___ Rockwell International

Rocketdyne Division

6633 Canoga Avenue

Canoga Park, California 91304

RI/RDBS-312

ATOMIZATION AND MIXING STUDY

30 DECEMBER 1985

Contract NAS8-34504

A. FerrenbergU

K. Hunt

3. Duesberg

Advanced Combustion Devices

i APpRovED8Y

Program Manager

Advanced Booster Propulslon Program

PREPARED FOR:

NASA/George C. Marshall Space Flight Center

Marshall Space Flight Center, Alabama 35812
V
FOREWORD

_7

This report describes the work performed during the period of

June 1983 through November 1985 under NASA contract NAS8-34504,

Atomization and Mixing Study. The Rocketdyne Project Engineer for

this work is Dr. Allan Ferrenberg of the Advanced Combustion

Devices group under the direction of Mr. James Lobltz.

Mr. Frank Kirby is the Rocketdyne Program Manager.

Mr. 3oseph Duesberg performed the majority of the atomization

testing, with some assistance from Mr. Ken Hunt. Mr. Hunt performed all of the mixing testing. Technician support was primarily provided by Mr. Gayle Steele. Other Rocketdyne personnel

supporting this work are Mr. Tony Exposito (preliminary droplet

Sizing Interferometry work), Mr. Harry Arbit (basic atomization

research literature review), Dr. Robert Jensen (computational

analyses in support of gas/llquld mixing methods assessment), and

Mssrs. Guido Defever and Robert Saxelby (design support). This

program was performed under the technical direction of Mr. Fred

Braam of the NASA Marshall Space Flight Center.

PRE NOr ALIm 

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CONTENTS

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I. Introduction • .

Liquld/Liquid Element Mixing Study

Gas/Liquld Element Mixing Study

Atomization Study

Appendix A: Liquid Rocket Atomization and Mixing lechnology

Appendix B: Basic Atomization Literature Review

If. Description of Injector Elements

Ill. Liquld/Liquld Element Mixing Study .

Single-Element, Unllke-Trlplet Study

Multlple-E1ement, Like-Doublet Study

IV. Gas/Liquld Mixing Testing

Measurement Capability Assessment_Tests "

Curtain Flow Effects Analysis

Conclusions

V. Atomization

Introduction • ..

Droplet Sizing Interferometry

Test Apparatus and Procedures

Atomization Characterization lestlng

Test Results and Findings

Analysis of Results

Atomization Testing Conclusions and Recommendations

VI. References

Appendix A

Appendix B

Basic Atomization Literature Review

Appendix C

Distribution List for Final Report Atomization

and Mixing Study

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ILLUSTRATIONS

Additional Liquld/Liquld-Trlplet Elements . .

Schematic Drawing of Orifice Assembly Employed in JPL Studies .

Rocketdyne Design to Match 3PL Trlplet-Assembly Drawing

Rocketdyne Design to Match JPL Trlplet-Oriflce Inserts

Basic Plate Design for Box Pattern Injectors

Baseline Elements, Fuel Box Around Oxidizer Doublet,

60° Impingement, 2.5 Scale

Small-Scale Element, Fuel Box Around Oxidizer Doublet,

60 ° Impingement, 1.25 Scale .

Small-Angle Element, Fuel Box Around Oxidizer Doublet,

40 ° Impingement, 2.5 Scale

Reversed Element, Oxidizer Box Around Fuel Doublet,

60° Impingement, 2.5 Scale

Test Apparatus Employed In the Liquld/Llquld Mixing Study .

Liquld/Liquld Mixing Test Apparatus ....

Water Flow Calibration Data for Element lO Inner Orifice

TRIC Flow Calibration Data for Element lO Outer Orifice Palr

Overall Mixing Efficiency vs Velocity Heat

Ratio for All Test Runs

Overall Mixing Efficiency vs Momentum Ratio for All Test Runs

Overall Mixing Efficiency vs Rupe Parameter for All Test Runs

Overall Mixing Efficiency vs Elverum-Morey

Parameter for All Test Runs

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Fluid Cavitation

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Repeatability Tests

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Collection Distance

Overall Mixing Efficiency vs Collection Distance for

Combined Rocketdyne and AeroJet Data

Overall Mlxlng Efficiency vs Elverum-Morey Parameter -

Effect of Diameter Ratio (O-F-O Configuration)

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Diameter Ratlo (F-O-F Configuration)

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Overall Mixing Efficiency vs ElverumMorey Parameter -

Effect of Orifice Scaling

Overall Mixing Efficiency vs Elverum-MoreY Parameter -

Effect of Impingement Angle .

Overall Mixing Efficiency vs Elverum-Morey Parameter Effect

of Injection Velocity at Constant Velocity Ratio (Throttling)

Overall Mixing Efficiency vs Elverum-Morey Parameter

Effect of Fabrication Method

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Rocketdyne vs JPL Data

Data Correlation for Liquld/Liquld Triplet Mixing Study

Schematic Drawing of the Four Cold-Flow Models

Cold-Flow Gas/Liquld Mixing Measurement System

New Gas/Llquld Mixing Measurