88-0236

(1988)

.

OXIDIZER

FUEL

• • 2 *y

FUEL

Fig. 1 Schematic of Impingement and Axes of Reference

826
Oxidizer riow Rate (ml/s)

1-81 2.72 3.63 3.18 3.18 3.1*

Fuel Flow Rat« (ml/s)

Fig. 2 Parameter Variation at (0,0): Mean Velocity () Ymin=7.126,

Ymax=15.987 m/s; Number Density (*) Ymin=780, Yinax=2800 /cc;

Volume Flux () Yinin=.36, Yniax=3.6 cc/sec/cm2; Diameter (•)

Ymin=23.377, Ymax=84.387 urn

Oxidizer flow Rate (ml/s)

0.0 0.0 0.0 2.27 3.18 4.08

1.0

0.75-

K 0.50-

0.25-

0.00

1.81 2.72 3.63 3.18 3.18 3.18

Fuel Flow Rate (ml/s)

Fig. 3 Parameter Variation at (0,2): Mean Velocity () Yinin=5.441,

Ymax=9.195 m/s; Number Density («) Yinin=5.7, Ymax=380 /cc;

Volume Flux («) Yinin=-0011, Yinax=-19 cc/sec/cm^; Diameter (•)

Ymin=36.489, Yinax=97.526 um

0.0 0.0 0.0 2.27 3.18 A. 08

c

E

I

1.0

0.75

0.50-

0.25

0.00

1.81 2.72 3.63 3.18 3.18

Fuel Flow Rate (ml/s)

3.18

Fig. 4 Parameter Variation at (2,0): Mean Velocity () Ymin=3.924,

Ymax=9.142 m/s; Number Density (*) Ymin^SSO, Yinax=2400 /cc;

Volume Flux (•) Yinin=-ll' Ymax=2.7 cc/sec/cm2; Diameter (•)

Ymin=24.343, Yinax=84.395 um

827
i

I

i

Fig. 5 Parallel and Perpendicular Views of Two-Jet Impingement with

Fuel Flow = 3.18 ml/s; Oxidizer =0.0 ml/s

i

Fig. 6 Parallel and Perpendicular Views of Four-Jet Impingement with

Fuel Flow =3.18 ml/s; Oxidizer =2.72 ml/s

Fig. 7 Parallel and Perpendicular Views of Four-Jet Impingement with

Fuel Flow = 3.18 ml/s; Oxidizer =4.08 ml/s

828
Diameter Contour (fxm) Number Density (j'/cc)

Fig. 8 Diameter Contour

Fuel=3.18 ml/s

Qxidizer=4.08 ml/s

Fig. 9 Number Density Contour

Fuel=3.18 ml/s

Oxidizer=4.08 ml/s

Volume Flux (cc/s/cm' ) Velocity (m/s)

2.00 -1.67 -1 i3 -1.00 -0»7 -OJJ -0 00 0.J3 0 «7 IJJO IJ3 1.67 J OO

-IJJ -1.00 -0 67 -OJJ -O.OO OJJ 067 1.00 1J3 167 7 00 X Position (cm)

-2 00 -1 67 -100 -0 67 -0 3J -0 00 0 35 0 67 1 OO 1J3 167 7 OO X Position (cm)

Fig. 10 Volume Flux Contour

Fuel=3.18 ml/s

Oxidizer=4.08 ml/s

Fig. 11 Velocity Contour

Fuel=3.18 ml/s

Oxidizer=4.08 ml/s

829

ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 93

ANALYTICAL AND EXPERIMENTAL STUDY OF

PRIMARY ATOMIZATION OF WATER JETS

A.J. Przekwas*, J.-G. Lee*, K.W. Gross^, N.A. Chigier* and H. Eroglu

*CFD Research Corporation, Huntsville, AL, U.S.A.

^NASA Marshall Space Flight Center, Huntsville, AL, U.S.A.

*Dept. of Mechanical Engineering, Carnegie Mellon University, Pittsburgh, PA, U.S.A.

ABSTRACT

The objectives of the present investigation are to predict primary atomization of liquid jets

employing the jet embedding technique coupled with jet instability analysis, and to compare the

prediction with experimental measurements. A novel jet breakup model is proposed, which reduces the

computation time by an order of magnitude as compared to the previously reported breakup model.

Experiments were performed to measure liquid jet core intact length and droplet size distribution.

Results of the validation study show that the predictions agree favorably with the experimental

measurements on liquid jet intact length and drop sizes at different flow conditions.

INTRODUCTION

Gas-liquid two-phase flow is common to practical combustion systems [1]. Primary atomization of

liquid jets has been studied for over a century, which includes classical works such as Rayleigh [2] and

Weber [3], and recent studies such as Reitz and Bracco [4], Meyer and Weihs [5], Lefebvre [1] and Lee and

Chen [6]. It has been recognized that the relative velocity between the gas and liquid is the cause of

the primary atomization resulting in small droplets detaching from the liquid surfaces. The present

investigation focuses on the numerical prediction of atomization of liquid jets with high speed air

flows.

Przekwas, et al. [7] studied the primary breakup of liquid oxygen jets with high speed gaseous

hydrogen annular jets. Predicted results were in favorable agreement with Sutton, et al. [8] (CICM

formula) and Liang, et al. [9]. The frequency scanning approach used in the original atomization model

[7] required long computational time. In the present paper, instead of scanning the entire wavelength

spectrum, the maximum growth rate is computed by analytic differentiation of the dispersion equation.

The breakup mass and drop sizes are calculated based on the local wavelength. An order of magnitude

gain in computational time was achieved with minor change in accuracy of the results. The breakup of

the surface into small droplets is assumed to occur when the local amplitude of the disturbance exceeds

one half the wavelength of the disturbance at the maximum growth rate. The results of the calculated jet intact length and droplet size are compared with the

complementary experimental measurements recently conducted by Chigier and Eroglu [10] for liquid jet

atomization under the influence of surrounding gas flows. Atomization characteristics such as liquid core intact lengths and droplet sizes were measured. The core intact length was determined from

photographs, while drop size was measured using the Phase Doppler Particle Analyzer (PDPA).

831
JET EMBEDDING TECHNIQUE

In the Jet Embedding (JE) technique, Przekwas [11], the liquid sheet is described in a one- dimensional form with a computational grid conforming to the liquid jet shape. At present, the

gravitational force and internal viscous effects are neglected. The friction between the liquid and gas is accounted for at the jet interface. The basic conservation equations of mass, axial- and cross-stream

momentum are expressed as follows:

aPf aPfUf

at

-^^="^6

t2 =

-^ai^

- (Uf - U^d) - ^s" Pf (Uf - Ug) 2 (2)

aPfVf aPfUfVf • o . « -7 + y = - mB (Vf - V°d) - ^n" Pf (Vf - Vg) 2 (3)

where mg is the droplet breakup rate, Uf and Vf are liquid velocities, U d and V d are initial droplet

velocities leaving the jet surface, Pf is the liquid density, and ^"s and ^"n are the gas-liquid friction

coefficients for the coaxial and cross flow directions, respectively.

The three conservation equations are approximated by the finite volume method on a one- dimensional adaptive, surface conforming grid (i.e., see Figure 1). A space marching technique is

employed to integrate the governing equations along the liquid surface trajectory from the nozzle exit

until the liquid jet mass is completely atomized. A schematic of an imbedded jet and its interaction

with the gas phase is shown in Figure 1.

Fig. 1 Schematic representation of the liquid jet breakup using an

Eulerian Computational grid for the gas phase equations.

LINEAR STABILITY ANALYSIS OF LIQUID JETS WITH HIGH SPEED AMBIENT GAS FLOWS

As a liquid jet issues from a nozzle, its surface is subjected to various disturbances, caused by

vibrations of the nozzle, shape of the jet, gas motion surrounding the jet, turbulence in the liquid, and

roughness of the wall surface. Disturbances can be represented by the superposition of Fourier

components. A Fourier component is expressed in the following form:

C = Co e ik2+at (4)

832
where Co is the initial amplitude of an infinitesimal axisymmetric displacement (disturbance) at the

interface, k is the wave number, and a is the growth rate of disturbances.

After algebraic manipulation (e.g., Levich [12] and Chuech, et al. [13]), a governing wave

disf)ersion equation is obtained in the general form:

2 ,9 I'l(l^a) 2kl Ii(ka)ri(la) " "^"f'' Wa) k2+l2lo(ka)Io(la)

gk ^ . l2-k2 Ii(ka) PgUr2k2 i2.k2 ii(ka) Ko(ka)

-Pfa2^ ^ ^S2+k2^l0(ka)-' S2+k2 ^IO(ka)Ki(ka)

where a is the jet radius, a is surface tension, pf is liquid density, Vf is liquid kinematic viscosity, 1^ = k^ + a^, and InW and Kn(x) are Bessel functions of the first and second kinds of order n, respectively. The

two terms on the LHS represent liquid inertia and viscosity, while the two RHS terms represent the

surface tension and gas-liquid relative velocity effects. Numerical solutions of the full dispersion

equation were obtained by Chuech, et al. [13]. However, considerable useful information can also be

obtained by analyzing the solutions of the simplified limiting forms from the full dispersion equation.

In the following, two such limiting forms are discussed.

High-speed Ambient Gas lets With Long Wavelength Disturbances ( 2. > a or ka < 1)

Ko(ka) k a For long wavelength disturbances,

^^^^^^^ = - ka In ( "2" )/ the dispersion equation can be

approximated as:

o o ok2 T ^ PaUr2k4a2 ka

a2 + Vf k2 a = (l-k2a2) - ^

In

( y) (6)

2 a Pf 2Pf

High-speed Ambient Gas Tets With Short Wavelength Disturbances ( ^ < a or ka > 1)

For short wavelength disturbances, Ii(ka) = lo(ka), Ii(la) == iQda), Ii'(ka) = Ii(ka), and Ko(ka) =

Ki(ka), the simplified equation is expressed as:

o . ak3 PgUr2k2

a2 + 2vfk2a = - —-^^-^

(7)

It is noted that for high-speed gas flows, for both cases, the dispersion equations reduced to quadratic

form which admits real and imaginary solutions. If the real part of the solution, ar > 0, the

disturbances grow exponentially in time, otherwise, the disturbances will be damped out.

MODEL OF PRIMARY ATOMIZATION

To improve the computational efficiency of the atomization model by Przekwas [7], an

alternative breakup model was postulated. In the original Jet Embedding model, the entire

wavelength spectrum, X, was employed. Here the wave dispersion equation is differentiated and the

maximum growth rate wave length, X^ax' is directly computed. Based on the local wave amplitude ^,

and the maximum wavelength Xmax, a surface breakup mechanism is invoked to calculate liquid

breakup rate. Since finding the wavelength at the maximum growth rate can be done much faster than

scanning the entire wave spectrum, much less time is required to obtain the results. Assuming Xmax as the characteristic length of the atomization process, i.e., Xmax = 2Ddrop/ where D^rop 'S the diameter

833
of the liquid droplets detaching from the liquid surface, the breakup rate per ur\it circumferential area (mass leaving the liquid surface due to breakup) is therefore calculated as:

• „ n V n Xmax ^max

^"aST " 16 In (Xmax / 2^0) ^

^

A X

where n = r is the number of wavelets per unit length.

^max

V = 71 D ( — ) is the volume of the liquid ring generated by a single wavelet, and

T= —-—In ( ^aH^^ ) is the characteristic time for a wavelet growing into a liquid ring,

"max 2^

COUPLING PROCEDURE

The main steps of the coupling procedure between the Jet Embedding technique and the proposed

primary breakup model in the present computation are sunnmarized below:

1 ) Calculate the jet velocity and diameter at a given axial location using the Jet Embedding technique

(see Figure 1).

2) Update the coefficients in the dispersion equation (7) and solve the equation for the wavelength

corresponding to the maximum growth rate. 3) Calculate the mass breakup rate (Eq. 8) and the droplet size at the same location.

4) Move to the next axial position along the jet axis, and go back to Step 1 with the updated mass

breakup rate.

EXPERIMENTAL TEST CONDITIONS

In order to validate the proposed atomization model, experimental measurements were

performed for the liquid jet atomization in coflowing air streams. Details of the experimental study

were reported by Chigier and Eroglu [10]. Two different nozzle geometries, namely, plane and round

jets, were investigated. For the round jet, two different central tube designs were tested to assess the

influence of nozzle geometry on the atomization process. The working fluids are air (in the annulus) and

water (in the central tube). Atomization characteristics such as liquid intact lengths and droplet sizes

were measured, the former was determined from photographs, while the latter was obtained using the

Phase Doppler Particle Analyzer (PDPA). Chigier and Eroglu [10] also reported the non-axisynunetric

breakup of liquid jet, and its theoretical analysis has been performed by Yang [14]. Since only the round

jet case is relevant to our study, the following discussion is limited to the round jet results. The

dimensions of the nozzle are given in Figure 3, and the test conditions are summarized in Table 1. The following empirical correlation was formulated from the photographs by Chigier:

^=0.5We-0-4ReO-6 (9)

Py Do(uL-Ug)2 py ut Do

where L is jet intact length, Dq is the diameter of the central jet, We= — —— and Re= ^ ,

and a is surface tension, pg, ug and pL, ul are densities and velocities of air and water, respectively.

The correlation implies that the intact length is not only controlled by the relative velocity (i.e.,

Weber number), but also by the momentum of the liquid jet (i.e., Reynolds number). Experimental results

also show that intact lengths are independent of the nozzle shape, but it does have an impact on the

droplet sizes.

834
10.36

0.971

1 ^6?

16.0

Fig. 2 Nozzle geometry for coaxial atomizer (dimensions in mm), Chigier and Eroglu [10].

Table 1 Experimental conditions of air-water jets by Chigier [10].

Water Velocity (m/s) Reynolds Number Air Velocity (m/s) Weber Ntimber

1.51

4.52

9.65

1456

4370

9328

36.7 - 159.0

39.8 - 162.0

44.9 - 167.0

50 - 200

50 - 200

50-200

RESULTS AND COMPARISON

Computations were made by employing the Jet Embedding technique. Equation (7), (8), and the

coupling procedure. Air and water properties at standard conditions were used, and the test conditions

are given by Table 1. The calculated liquid jet lengths are shown in Figure 3 for Re=1456 and We=10-

200. It is obvious that the jet length is decreased as Weber number increases. Increase of Weber number

means increase of relative velocity between the air and liquid, assuming fluid properties remain

constant. A higher relative velocity results in a higher breakup rate, i.e., a shorter jet. It should be

noted that the steps on the jet surfaces in Figure 3 are due to the resolution of the graphics software. Jet velocities as a function of axial distance are plotted in Figure 4, which shows that higher gas velocity

reduces jet lengths and results in a higher jet velocity as the jet diameter approaches zero. Figure 6

depicts the droplet size variation under the same flow conditions. It appears that the droplet size

decreases significantly as Weber number number increases, particularly when Weber number is changed

from 10 to 50. The droplet sizes are under 50 |jjn for Weber number =50 to 200. The size of the droplets at the separation increases along the jet direction axial direction (Figure 5), as the relative velocity

between gas and liquid gradually diminished.

Calculations were also performed for Reynolds numbers 3270 and 9328, and Weber numbers ranging

from 10 to 200. Results with similar trends were obtained. Figure 6 shows the calculated jet intact

lengths under different flow conditions, including Reynolds numbers equal to 3270 and 9328. The

experimental observations of the jet intact length by Chigier are included for comparison. Comparison

shows good agreement at low Reynolds number, and is less satisfactory as Reynolds number is increased.

It is noted that the empirical correlation shown above for the measured breakup lengths was obtained

from the scattered experimental data (i.e., see Figure 6). Figure 7 shows the comparison between the

experimental correlation and the prediction from the CICM code [8]. It is seen that the present

prediction compares favorably with the jet intact length calculated from the CICM code for Reynolds

numbers 1456 and 3270. For Re=9328, the present prediction shows better prediction in the low Weber

number regime, and less satisfactory in the high Weber number regime than CICM prediction.

However, it is noted that an empirical constant has to be determined in the CICM mass breakup rate

correlation:

Mass breakup rate mg = Ca [" (9)

where Liang, et al. (1986) use Ca = 0.0378 in the injector element cup region, and 0.14 in the main

chamber. The constant has to be estimated from experimental measurements. The present CICM results

shown in Figure 8 use Ca = 014 in the mass breakup rate equation (9), and it is also noted that previous

835
CICM results from the SSME LOX jet calculation (Chuech and Przekwas, 1988) used Ca = 0.0378. The

present prediction, however, using the proposed breakup model performs faster and without adjusting

the constants.

The predicted drop sizes and measured Sauter Mean Diameter (SMD) are presented in Figure 8. It should be noted that the measured drop sizes are obtained approximately 2 mm downstream of the jet's

tip location, and SMD variation in the radial direction was recorded. Since the present numerical

analysis predicts droplet sizes leaving the liquid surface, but not beyond the jet tip, a direct comparison

is not possible. To circumvent the difficulty, we assume that all droplets leaving the jet surface reach

the downstream location with negligible coalescence. The predicted droplet size along the jet axis

remains relatively unchanged {e.g., see Figure 5), and only the averaged droplet sizes are plotted in

Figure 8. The measured droplet sizes show variation in the radial direction, and only the minimum and

maximum droplet sizes in the radial direction are shown in Figure 8. It appears that the present

prediction agrees reasonably well with the measured droplet size for low Weber numbers, and

underestimates the droplet size as Weber number increases.

2

-22 D 0

-22

-2

10 x/D

—I 1—1—I 1 Re « 1456, We » 10

Re = 1456,

Re = 1456,

Re = 1456,

Re = 1456

We = 50

We = 100

We > ISO

We = 200

20

10 x/D

Re = 1456

IS 20

Fig. 3 Liquid jet profiles for Re=1456 and

We=10-200.

Fig. 4 Axial liquid jet velocities for

Re=1456 and We=l0-200.

300

§ 250 We = 10

e 200 :

Re = 1456

15 20

s

s00

U

125

100

1

'

Present Prediction Experiment (Chigier, 1990)

Actiul diu

x/D. Re = 1456

x/D, Re = 3270

x/D, Re = 9328

• L/D.Re = 1456 L/D,Re = 4370 * WD. Re = 9328

CoTTcUtcd curvei UD, Re= 1456

L/D, Re = 4370

I7D. Re = 9328

Weber Number

Fig. 5 Estimated droplet diameters in the axial

direction for Re=1456 and We=10-200.

Fig. 6 Comparison of liquid jet intact

lengths between present prediction

and experiments by Chigier [10].

836
Weber Number

Fig, 7 Comparison of liquid jet intact lengths

between CICM prediction and

experiments by Chigier [10].

80

I

5 40

o.

20

- 1 • 1 r—, . 1 . 1 r—

• Present Prediction Chigier (1990) A — 1 1 1A

•

Rc = 1349 - O Re = 2502 -A a Rfi = 4370 -

-a

-A •

•

• •

-—1—.—. 1 . 1 .__.—.

.—1 ....

0 50 100 150 200

Weber Number

Fig. 8 Comparison of predicted

and measured droplet sizes.

CONCLUSIONS

A novel breakup model is proposed and implemented in the present numerical analysis exploiting

the Jet Embedding technique. Complementary experimental measurements of water jet atomization

were performed. The comparison shows favorable agreement between numerical data and experimental

measurements of liquid jet intact lengths and droplet sizes for water jets with ambient air flows. The

computation time has been improved by an order of magnitude compared to the previous prediction

method by the authors. The model will be further validated for the SSME LOX jet flow conditions.

ACKNOWLEDGEMENT

The present work was performed under an SBIR study for NASA MSFC (Contract NAS8-38425).

The authors express their gratitude to Dr. A. K. Singhal, President and Technical Director of CFD

Research Corp. for his active participation and valuable guidance.

REFERENCES

1. Lefebvre, A. H., Atomization and Sprays, Hemisphere (1989).

2. Rayleigh, L., Theory of Sound. Dover. Vol 2, pp. 351-365 (1941).

3. Weber, C, "Disintegration of Liquid Jets", Z. Angew. Math. Mech., Vol 11, No. 2, pp. 136-159

(1931).

4. Reitz, R. D. & Bracco, F. V., "Mechanisms of Breakup of Round Liquid Jets" in Encyclopedia of Fluid

Mechanics (N. P. Cheremisinoff, Ed.), Vol 3, Gulf, pp. 233-249 (1986).

5. Meyer, J. and Weihs, D., "Capillary Instability of an Annular Liquid Jet", J. Fluid Mech., Vol 179,

pp. 531-545 (1987).

6. Lee, J.-G. and Chen, L.-D., "A Linear Stability Analysis of Gas-Liquid Interface", AIAA J. (in press)

(1991).

837
7. Przekwas, A. J., Chuech, S. G. and Singhal, A. K., "Numerical Modeling for Primary Atomization

of Liquid Jets", AIAA J. of Propulsion and Power (1991).

8. Sutton, R. D., Schuman, M. D. and Chadwick, W. D., "Operating Manual for Coaxial Injection

Combustion Model", NASA CR-129031 (1974).

9. Liang, P. Y., Jensen, R. J. and Chang, Y. M., "Numerical Analysis of SSME Preburner Injector

Atomization and Combustion process", AIAA-86-0454 (1986).

10. Chigier, N. and Eroglu, H., "Atomization of Liquid jets from Injection Elements in Liquid Rocket

Combustion Chambers", NASA NAG8-126 (1990).

11. Przekwas, A. J., "Spray Formulation in Liquid Propellant Atomization Systems", TANNAF

Workshop on Atomization. NASA MSFC, Huntsville, AL (1987).

12. Levich, V. G., Physicochemical Hydrodynamics, Prentice-Hall, pp. 591-668 (1962).

13. Cheuch, S. G., Przekwas, A. J. and Singhal, A. K., "A Computer Code for Liquid Jet Atomization in

Rocket Thrust Chambers", Appendix C, NASA NAS8-37620 (1988).

14. Yang, H. Q., "Non-axisjmnmetric Breakup of a Liquid Jet during Atomization", AIAA 91-0693, 29th

Aerospace Sciences meeting, Reno, Nevada (1991).

838
ICLASS-91 Gaithersburg, MD. U.S.A. July 1991 Paper 94

DISINTEGRATION PHENOMENA OF METALIZED SLURRY

FUEL JETS IN HIGH SPEED AIR STREAM

T. Inamura, N. Nagai, T. Hirai and H. Asano

Faculty of Engineering

Tohoku University

Sendai, Japan

ABSTRACT

The metalized slurry fuels are anticipated to enhance the combustion efficiency and the

specific thrust of a ramjet and a scramjet engine. The disintegration mechanism of slurry fuel

is expected to differ from that of single phase liquid, such as water. Because the slurry fuel is generally non-Newtonian liquid and it has rheological nature. Consequently it is difficult to apply

the experimental results obtained for the single phase liquid to slurry fuels. This study aims to

clarify the disintegration mechanism and spray characteristics of slurry fuels in high speed air flow.

In this report, the empirical equations of the jet penetration and the jet width for water and slurry

fuels were deduced. Furthermore, break-up lengths of liquid jets and mass flux of water spray drops

were measured. And compared the experimental results of slurry fuels to those of water, it was

clarified that the measuring results of slurry fuels slightly diff^er from those of water.

INTRODUCTION

As the fuel of the air breathing engine such as ramjet and scramjet engine, metalized slurry

fuels have the advantages of high density and volumetric heating value, they contribute to the

compactness of the fuel storage [1]. Yatsuyanagi [2] showed experimentally that the kerosene fuel

including aluminum powder has a good combustion efficiency for a rocket engine. And highly

metalized slurry fuels are expected to enhance the combustion efficiency and the specific thrust of

the ramjet and scramjet engine.

In order to ignite the fuel spray stably, to hold the flame in a high speed air stream and to

complete the combustion in a combustion chamber, improvement of the spray characteristics and

the control of the mixing between fuel sprays and combustion air are required. To accomplish these

purposes, it is needed to clarify the disintegration mechanism of liquid jet and the behavior of the

fuel spray in a high speed air stream.

On the other hand, high metalized slurry fuel is generally non-Newtonian liquid. And its

disintegration mechanism seems to differ from that of single phase liquid fuel [3]. For example the

viscosity of slurry fuel depends on the shear velocity, this fact contributes to different disintegration

mechanism.

There are many studies on the disintegration of single phase liquid jet in high speed air

stream [4,5], however concerning to the slurry fuel jet there are few studies [6]. The present study

aims to clarify the disintegration mechanism of metalized slurry fuel jet in a high speed air stream,

the spray characteristics and the mixing characteristics between fuel sprays and air stream. In this

report the empirical equations of the jet penetration and the jet width for metalized slurry fuel and

water were deduced. Furthermore, jet breakup lengths of liquid jets and the mass flux of water

spray drops were measured.

839
APPARATUS AND PROCEDURES

The subsonic air flow was selected in order to observe the disintegration mechanisms and to

measure the atomization characteristics easily. Liquid jet was injected transverse to a subsonic air

crossflow.

Water was used to simulate hydrocarbon fuels and a mixture of water and aluminum particles

was used as a slurry fuel. The aluminum particles have a density of 2.67 g/ml and of irregular shapes.

Their average diameter is 3.1 ^m, and the surfactant was included to prohibit the deposition of

aluminum particles.

Figure 1 shows the experimental apparatus. The air blown off from a blower 1 is supplied to

the test section 4 passed through a settling chamber 2 and a nozzle 3. The air/liquid two phase

flow is decelerated by diffusor and then exhausted. The slurry particles are eliminated by cyclone

separator 5, before they are exhausted. The blower is driven by a motor 6, and the rotational

speed of the motor is controlled by a inverter 7. The pressure vessel 8 for water supply and 9

for slurry supply are pressurized by the air from a screw compressor 10. The concentration of

the water/ aluminum slurry was kept constant during the experiments by a stirrer in the pressure

vessel. The water is metered by orifice flow meter 11, and the metalized slurry is metered by

electro-magnetic flow meter 13 through a trap 12 for a verification of density.

The disintegration phenomena were observed by back-lighted instantaneous photograph. The

microflash 15, exposure time is less than 1 fis was used as a light source in this case. The jet

penetration and jet width were measured by side-lighted streak photograph, exposure time is 1/30

s to obtain the time-averaged contour of the unsteady spray flow.

Fig.l Experimental apparatus Fig. 2 Details of test section

Figure 2 shows the details of the test section. This is 40 x 40 mm subsonic blowdown wind

tunnel up to 140 m/s. Its length is 350 mm. The pressure taps are installed at the top wall of wind

tunnel every 30 mm to measure a static pressure on the wall surface. The Pyrex plates are attached

to both sides of wind tunnel to take a picture. The liquid injection nozzle, directed perpendicularly

to air stream is located at 82 mm downstream of the entrance of wind tunnel. And liquid nozzle

edge is in same plane with the bottom wall surface. Three sizes of liquid nozzle, inner diameter is 0.5, 1.0, 2.0 mm were used. Coordinate system is right-hand system. X coordinate is in a direction

of liquid jet at liquid nozzle exit, Y coordinate is in a direction of width of wind tunnel and Z

coordinate is in a direction of a air stream. And the origin of coordinate system is placed on the

upstream edge of liquid nozzle exit.

Figure 3 shows the contact needle probe to measure the breakup length of liquid jet by

the electrical resistance method. Needle probe is movable up and down and back and forth by

micrometer head 3 and precisely movable stage 4. A direct current voltage, which amplitude is 15

V was applied between the liquid nozzle and the needle probes. If the liquid jet contacts to the

840
Fig. 3 Details of contact needle probe vicinity of liquid nozzle exit

needle probe, then circuit is closed and the output is obtained. The needle probe 1 was used to

measure the breakup length of downstream side of liquid jet, and the probe 2 upstream side.

The local mass flux of droplets was measured by the iso-kinetic sampling probe, which inner

diameter is 0.7 mm and outer diameter 2.5 mm.

Tap water and water/aluminum slurry fuels, 20 wt% and 40 wt% were used as a injectant.

The range of injection velocity of the injectant is 7 ~ 26 m/s and air velocity 55 ~ 140 m/s. The

viscosity of 20 wt% slurry at low shear velocity is 2.1 X 10~^ Pa-s. And the surface tension is 62.5 X 10-3 ^/j^

RESULTS AND DISCUSSIONS

Air Velocity Distribution

Figure 4 shows the air velocity profiles in the vicinity of the liquid nozzle exit measured by

hot wire anemometer. Boundary layer thickness is about 2 mm under these operating conditions.

The turbulent component of the air flow is less than 2 % of the average air velocity.

Disintegration Phenomena of Liquid Jet

Typical pictures of disintegration phenomena of water and slurry jet are shown in Fig. 5. Figure 5(a) and 5(b) are the side view photographs and Fig. 5(c) is the top view photograph. In

the case of water jet (Fig. 5(a)), near the bottom wall liquid jet moves perpendicularly and then

is bent downstream rapidly. The cross section of liquid jet is circular at the nozzle exit, then it

is transformed into kidney shape and the drag by the air stream increases rapidly. This increase

of the drag results in the rapid downstream bend of liquid jet. On the upstream surface of liquid

column, the wave originated at liquid nozzle exit grows rapidly and the small particles are produced

from the side surface and downstream surface of liquid column. On the other hand, the large liquid

lump detaches at the trough of surface wave from the liquid column. Then it is atomized into finer

particles further downstream. The location where the liquid lump detaches from the liquid column

is about three times of nozzle diameter downstream of the origin of the coordinate system. At this

location the fine spray particles are torn off by the shear force of air stream at the surface wave

crest.

In the case of slurry jet(Fig.5(b)), the disintegration phenomena are almost similar to those

of water jet. But the surface waves are slightly smaller than those of water jet. This seems to be

due to the viscosity increase of the slurry or the rheological nature of slurry jet.

841
-10-^

(c) Top view, Water

Vj=11.0m/s,Va=96m/s

Fig. 5 Disintegration phenomena

mm

(a) Water (b) Slurry (20 wt%)

Vj=7.5m/s,Va=100m/s Vj=7.0m/s,Va=100m/s

Fig. 6 Magnified picture of disintegration phenomena

For both of liquid jets the penetrations of the spray increase as the liquid injection velocity

increases, as the air velocity decreases and as the the nozzle diameter increases. And under same

liquid injection velocity, the penetrations increase with increasing the loading of aluminum particles,

that is the increase in the liquid density.

To take the top view photograph(Fig.5(c)) the bottom wall of test section was changed to the

grass plate and the picture was taken back-lighted by the microflash. This picture is in the case of

large injection velocity of the liquid jet. The large liquid lumps detach alternatively from the both

sides of liquid column such as Karman's vortex streets.

Figure 6(a) shows the magnified picture of water jet at nozzle exit. Taking notice of the surface

waves, near the nozzle exit the waves incline downstream. And then they incline upstream. These

842
phenomena mean that the relative velocity of liquid surface to surrounding air directs downstream

first and then it directs upstream. This point where the shapes of surface waves change leaves from

the nozzle exit with increasing injection velocity.

Figure 6(b) shows the magnified picture of slurry jet. The surface waves are obviously smaller

than those of water jet (Fig. 6(a)). This is caused by the stability of the slurry jet due to high

viscosity [6].

Jet Penetration

The jet penetration and the jet width are very important for the design of combustor. The

combustor dimensions are also affected by the jet penetration and the jet width. Furthermore,

combustion characteristics are affected by those.

In this report the jet penetration and jet width were measured by side-lighted streak photographs and the empirical equations of those were deduced.

The jet penetration and jet width were defined as shown in Fig. 7. The jet penetration was

defined as the length from the spray axis to the contour of side view photograph of spray(Fig.7(a)).

And the jet width was defined as that of top view photograph(Fig.7(b)).

( a ) Penetration ( Side view ) ( b ) Jet width ( Top view

)

Fig. 7 Definitions of the jet penetration and the jet width

By the observations of streak photographs the jet penetration increases with increasing the

injection velocity of liquid jet and with decreasing air velocity. And the penetration increases as liquid density increases due to the momentum increase of liquid jet. Therefore the jet penetration

seems to be the function of momentum ratio, nozzle diameter and the streamwise distance from the

liquid nozzle [7]. From these considerations and the preparatory experiments the jet penetration

was assumed as follows:

X,out d = ^(d)rin{l+5(d)|} (1)

where a is constant and A{d), B{d) are function of liquid nozzle diameter, d.

a, A{d) and B{d) were determined by least square method as follows:

for water jet;

^out _

d

for both of slurry jets

^out

(1.18 -F- 0.24 d)f-^^ In |l -H (1.56 + 0.48 d) ^| (2)

(1.17 + 0.16 d) " In |l + (0.75 + 0.95 d) ^| (3)

Figure 8 shows the variation of the jet penetration curve with liquid nozzle diameters. Under

same momentum ratio the larger the nozzle diameter the larger the jet penetration is. Figure

9 shows the variation of the jet penetration curve with various kinds of liquids. Under large

843
15

Injectant: Water

q=10

15

10

•o

q=10

d=2.0mm

ijectant'Slurry

^"^^-^ 1 njectantiWater

5 10 15

Z/d Z/d

Fig.8 Variation of jet penetration curve with Fig. 9 Variation of jet penetration curve with

liquid nozzle diameters various kinds of liquids

momentum ratio (g = 10.0), the jet penetration of slurry jet is larger than that of water. This

trend is different from the results obtained by Less and Schetz [6]. They observed the phase

separation, and they concluded that the separation of the phases results in reduced penetration of

liquid portion of the injectant as the loading is increased. However, in our studies phase separation

was never observed. This is due to the differences of air velocity. They used the supersonic air flow,

and in our case the subsonic air flow was used. In our case the large jet penetration of slurry jet is caused by the stability of slurry jet due to high viscosity.

Jet Width

The empirical equation of the jet width was assumed as follows from the preparatory experiments.

Yrout

where C, 7 is constant.

C, 7 were determined by least square method as follows:

for water jet (for all nozzle diameters);

(4)

f)

0.49

(5)

for slurry jet (for 20 wt %, d=1.0 mm); —^out— r, -, -0.06 = 0.1 q (6)

For the water jet, the nozzle diameters donot influence on the jet width, and this trend is contrastive with that of the jet penetration (Fig.8).

Break-up Location of Liquid Jet

The examples of break-up length of liquid jets measured by the contact needle probes are shown in Fig. 10. The values in percent in the figure indicate the rate of the period in which the

probe detaches from the liquid column. Here 5% break-up location is defined as the break-up

location of liquid jet.

The break-up location in X-direction moves downstream as the momentum ratio increases,

whereas that in Z-direction is almost constant in spite of the momentum ratios and Zjd = 3 ~ 3.5.

This break-up location is almost coincident with the point at which the large liquid lumps detach

from the liquid column.

844
Break Up Rate

Water Water.AI(20wf/.)

5'/. 5°/. 5080

Equations (2),(3)

d = 2.0mm : ^=31

Fig. 10 Break-up length of liquid jet

26

23

20 )0030

17|Q0360 00309 00120^ 0

11

LEGEND

0.001 001 010

016

peak m/Aal

Y/d core region

K|00BS7 QOTOffl 00292 00095

ffisbV " "ffl^«3 - -. Q0405 Q0I6I I 00012

5,1551 OJI7,0| ' 00633 00255/

00873 00837 00390 Q015I Og/smin'

•0 '3 '9 '\2 -15 -18 -20 Y/d

Injectant:Water . cl = 1.0mm

Vj=17m/s,Va=R2.Am/s. Z/d=30

Fig. 11 Mass flux of spray drops

In the figure the break-up location of slurry jet is shown compared to that of water jet. And

the momentum ratio of water jet and slurry jet is kept constant. The break-up location of slurry jet

slightly moves downstream compared to that of water in the case of same momentum ratio. This

seems to be caused by the stability due to the viscosity increase of slurry jet and further researches

on the disintegration mechanism of slurry jets are required.

Mass Flux of Spray Drops

Figure 11 shows the mass flux of spray drops measured by the iso- kinetic sampling probe at

Z/d = 30. Defining the contour of spray as the location of 0.001 g/s-mm^, Xout/d is equal to 20

and Yout/d is equal to 12. Compared this values to the values measured by the photograph, the

penetration measured by the photographs is underestimated about 20 %. And the width measured

by the photographs is underestimated about 5 %.

In the region of X/d = 5 ~ 14 and Y/d = 0 ^ 3, the mass flux is almost constant and

this region is seemed to form the core region of the spray. Furthermore by the measurements of

the mass flux variation in Z-direction, the contour line of 0.001 g/s-mm^ spreads and that of 0.1

g/s mm^ narrows further downstream. By the comparison of the mass flux measurements to the

break-up location measurements, the large liquid lumps detached from the liquid column form the

core region of the spray after those are atomized into spray drops, whereas the fine drops atomized

from the spray lumps in the core region and torn off from the liquid surface by the shear force of

air stream form the surrounding of the core region. Consequently mean diameter of spray drops in

the core region seems to be larger than that in the surrounding of the spray.

CONCLUSIONS

In order to clarify the disintegration mechanism of metalized slurry fuel jet in a high speed

air stream, the jet penetration and the jet width of metalized slurry fuel and water are measured

and the break-up location of liquid jets and the mass flux of water spray drops were measured.

Consequently the following results were obtained.

1. The waves on the liquid surface are amplified rapidly as they move downstream and the large

liquid lumps detach from the liquid jet at the trough of surface wave. These liquid lumps

845
are atomized again into further fine spray and then form the core region of the spray further

downstream.

2. The jet penetration and the jet width can be represented by the empirical equations (2), '(3)

and (5), (6) respectively.

3. The slurry jets are more stable than the water jet, and it results in larger penetrations and

longer break-up lengths.

4. The distance from the liquid nozzle exit to the point where the liquid lumps detach from

liquid jet in a direction of the air stream is almost constant over the whole momentum ratios

for both of slurry jets and water jet.

5. The spray consists of core region in which mass flux of spray drops is constant and the

surrounding region in which the mass flux is smaller than that in core region. The core region

becomes small and the surrounding region becomes large as one moves downstream.

NOMENCLATURE

d liquid nozzle inner diameter

q momentum ratio(= Pj^] / Pa^a)

V velocity

X,Y,Z coordinates

jet penetration

V^ out jet width

p density

Subscripts

a : air

j : liquid

ACKNOWLEDGEMENTS

The authors would like to thank Dr. N. Yatsuyanagi and Mr. A. Kumakawa of National

Aerospace Laboratory for their valuable advice. They also acknowledge the help in the experiments

with Mr. N. Fukuda, Mr. M. Kuroda and Mr. M. Tomita of Tohoku University.

REFERENCES

1. Northam, G. B., "Combustion in Supersonic Flow", Proc. 21st JANNAF Combust. Meeting , pp. 399-410, JPL,Pasadena,CA (1985).

2. Yatsuyanagi, N., "Combustion Characteristics of Metalized Hydrocarbon Fuels", Journal

J.S.A.S.S., Vol.37, No.427, pp.393-399 (1989) (in Japanese).

3. Inamura, T., Nagai, N. and Inagaki, H., "A Study on the Twin-Fluid Atomization of a Highly

Concentrated Coal- Water Mixture", Trans. J.S.M.E., Ser.B, Vol.53, No.490, pp.1846-1851

(1987) (in Japanese).

4. Catton, I., Hill, D. E. and McRae, R. P., "Study of Liquid Jet Penetration in a Hypersonic

Stream", AIAA Journal, Vol.6, No.ll, pp.2084-2089 (1968).

5. Kush Jr., E. A. and Schetz, J. A., "Liquid Jet Injection into a Supersonic Flow", AIAA

Journal, VoLll, No.9, pp.1223-2114 (1973).

6. Less, D. M. and Schetz, J. A., "Penetration and Breakup of Slurry Jets in a Supersonic

Stream", AIAA Journal, Vol.21, No. 7, pp.1045-1046 (1983).

7. Thomas, R. H. and Schetz, J. A., "Distributions Across the Plume of Transverse Liquid and

Slurry Jets in Supersonic Airflow", AIAA Journal, Vol.23, No. 12, pp. 1892-1901 (1985).

846
ICLASS-91 Gaithersburg. MD, U.S.A. July 1991 Paper 95

THERMODYNAMICS AND APPLICATIONS OF A SPRAY INTERACTING

WITH A COMPRESSIBLE FLOW

F. Ruiz

Department of Mechanical and Aerospace Engineering

Illinois Institute of Technology

Chicago, IL, U.S.A.

ABSTRACT

The final state of the interaction of a compressible flow with a spray (or solid particle dispersion)

is analyzed, using elementary fluid mechanics and a control-volume approach, in order to detect

variations of the flow Mach niunber at low values of the liquid volume fraction. It was found that the

flow can be choked by effect of the spray in a way analogous to that of a heat addition into the flow.

The effects of downstream pressure changes are discussed. Two applications are presented: as a flow

control device, and as the basis of a ramjet engine.

INTRODUCTION

It is well known that the addition of a solid or liquid phase to a gas can lower the speed of sound

of the resulting two-phase mixture to very low values [1]. This phenomenon has been used to generate

underexpanded air-Uquid jets which diverge on contact with the lower ambient presswe, and cause the

liquid phase to be finely atomized [2]. Critical flow effects are also encountered in flash-boiling

systems, where the separation of vapor bubbles from a Uquid flow can lead to choking of the flow [1].

Ramjet engines rely on the injection of a liquid fuel into a high subsonic or supersonic flow. In the

latter ciise, imlike the others, the volimie occupied by the liquid is negligible (although its mass is not)

compared with that of the gas phase. In most cases analyzed in the literature, however, the liquid

volume fraction is considered to be significantly larger than that of the gas. This paper intends to present an analysis for the opposite case, where the liquid volume fraction is much smaller than that of

the gas, which is the situation encountered in spray systems. We will see how a spray interacting with a

high speed flow can change its Mach number, and open the possibiUty of some interesting appUcations.

DISPERSION OF A SPRAY IN A COMPRESSIBLE STREAM

A gas flows at an initial velocity Vj (and Mach number Mj) through a constant-area duct. From

the initial section (section 1) to the final section 2, liquid particles are added to the flow, so that at

section 2 both phaises have reached the same velocity Fj. Solid particles would have the same effect;

the name "spray" used throughout the paper is imderstood to mean also a dispersion of solid particles.

Now, the addition of pzirticles to the gas alters the speed of soimd, since the particles add to the inertia

of the two-phase mixture, but leave its compressibihty imtouched, if the hquid volume fraction is much

847
smaller than that of the gas. For this limit, and assuming that the particles are small enough to instantly

follow the velocity variations of the gas flow, Wallis [1] gives the following expression for the speed of

soxmd of the two-phase mixture:

where Rg is the original ideal gas const£mt and ms is the liquid mass flow rate divided by the gas mass

flow rate. Equation (1) is based on the assumption that the only effect of the liquid particles is to increment the inertia of the mixture.

When the injecting liquid velocity, Vj^, is other than Vj, there will be momentum transfer between

the phases, and perhaps some mass and heat transfer. To simplify the problem, it will be assumed that

there is no mass nor heat transfer between the phases, no heat nor momentum transfer with the duct

walls. If a solution involving these processes is desired, it can be obtained by adding them at a section

after section 2, using the classical control volume solutions [3]. In other words, it will be assumed that

the two-phase liquid addition and momentum transfer between the phaises proceeds at a much faster

rate than the other processes. The final pressure, velocity, and temperature must be determined by the

conservation laws between sections 1 and 2:

Continuity in the gas phase:

(1)

(2)

Momentum, with friction allowed only between phases:

(Pl-P2)A=ml(V2-ViJ+piAVi(V2-Vi) (3)

Energy, assuming no heat transfer between phases:

PiAVi(CpT2 + ^V^-CpTj.^V]2).,^1(^Vl2.^V^)=0 (4)

Equations (2) to (4) can be put in dimensionless terms using the Mach numbers A/^ and Mj, and

the relative temperature t, defined by:

(5)

^2

(6)

T =

(7)

848
The equations, in dimensionless terms, and substituting continuity into the other equations,

become:

Momentimi:

T(kM2 + Tr)^{l + = (kMi^ms + kMj + j-)^ (8)

Energy:

'•r]^ + A//; = ^ + + Mj}ms (9)

These equations are of second degree in ms and Mj^ (Mj is a fixed parameter, in all that follows).

For each value of ms and M2 there will be two (or none) values of A/^. The same can be said of ms.

Let us first determine the conditions that lead to a final sonic state (M2 = 1). Substituting into

equations (8) and (9) the following relationship is obtained:

(1+k) (2 + (k-1) {Mf^ Ml} ms)) (1 + ms) = (k Mi^ms + k Mj + (10)

Equation (10) is plotted in Fig. 1 for Mj = 0.8 and physically meaningful values of ms (ms > 0).

There is no solution for an injection Mach number Mj^ between Mj and another high value of Mj^,

normally in the hypersonic region. It seems paradoxical that it is possible to reach sonic conditions by

injecting liquid at a velocity Vj (and rate ms = - i or less), but not at a higher speed (which

would tend to increase the final velocity ^2), nor at a higher rate (which would tend to lower the speed

of sound to an even lower value). For these values of Af^, the increase of the speed of sound caused by

the heat generated by friction between the phases is larger than the drop due to the increased inertia of

the mixture (eq. (1)). The paradox can be better understood by replacing M2 in the equations. Given

certain values of ms and A/^, the final Mach number A/2 is given by the following equation (from eqs.

(5) and (6)):

^^^2 + ir]^ i^^L ms^kMi^

^ + M2^ (1 + ms) + A// + Mi^^^s)

(11)

Equation (11) is of fourth degree in M2; for each pair of values of ms and A/£ there are two real

values of A/2 (one subsonic and one supersonic, the other two solutions are not real) or none. Regions

II and rv in Fig. 1 allow two values of A/2, while regions I and III do not allow any real solution. The

two solutions in regions II and IV, A/2 and A/2', are related by:

849
1 + 1 +

yielding:

which binds the different values of the Mach number across a normal shock wave. If one solution is

subsonic, then the other is supersonic. Normally the only physically possible solution is the subsonic

one, because the supersonic solution would involve a spontaneous decrease of entropy, against the

second law, as we shall see later.

In regions I and III there is no solution for It seems, however, that since ms and Mj^ are

independent variables, there should be at least one value of M2, if there is any flow at all. The problem

is similar to that of arbitrary heat addition or flow contraction: a point is reached in which any further

variation in a certain direction does not change the Mach nimiber, which continues at its sonic value,

but will cause the initial conditions to be changed. In the case of this study, any attempt to inject a Uquid with a combination of the parameters ms and Mj^ within regions I and III will result in a choked

exit flow, with M2 = 1, and a decrease of Mj below the original value, if Mj was subsonic, or the

upstream propagation of a shock wave that brings Mj to a subsonic value, if it was originally

supersonic.

It is interesting to see how the choking state will evolve as mj is increased, at a constant Af^ < A/j

(Mj subsonic). First M2 will increase with ms, until M2 = 1 is reached (point "a" in Fig. 2). A further

increase of ms will continue choking the flow and Mj will be decreased. Physically, this represents a

sonic compression wave travelling upstream. The change in Mj will continue xmtil the higher choking

point ("b" in Fig. 2) is reached and Mj = M^. Another increase of ms will maintain Mj = A/^, but the

final state will be supersonic, and the phase interaction fiictionless (point "c" in the figure). The final

pressure will be adjusted by an expansion wave or a system of obhque shock waves, like in a supersonic

nozzle.

If Mj is subsonic and Mj^ > Mj, the exit Mach number will never reach imity: it will start

decreasing d& ms increases, after an initial increase. If Mj is supersonic and Af^ > Mj>, the exit Mach

number will stay supersonic, as long as the outside pressure is low enough to prevent the formation of a normal shock wave.

ENTROPY VARIATION IN ADIABATIC SPRAY MIXING

Another way to look at the choking problem is to consider the change in entropy of the air,

before and after it interacts with the spray, and acquires the same velocity. The variation of entropy of

the gas (the thermodynamic state of the liquid does not change) from 1 to 2, after substituting the

850
appropriate equations, is given by:

^2-^1 k+1 ^2 Mj J

2(k-l) Tj M2 2

' (14)

Substituting sample values for ms, Mj^, and the two corresponding solutions of A/2, one can see

that the subsonic M2 normeilly corresponds to a positive A5, while the supersonic M2 solution normedly

gives a negative A5 and is, therefore, impossible by the second law of thermodynamics. Obviously,

when Mj^ = A/^, the phase interaction is frictionless and we obtain, from equations (8) and (9):

M2 = Mj(l + ms)^/2 (15)

T2=Tj (16)

and, therefore:

AS

Rg

= 0 (17)

Thus, it is physically understandable how the flow will be able to pass from subsonic to supersonic

by adding spray droplets, if these are injected at the same speed as that of the gas flow.

Following a constant M2 line, the entropy variation looks as shown in Fig. 3. For a subsonic M2

value, the entropy change has a zero and a local minimum at = Mj, on the left side of the higher

choking point ("b" in Fig. 2). The conjugated supersonic Mach number, A/2', which gives the same line

in the ms-Mj^ plane, gives a local maximum of the entropy at the crossing with A/^ = Mj, this time on

the right side of the higher choking point. Both curves are separated by a constant offset of value:

^sub ' ^

super

k-1

i + ^A/^^ \] + j,M2^

1 + kM2^ 1 . '4^2^

-hi

1 + kM2^

1 + kM2^

(18)

which is the entropy generated by a normal shock wave, lowering the Mach number from an initied A/2'

> i, to its conjugated A/2 < i, given by eq. (13). It appears, thus, that the solutions in the subsonic

sheet are equivalent to those in the supersonic sheet, followed by a normal shock wave. It makes sense,

since the eissimiptions made for the phase interaction process (adiabatic, conservation of mass and

momentum) can also give a normal shock wave as a possible solution. The subsonic solutions physically

correspond to a supersonic solution followed by a shock wave, but with both processes (mixing and

entropy generation) occurring simultaneously.

Many supersonic A/2 cases would yield AS < 0,'m violation of the second law of thermodynamics.

If the second law is to be preserved, then AS will need to be greater or equal than zero: A/2 > 1 can

only be possible on the line AS = 0 when Mj < 1. For this to occur (given that A/^ and ms are

arbitrarily chosen, and the downstream pressure can be lowered to whatever value is required) the

851
upstream conditions must change, as in the cases of choking. In other words, Mj will remain locked to

the value M^.

Which solution for M2 appears in a real case will be determmed by the downstream pressure. If

the pressure is high enough to support a normal shock wave, then the exit flow will be subsonic. If, on

the contrary, the downstream pressure is low enough, then the shock cannot be supported, and M2 will

be larger than one. Between these, other cases can appear in which the mixing is followed by an obhque shock wave. In the subsonic M2 cases, a variation in the exit pressiu-e will automatically induce

a change in Mjy whether < Mj or not. When A/2 reaches unity, variations in the exit pressure vwll

still be transmitted upstream, so that the choking condition is maintained. Only when Mj has reached

the value Mj = Mj^ will M2 become supersonic. Physically, this can be viewed as the separation of the

shock component of the two-phase interaction, which is decomposed into an isentropic spray

dispersion, followed by a normal shock wave. If the exit pressiue falls below the value required for a normal shock, the isentropic interaction with Mj = Mj^ will be maintained, but followed by a system of

obhque shock or expansion waves, to reach the appropriate external pressure.

In the supersonic Mj and M2 cases, the two-phase interaction will be fully supersonic, followed by

obhque shock or expemsion waves. If the downstream presstu^e reaches the level that would produce a normal shock wave, this shock wave will propagate upstrejun imtil the whole flow is subsonic. A further

increase in the downstream pressure will lower Mi accordingly.

The temperature-entropy locus (t, ^/Rg) of the final state of the interaction, as the relative

hqtiid flow, ms, is increased from zero, for constant values of M^, is plotted in Fig. 4, for Mj = 0.8, and

Mj^ = 0.5, 0.6, and 0.7. The exit Mach number, M2, is subsonic in the upper branches of the curves, and

supersonic in the lower branches. The analogy with the Fanno line (of flow with friction) or the

Rayleigh hne (of flow with heat addition) can be seen easily [3]. The choking condition corresponds to

the maximimi of the entropy generation. Unlike in those classic cases, however, choking is not always

reached as ms increases. For a subsonic Mj^ sUghtly larger than Mj, the flow will always remain

subsonic, M2 reaching a maximum for a certain value of ms and decreasing as ms increzises beyond this

vjdue. For Mj^ = Mj, the flow will be able to pass from subsonic to supersonic, as ms increases, since

the interaction would be isentropic.

APPLICATIONS

The choking effect can be used for flow control in exhaust ducts. Water (or other fluid) can be

injected at low speed (A/^ ~ 0) into the duct. The relative flow of water, ms, necessary to choke the

flow of gas is given by:

ms =

(19)

(l+k)(2+(k-l)Mf)

The liquid flow can control a much larger gas flow (given by Mj) with great accuracy. The relative

spray flow required to control a given gas flow is represented m Fig. 5. This device could by particularly

852
suitable for discharge of hot gases, which could damage other control systems, such as dampers,

Venturis, or critical flow orifices.

Another practical application results from this jmalysis. A spray-driven ramjet engine would be

composed of a diffuser, a spray-mixing area, and a nozzle (optional) as represented in Fig. 6. The

increment of the flow Mach nimiber caused by the spray would be equivalent to that caused by

combustion, which is the method universally applied to generate thrust in ramjet engines. If the Mach

nimiber is increased between the diffuser and the nozzle (not necessarily by heat addition), the engine

will deliver a thrust. An analysis has been performed which shows that the thrust comes from the

kinetic energy of the hquid stored in the structiu-e, with energy conversion efficiencies potentially

higher than 60%. This kind of engine would be advantageous in those situations where a cool exhaust,

without any infrared emissions, would be desirable.

CONCLUSIONS

1. There is an analogy between the interaction of a spray with a compressible gas flow, and heat

addition into a gcis flow: Mach number changes leading to eventual choking are predicted.

2. Unlike in the case of heat addition, a point will be reached where the interaction would be

isentropic, leading to a continuous change of the Mach number from subsonic to supersonic.

3. Applications of this phenomenon to control a subsonic or supersonic exhaust flow and as the

basis of a spray-driven ramjet engine are envisioned.

REFERENCES

1. Wallis, "One-Dimensional Two-Phase Flow", p. 209, McGraw-Hill. 1964.

2. Chawla, J.M., "Atomisation of Liquids Employing the Low Sonic Velocity in Liquid/Gas

Mixtures", Proceedings of ICLASS-85. p. LP/L\/5/l. 1985.

3. Shapiro, A.H., "The Dynamics and Thermodynamics of Compressible Fluid Flow". Ronald,

1953.

NOMENCLATURE

A local section area T gas temperature

C speed of soimd V velocity

Cp specific heat at constant pressure P density

k adiabatic exponent T temperature ratio T2/T1

M Mach number

M' conjugated Mach number subscripts: ms relative hquid/air flow ratio 1 before phase interaction

m mass flow rate 2 after complete phase interaction

P pressure g

of the gas

Rg ideal gas constant L of the liquid

S gas entropy

> conjugated (across a shock wave)

853
854
Entropy

Fig. 4 Lines of constant M in the T-S diagram

855
ms

9B-|

88-

78-

£>e58

48-

38-

28-

18-

8 -8 T r

.8 r.9 Ml

Fig. 5 Relative liquid flow required to choke the gas flow,

vs. inlet Mach number M,

Fig. 6 Schematic diagram of the spray-driven ramjet engine

856
ICLASS-91 Gaithersburg, MD, U.S.A. July 1991 Paper 96

SPRAY CHARACTERISTICS OF PINTLE TYPE INJECTOR

UNDER LOW-PRESSURE FIELD

J. Senda', M. Yamaguchi*, T. Wakashiro^ T. Tsukamoto*,

K. Hojyo^ and H. Fujimoto*

*Dept. of Mechanical Engineering, Doshisha University, Kyoto, Japan

^HONDA R&D Co., Ltd, Tochigi, Japan

*HORIBA Ltd., Automotive Analysis Dept., Kyoto, Japan

§Dept. of Mechanical Engineering, Osaka Institute of Technology, Osaka, Japan

ABSTRACT

The purposes of this study are to clarify the atoraization mechanism of

the spray injected into low-pressure field through a pintle type electronic

control injector, especially the change in spray characteristics dependent on

the surrounding back pressure. In the experiments, the fuel oil is injected

with the relative low-pressure into the quiescent gaseous atmosphere under

the atmospheric pressure. The spray is observed by taking photographs at an

arbitrary time varying the back pressure, using 35mm camera and CCD camera

system. The results show that the back pressure has a great influence on the

spray characteristics. In particular, the saturated vapor pressure of the

fuel oil is the most significant factor. Spray pattern such as the spray

angle, and the atomization characteristics such as Sauter mean diameter of

breakup droplets are almost constant in the range of back pressure from

atmospheric pressure to the vapor pressure. As the back pressure reaches to

the vapor pressure, the vapor bubbles start to grow due to flash boiling.

Further decreasing the back pressure below the vapor pressure results in an

increase in the spray angle and in a decrease in Sauter mean diameter

markedly, since rapid expansion of the vapor phase may them atomize the fuel

jet. Consequently, in the case of back pressure below the vapor pressure,

the characteristics can be explained in terms of the growth rate of vapor

bubbles corresponding with the pressure difference between the back pressure

and the vapor pressure.

INTRODUCTION

In SI Engines, a multi-point injection system, which is controlled

electrically, is predominant at present, owing to high responsibility and

high control quality of fuel flow. However, in some operating conditions, a

great deal of unburnt hydrocarbon is emitted, because of low quality of the

spray atomization. Therefore, it is necessary to reveal spray formation and

atomization processes . Particularly, spray characteristics under the lowpressure field is attractive with reference to the reduced pressure state in

an actual suction manifold.

Conventional studies on fuel injection system for gasoline engines have

dealt with the mixture formation process considering the droplet deposition

857
against the wall in order to improve fuel economy and emission

characteristics [1], And in some cases, fuel flow analysis and the droplet

size distribution for the spray have been reported [2] as to engine

performance. However, no investigations have been attempted to clarify the

characteristics of the spray, injected into lower ambient pressure field than

the saturated vapor pressure of the fuel, where it is necessary to consider

the effect of flash boiling.

When liquid fuel is injected into an ambient pressure lower than the

equilibrium pressure, so-called vapor pressure for the liquid temperature,

the liquid spray disintegrates into fine droplets by partial evolution of

vapor. This flash boiling phenomena occurs when a liquid is rapidly

depressed to a pressure below saturated vapor pressure to initiate a rapid

boiling process. And then, in connection with flash boiling fuel injection,

superheated liquid jets generally have been researched. Brown et al. [3]

have described the relation between the droplet mean size and the growth rate

constant of the bubbles, and Lienhard et al. [4] have presented the breakup

length of the flashing superheated liquid jet. Further, the effects of the

nuclei number and the idle time of bubble growth rate on the atomization

process [5], and research on analytical bubble growth rate for practical

injectors [6] have been reported also.

EXPERIMENTAL APPARATUS AND PROCEDURE

Figure 1 shows a schematic diagram of experimental apparatus. A constant

volume vessel(T)( 170mm x 210mm x 350mm width ) has reinforced glass windows

(2) on both sides to photograph the spray by transmitted light. Fuel was

injected vertically through an electromagnetic pintle type injector(3)(

0.798mm in pintle diameter and 0.931mm in inside diameter of a hole )

installed in the top of the vessel. A pressure inside the vessel, that is

an ambient back pressure P^, could be regulated arbitrarily reducing the

pressure by using a vacuum pump(3« This back pressure was measured by a

vacuum pressure sensor(3). Pressure was applied to a fuel tank(6)by an air

bomb(7)) and this fuel injection pressure P^ was picked up with a pressure

gage(§). And fuel was injected through the injector, operated with a control

circuit(9), during valve opening period t^= 4 ms. at constant.

The photographs of the spray were taken by

CCD Camera system (id) using the transmitted

light of a micro-flash (ij). Thereafter, images

obtained withCCD Camera were recorded on a

image memory Q^, and the spray characteristics,

that is spray pattern and droplet mean diameter

and so on, were analyzed with an image analyzer

equipment (O) via a micro-computer (I3).

In this study, a pressure drop for

injection AP ( AP=P£-Pj^ ) was held constant at

250 kPa. And we introduce a new parameter

defined by the pressure difference APi^^ between

back pressure P^^ and saturated vapor pressure

P^ ( APj^y=P|^-P^ ) to express the intensity of

flash boiling phenomena. Experiments were

conducted at room temperature under the back

pressure P|^ ranged from atmospheric pressure to

the pressure less 'than the vapor pressure P^.

858
Shell-LAWS, n-pentane and n-hexane were used for test fuel at room temperature, T£=20 °C. Fuel properties for each oil were shown in Table 1,

respectively.

Table 1 Fuel properties

Shell-LAWS n-pentane n-hexane

Specific gravity p-^ ( kg/m^ ) 793 626 662

Surface tension 0 x 10^ ( N/m ) 25.1 16.05 18.4

Viscosity y x 10^ ( Pa s ) 121 236 320

Saturated vapor pressure ( kPa ) 62.5 20.5

RESULTS AND CONSIDERATION

![

Fig. 2 Atomization process for Shell-LAWS spray

( AP=150kPa, Pf^=101kPa, t=2.0ms. )

Spray Pattern

Injected fuel from

an injector appears in

the form of a liquid

film with conical shape

near the injector and

the breakup of the film

flow into droplets is

not completed

immediately. In fact,

aerodynamic forces and

the instability of fuel

film cause unstable

wave growth leading to

the disruption of the

film flow, resulting

to liquid ligaments and

to droplets.

Figure 2 shows a unique atomization process for Shell-LAWS spray, under

the conditions of AP=150 kPa, Pjj=101 kPa and t=2.0 ms. Injected fuel forms a

conical shaped liquid film, and thereafter they contract gradually in

progress. Then, the wavy hollow cone jet is formed by the collision of the

film, and it eventually breaks up into relative large droplets from the

circumference of the wavy jet. Finally, in the down stream region, the wavy

flow disrupts leading to large droplets near the spray axis with increasing

disturbance as time passes. In this case, the atomization is suppressed by

the spray contraction.

Figure 3 shows the change in spray pattern with the back pressure Pj^ for

n-pentane, at the time from injection t=3.0 ms. In n-pentane spray,

distinct changes in spray patterns are not observed in the back pressure

range 101 to 61 kPa. At P5=48 kPa, which is below the vapor pressure P^, the

spray atomization is little promoted with contracting spray expanse slightly.

Also, the vapor bubbles start to grow due to the flash boiling inside the

liquid film, and this film length becomes longer a little. On the contrary,

as the back pressure is reduced to 35 kPa, the spray starts to expand in

radial direction again markedly and high atomization quality is revealed,

activated by flash boiling. In this condition, a large number of bubbles are

observed inside the liquid film near the injector. With a further decrease in

859
mm

Fig. 3 Spray patterns with a change in back

pressure ( n-pentane, t=3.0ms. )

the back pressure, the spray angle increases to a greater extent, and it is

found that the liquid spray breaks up promptly into small droplets since the

fuel is flash boiled vigorously at the hole exit of the injector. And then,

the spray vanishes near the injector because of an rapid evaporation of the

small droplets and the film. Thus, it seems to suggest that the spray

characteristics may be assessed in relation to the bubble growth rate, due to

flash boiling, corresponding with the pressure difference AP^^ in the case

of the back pressure below the vapor pressure P^.

In n-hexane spray, spray patterns are almost similar in its properties

to the n-pentane spray, considering the pressure difference AP|^^.

Accordingly, flash boiling phenomena has a remarkable effect on the spray

pattern and the atomization mechanism such as the spray injected into lowpressure field nearby the vapor pressure.

In Shell-LAWS spray, with decreasing the back pressure P^^, the liquid

film length along the spray axis increases and the atomization becomes low

quality. This result appears to be due to the decrease in ambient density.

The film intact length is longer apparently than that of the sprays of npentane over the range of the back pressure independent of the flash boiling.

Spray Angle and Breakup Length

In the following, spray characteristics is described for n-pentane and

n-hexane. In this study, a spray angle 6 and a breakup length 1^ are defined

as shown in Fig. A. The spray angle 0 indicates the cone angle of the liquid

film at the edge of the pintle, and the breakup length 1|^ presents the

distance of the film intact length from the injector.

Figure 5 shows the dependence of spray angle 6 and the breakup length 1^

on pressure difference AP|^y under the condition of AP=250 kPa and t=3.0 ms. For each fuel, 0 becomes almost constant in the region of tsF^^>0, and has a

minimum value near AP|jy=0, then 6 increases markedly with a further decrease

in AP|j attributed to violent flash boiling in the spray, also shown in

Fig.3. Here, the minimum value of 6 near APl^=0 appears to be a transitional

region with the spray contraction owing to the bubble initiation in the film.

Therefore, it seems that the atomization mechanism transits from the process

corresponding to aerodynamical instability to the process controlled by flash

boiling in the case of AP|^^=0. And the value of 0 over the range of AP^^ is

depend on the fuel own properties.

The breakup length 1^^ shows a symmetrical tendency with the spray angle

0 against APj^^. Accordingly, in the region of APjj^<0, Ij^ decreases promptly

860
mm

Fig. 4 Measuring region of fuel spray

-50 -25 0 25 50 75 90

pressure difference APj^^ ( kPa ) Fig, 5 Dependence of spray angle 6 and

breakup length lu on pressure

. ( t=3.

by the effect of flash boiling,

corresponding with an increase in the

spray angle 6. And then it appears

that the bubble growth rate exerts a

great influence on the rapid decrease

in 6 in the APt^ range of less than 0

kPa.

Atomization Characteristics

Here, as the atomization

characteristics, droplet size

distribution and the meandiameter

of breakup droplets are took up for

n-pentane and n-hexane sprays.

Figure 4 illustrates the measuring

region of fuel droplets.

Figure 6 shows the dependence of

Sauter mean diameter 6.-^2 of breakup

droplets on the pressure difference

AP|^^ under the condition of AP=250

kPa and t=3.0 ms. In this figure,

the region drawn with hatched line

indicates the range of impossible

measurement of mean droplet diameter

in presence of very large liquid with

irregular massive state for n-pentane

spray. The results of Sauter mean

diameter d 32 are similar in its

properties to the breakup length 1^

over the range of AP^^.

300

J-1 p

CO

0

-50

O-n-pentane

zVn-hexane

TT

immeasurable area

for n-pentane spray

-25 0 25 50 75 90

pressure difference APj^^ ( kPa )

difference APbv . 0ms . )

Fig. 6 Dependence of Sauter mean droplet

diameter d-^o on pressure

difference APj^^ ( t=3.0ms. )

861
The variation in droplet size distribution for n-pentane spray with

pressure difference AP^^ is revealed in Fig.7. The results indicated that the

distribution was relatively non-uniform having large droplets up to 200 um in

diameter in AP^^^ range from 38.5 to -1.5 kPa. With a decrease in AP^^, the

large droplets reduced gradually the number. Further, in the case of -48.5

kPa in APj^^, it became uniform obviously due to flash boiling, showing the

shape of high frequency of small droplets less than 80 ym in diameter and the

narrow range of the distribution.

Thus, with a further decrease in back pressure P^^ less than the vapor

pressure P^, rapid flashing of grown up cavitation bubbles provides the

uniform droplet size distribution as shown in Fig. 7 and the uniform

distribution of d22 in radial direction in the spray.

100^ ^//^////y/////////////////

C

a 50

>,o

c

0)

cru 0

38.5 11.5 1.5 -1,5 -19.5 -27.5 -41.5 -48.5

pressure difference APf^^ ( kPa )

Fig. 7 Variation in droplet size distribution

for n-pentane spray with

pressure difference APj^^

Fig. 8 Spray outside shape

near the pintle for

n-pentane spray

( t=3.0ms. )

Effect of Flash Boiling

The outline of spray outside near the pintle obtained by the micrographs

illustrated in Fig. 8 for n-pentane spray. In the pressure difference AP^^

range above 0 kPa, the disturbance in liquid film, which is produced at the

pintle edge, has a marked influence in the spray atomization since the

outline of spray outside flows along the pintle shape. With a decrease in

AP|j^ in the range of APjjy<0, the spray expansion, that is the spray angle 6,

decreases gradually due to the spray contraction, not along the pintle edge.

With further decreasing AP|^^, the spray angle 0 markedly increases and the

start point of spray expansion approaches gradually to the injector. From

these results, it seems that the spray atomization is attributed to the

intensity of flash boiling, that is the bubble growth rate and so on, not to

the shape of injector outlet in the range of AP|^y<0 kPa.

Therefore, the behavior of a cavitation bubble growth due to flash

boiling was investigated in this section. The bubble is assumed to grow

spherically, and the Rayleigh-Plesset equation [7] is used to indicate the

growth process. The growth rate of cavitation bubbles is controlled by the

hydrodynamic forces as shown following equation.

RR + -|r'=-^(Pw -Pr ) ( 1 )

862
where, R is the bubble radius, p-|^ the liquid density, the pressure in

liquid at the bubble wall and Pj. is the pressure of liquid surrounding the

bubble. The pressure P^ is defined by the following equation in this study.

Fw -1 V +Pj,„( — ) ^ 4u-^ ( 2 )

where, P^ is the saturated vapor pressure, P the initial pressure of

noncondensable gas inside the bubble, R^ the initial radius of bubble, K the

ratio of specific heat of gas, O the surface tension and ]A is the liquid

viscosity

.

For applying the bubble growth analysis to flash boiling injection, it

is necessary to estimate the pressure profile inside the injector. In this

study, pressures and velocities in steady state were assessed using the

Bernoulli equation with loss terms, considering the cross sectional area at

each point inside the injector.

Figure 9 shows the temporal change in the bubble radius R with each

pressure difference APl^ calculated from equations (1) and (2) for n-pentane

spray, in the case of Rq=15 ym. The bubble radius R increases abruptly with

decreasing AP^^. In particular, in the conditions of AP^y=-41.5 and -48.5

kPa, the bubble nuclei starts to grow inside the orifice of the injector. In

this calculation, we can obtain the bubble growth rate R which appears to be

controlling the spray atomization. Here, the growth rate Rj-^t^) at the

breakup time t^ corresponding with the breakup length 1^ was considered as a

typical rate.

Figure 10 shows the relation between the bubble growth rate Rj-_^^) at t^^ and the pressure difference AP^^ assessed with bubble growth analysis. The

bubble grow rate R^-^^k increases gradually with decreasing the pressure

difference AP|^^. Therefore it is found that vigorous flash boiling phenomena

500

450

400

350

e

V ' 300

pa

CO3 250

•H-o

CO 200 U

0) rH 150

X)

D

100

50

10

COE

U

II

•OS

OJ

4-J CO

x;

o

u

00

CD I—

I

X3

-Q

XI

8

6

R

0 50 100 150 200 250 300

time ( ys ) Fig. 9 Bubble growth calculation

for n-pentane spray

50 -40 -30 -20 -10 0

pressure difference AP^^ ( kPa )

Fig. 10 Relation between bubble growth

rate R and pressure difference

APjjy for n-pentane spray

863
occurs with a decrease in APj^^ from this

bubble growth analysis. And we could

convert the experimental parameter such

as the pressure difference AP|^^ into the

physical quantities such as the growth

rate R in connection with a description

of spray characteristics with flash

boiling.

Finally, several spray characteristics mentioned above are arranged

qualitatively with the bubble growth

rate. Thus, spray characteristics, such

as the spray angle 6, the breakup

length 1^ and the droplets mean diameter

^32' were plotted against the growth

rate as shown in Fig. 11. With increasing

the rate ^t = tb' spray angle

increases rapidly, and the breakup

length 1^ and Sauter mean diameter d22

decrease rapidly, respectively.

Therefore, it seems to suggest that the

spray characteristics in the back

pressure range of less than the vapor

pressure can be described

qualitatively by the bubble growth rate

in terms of flash boiling phenomena.

However, it is necessary to consider the

distribution of the bubble nuclei and

the nucleation process to perform

quantitatively this flash boiling

analysis.

REFERENCES

40

CD 20

21

300

E 200

P^ioo

1

\ O

> (

o

1 s

t r

\ O

) ^ o

C

O

o

( ) o

u

o

oo o

2 A^ . 6 8

bubble growth rate R^--^-^ ( m/ s )

Fig.11 Dependence of spray

characteristics on bubble

growth rate R for n-pentane

spray ( t=3.0ms. )

[1

[2

[3

[4

[5

[6

[7

Nogi,T., Ohyama,Y., Yamauchi,T. and Kuroiwa,H.,

"Mixture Formation of Fuel Injection Systems in Gasoline Engines",

SAE Technical Paper, No. 880558 (1988).

Kashiwaya,M. , Kosuge,T., Nakagawa,K. and Okamoto.Y.,

"The Effect of Atomization of Fuel Injectors on Engine Performance",

SAE Technical Paper, No. 900261 (1990).

Brown, R. and York, J. L., "Spray Formed by Flashing Liquid Jets",

A.I.Ch.E. Journal, Vol.8, No. 2 (1962), pp. 149-153.

Lienhard, J.H. and James, B.D.,

"The Breakup of Surperheated Liquid Jets",

Trans.of ASME, J. of Basic Engineering, (1970), pp.515-522

Suzuki, M., Yamamoto,T., Futagami,N. andMaeda,S.,

"Atomization of Superheated Liquid Jet",

ICLASS '78, 2-2 (1978), pp. 37-43

0za,R.D. and Sinnamon, J.F.

,

"An Experimental and Analytical study of Flash-Boiling Fuel Injection",

SAE Technical Paper, No. 830590 (1983).

Plesset,M.S. , "Cavitation in Real Liquids",

Amer. Elsevrer Pub. Co., (196A), p.l.

864
ICLASS-91

INDEX OF AUTHORS

(Numbers refer to paper numbers)

Abe, K. 62 De Michele, G. 87

Ahmad, C.N. 83,89 Dombrowski, N. 16

Akamatsu, F. 58 Domnick, J. 9

Aoki, H. 53 Dong, Y. 29

Arai, M. 13, 26, 47, 61 Dressier, J.L. 41

Asano, K. 2 Dumouchel, C. 12

Asano, H. 94 Dunn, P.F. 38

Ashgriz, N. 92 Durst, F. 31

Avedisian, C.T. 56 Edwards, C.F. 72

Bachalo, W.D. 44 El-Emam, S.H. 58

Balachandran, W. 83, 89 Elbahar, O.M.F. 84

Bardey, X. 74 Elkotb, M.M. 84

Barry, J.W. 32 Eroglu, H. 75, 78, 93

Barton, S.A. 83 Evers, L.W. 91

Bauckhage, K. C, 23, 51, 52 Farag5, Z. 73

Bellagamba, B. 88 Fargfere, N.M. 14

Bellan, J. 82 Felton, P.G. 67

Beretta, F. 35 Filipovic, I. 25

Bhattacharyya, S. 59 Fritsching, U. 51, 52

Bossard, J. A. 57 Fuchihata, M. 58

Brena de la Rosa, A. 44 Fujii, H. 47

Brunello, G. 71 Fujimoto, H. 96

Callera, S. 71 Fujiwara, T. 4

Camatte, P. 77 Furuhata, T. 53

Care, I. 85 Gakkhar, R.P. 5

Chandrasekhar, R. 59 Gamma, F. 71

Chang, X.-Y. 4 Ghosh, A.K. 59

Chauveau, C. 3 Glancy, C.W. 37

Chesneau, X. 3 Gokalp, I. 3

Chiang, C.H. 80 Gomez, A. 86, 90

Chigier, N. A, 15, 48, 73, 75, 78, 93 Gonzales, M.A. 28

Choi, K.J. 36 Graziadio, M. 87

Chongji, Z. 30 Greenberg, J.B. 33, 81

Chuech, S.G. 20 Grissom, W.M. 69

Coghe, A. 71 Gross, K.W. 93

Corbeels, P.L. 8 Gupta, A.K. 56

Cossali, G.E. 71 Haase, B. 23

D'Alessio, A. 35 Han, Z. 91

Dahl, H.D. 18 Hang, X. 30

Daikoku, M. 53 Harstad, K. 82

Daniels, C.H. 91 Hebrard, P. 74

865
XJiroi xllral,

T1 .IV/frFnrlnnH ivivi^oiiaiiu, /\.iv. A P 10lU

niiiciiicuij Ci.Lj, 43 57 67 Mehra

DS 60

riiroyasu, ti. TV^irTH iVlll^id, iVl.IV. P zj

110) yo, jv. 53

rloy, Jv.iv.

'^7 A^iviitcini XVll^UldJll,

VI . Jo

66 Morelli

F 87o /

II,, null,h J,T -P 27 18iO

riUZaJcWlCZ, '^4 IN agdi

J IN • fxl 04

TflOTYlllfQ T* XlldiiiUKl) 1 • 63 94 Nakabe

K .JO 58

mgeoo, K.u. 54 (y*

irons, \j. 7

24 NieKon TC

A 37

60 Nishida

K 47

IfJVdJl&JVUllClI, anWiinpfi rv. A 46 Novel l^VJVVXXX, H

GvJ .87o / lS.£uaSaWd, 1 .

fO 70 1^JJ

^ciiiicuy, J. 67 OyHemir VJZiii^lXXXX

T , J.. 10

ivcnncy, o.c 43 46

^111111, r%..lJ. 10 Ppnn(Tir»tr*ii X oiiagivjiuu, 1 .

40'ry

Vimn TJ 10 Xian\r xCCK,PIv.xji. F ^7

jvim, 17. J. 17 Pprrv rCiry, \C^.xVl, \A 37

XVllllUiCl, IV* 40 Pietsrh X XWtoVXX,

PXV. i^. A 68

VinonR Do PraVocli

C i^rdKasn, o.

cJ

^IliUlLfC, J. IN.. 7 X iCooCl, v^. 34Jt, 48T^O, JKJ 56

ITncliQn IVllMlall, VJ. \^T'7p\r\Ud^AT

X l^C&Wao, r\.3 .

03yj

^UiaUdyaMll , 1. 69 70 Poftriiinon/ion JvagllUIldJlUdll, IJ.i>. 60

KwnkK -C PnmpcVi XV<lllXWOil,

MX>.XV. P 42

rv. 23 Ppil-7 ivCil^,

PIV.1.J. Fi 28^o

12 77 85 Pnv

R 50

T op

n -Y 27 Pll17

F 65\jj , 057J

17 27 oaury , 1 .1. 84

03 oaQdKdul, xvl. 40

i_<cicuvic, /\.n.. Vi 8 OdIlO, JVl. 40

l^UCKCi, VY . 4^'fJ oaKai, Col/-ot

T1 .

40

T /»vpnHic

YA 40 oalllUClbCll, O.O. 76

DO odlO, iVl. 40

T in ^ P '10 ocnuiic, o. 10ly T ill H <\9 JC ocllcns, QAllf»nc PK.Ww . 1 1 1 T^XU, ill RXJ. YX .XX. H OClIxClJldll, xl.VJ. 56

I-AJXIXCI iVldllg, iVl. 45 oCHUd, J. 06yv

M-AJllly VJ. oo ocnscr, u.

w . Qo

A^ono/^iiT* IVlalliSUUIA , r\. 1 c1 J

onen, CVion

Vi.d.R '^6

MaoC -P 20 55 62 70

Marcano, N. 50 Shimizu, M. 26, 61

Marx, K.D. 72 Silverman, I. 33, 81

Mattachini, F. 88 Simon, 0. 3

McDonell, V.G. 76 Singh, J. 16

866
Sirignano, W.A. OA80

Skyler, PJ. 32

Smith-Wackerle, V.L. 7

SnarsKi, S.R. 38

Sojka, P.E. 68

Sowa, W. 76

Stewart, G.W. 34

Sugiyama, H. 13

Tabata, M. 47

Takada, M. 21, 22

iaicaiiasni, 1.

Tambour, Y. 1, 33, 81

Tanaka, M. 62

Tang, K. 00, 90

Taniguchi, I. 2

Tanno, S. 53

Teske, M.E. 32

Tognotti, l.L. 0000

Tokuoka, N. 21, 22

Tregrossi, A. 35

Trichet, P. 74

Tropea, C. 9, 31

Tsao, K.C. 29

Tsukamoto, T. 96

Uhlenwinkel, V. 51

T T A Umemura, A. 4

T T Ik. T Uno, N. 13

Varacalle, Jr., D.J. /

Vassallo, P. 92

Wakashiro, T. 96

Walter, J.A. 7

Wang, H.Y. 76

Ward, M.J. 57

11 7*11 " A Williams, A. 50

XT 7*1 OA Wilson, S.A. 84

Woods, U.K. j9

Xin, J. Jo

Xu, T.-H. 9, 31

Xu, Y. 29

Yamaguchi, Y. 21, 22

Yamaguchi, M. 96

Yang, O.-Y. 27

Yao, S.C. 79

Yoshizu, F. 64

Yuanji, M. 30

Yule, A.J. 25

Yunbiao, S. 30

Zahavi, S. 1 Zanelli, S. 88

Zhang, F. 21, 22

Zhang, L.-F. 66

Zhou, Q. 79

Zurlo, J. 48

867 *U.S. GOVERNMENT PRINTING OFFICE: 1991-281-618/42749

NIST-114A

(REV. 3-89)

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

BIBLIOGRAPHIC DATA SHEET

1. PUBUCATION OR REPORT NUMBER

NIST/SP-813

2. PERFORMING ORGANIZATION REPORT NUMBER

3. PUBUCATION DATE

July 1991

4. TITLE AND SUBTITLE

Proceedings of the Fifth International Conference on Liquid Atomization and Spray Systems

(ICLASS-91)

S. AUTHOR(S)

Hratch G. Semerjian (Editor)

6. PERFORMING ORGANIZATION (IF JOINT OR OTHER THAN NIST, SEE INSTRUCTIONS)

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

GAITHERSBURQ, MD 20899

7. CONTRACT/GRANT NUMBER

8. TYPE OF REPORT AND PERIOD COVERED

Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (STREET, CITY, STATE, ZIP)

Institute for Liquid Atomization and Spray Systems (ILASS-Americas)

and

National Institute of Standards and Technology

10. SUPPLEMENTARY NOTES

DOCUMENT DESCRIBES A COMPUTER PROGRAM; SF-185, FIPS SOFTWARE SUMMARY, IS ATTACHED.

11. ABSTRACT (A 200-WORD OR LESS FACTUAL SUMMARY OF MOST SIGNIFICANT INFORMATION. IF DOCUMENT INCLUDES A SIGNIFICANT BIBUOGRAPHY OR

LITERATURE SURVEY, MENTION IT HERE.)

This NIST special Publication contains the Proceedings of the Fifth International Conference on

Liquid Atomization and Spray Systems, sponsored and organized by the Institute for Liquid

Atomization and Spray Systems (ILASS-Americas) and the National Institute of Standards and

Technology (NIST), held at NIST, in Gaithersburg, Maryland on July 15-18, 1991. One hundred

papers were presented at the Conference, including four keynote lectures on "Physics of

Atomization", "Experimental and Theoretical Studies on the Structure of Fuel Sprays in Diesel

Engines", "Atomization of Liquid Metals for the Manufacture of Metal Powder and for Spray

Forming", and "Twin Fluid Atomization: Factors Influencing Mean Drop Size". Technical

sessions were organized on Evaporation and Mass Transfer in Sprays, Spray Coating, Basic

Atomization Processes, Pressure Atomizers, Diesel Sprays, Drop Size Distribution of Sprays,

Novel Atomizers, Diagnostics, Atomization of Special Liquids, Spray Combustion, Twin Fluid

Atomization, Spray Modeling, Airblast and Electrostatic Atomization, and Airbreathing and

Rocket Engines.

12. KEY WORDS (6 TO 12 ENTRIES; ALPHABETICAL ORDER; CAPITALIZE ONLY PROPER NAMES; AND SEPARATE KEY WORDS BY SEMICOLONS)

Atomization, Atomization of Special Liquids, Diagnostics, Drop Size Distribution, Electrostatic

Atomization, Fuel Atomization, Pressure Atomizers, Spray Combustion, Spray Modeling, Sprays,

Twin Fluid Atomizers.

13. AVAILABIUTY

X

UNUNITED

FOR OFFICIAL DISTRIBUTION. DO NOT RELEASE TO NATIONAL TECHNICAL INFORMATION SERVICE (NTIS).

ORDER FROM SUPERINTENDENT OF DOCUMENTS, U.S. GOVERNMENT PRINTING OFFICE,

WASHINGTON, DC 20402. ORDER FROM NATIONAL TECHNICAL INFORMATION SERVICE (NTIS). SPRINGFIELD, VA 22161.

14. NUMBER OF PRINTED PAGES

869

15. PRICE

ELECTRONIC FORM

1 1 J.kJ JL Technical Publications

Periodical

Journal of Research of the National Institute of Standards and Technology—Reports NIST research

and development in those disciplines of the physical and engineering sciences in which the Institute

is active. These include physics, chemistry, engineering, mathematics, and computer sciences.

Papers cover a broad range of subjects, with major emphasis on measurement methodology and

the basic technology underlying standardization. Also included from time to time are survey articles

on topics closely related to "the Institute's technical and scientific programs. Issued six times a year.

Nonperiodicals

Monographs—Major contributions to the technical literature on various subjects related to the

Institute's scientific and technical activities.

Handbooks—Recommended codes of engineering and industrial practice (including safety codes) developed in cooperation with interested industries, professional organizations, and regulatory bodies.

Special Publications—Include proceedings of conferences sponsored by NIST, NIST annual reports,

and other special publications appropriate to this grouping such as wall charts, pocket cards, and

bibliographies.

Applied Mathematics Series—Mathematical tables, manuals, and studies of special interest to physicists, engineers, chemists, biologists, mathematicians, computer programmers, and others engaged in

scientific and technical work.

National Standard Reference Data Series—Provides quantitative data on the physical and chemical

properties of materials, compiled from the world's literature and critically evaluated. Developed un- der a worldwide program coordinated by NIST under the authority of the National Standard Data

Act (Public Law 90-396). NOTE: The Journal of Physical and Chemical Reference Data (JPCRD)

is published quarterly for NIST by the American Chemical Society (ACS) and the American Institute of Physics (AIP). Subscriptions, reprints, and supplements are available from ACS, 1155 Sixteenth St., NW., Washington, DC 20056.

Building Science Series—Disseminates technical information developed at the Institute on building

materials, components, systems, and whole structures. The series presents research results, test methods, and performance criteria related to the structural and environmental functions and the

durability and safety characteristics of building elements and systems.

Technical Notes—Studies or reports which are complete in themselves but restrictive in their treat- ment of a subject. Analogous to monographs but not so comprehensive in scope or definitive in

treatment of the subject area. Often serve as a vehicle for final reports of work performed at NIST

under the sponsorship of other government agencies.

Voluntary Product Standards—Developed under procedures published by the Department of Commerce in Part 10, Title 15, of the Code of Federal Regulations. The standards establish nationally

recognized requirements for products, and provide all concerned interests with a basis for common

understanding of the characteristics of the products. NIST administers this program as a supplement

to the activities of the private sector standardizing organizations.

Consumer Information Series—Practical information, based on NIST research and experience, covering areas of interest to the consumer. Easily understandable language and illustrations provide useful background knowledge for shopping in today's technological marketplace.

Order the above NISTpublications from: Superintendent of Documents, Government Printing Office,

Washington, DC 20402.

Order the following NIST publications—FIPS and NISTIRs—from the National Technical Information

Service, Springfield, VA 22161.

Federal Information Processing Standards Publications (FIPS PUB)—Publications in this series collectively constitute the Federal Information Processing Standards Register. The Register serves as the official source of information in the Federal Government regarding standards issued by NIST

pursuant to the Federal Property and Administrative Services Act of 1949 as amended. Public Law

89-306 (79 Stat. 1127), and as implemented by Executive Order 11717 (38 FR 12315, dated May 11,

1973) and Part 6 of Title 15 CFR (Code of Federal Regulations).

NIST Interagency Reports (NISTIR)—A special series of interim or final reports on work performed

by NIST for outside sponsors (both government and non-government). In general, initial distribution is handled by the sponsor; public distribution is by the National Technical Information Service,

Springfield, VA 22161, in paper copy or microfiche form.
U.S. Department of Commerce

National Institute of Standards and Technology

(formerly National Bureau of Standards)

Gaithersburg, MD 20899

Official Business

Penalty for Private Use $300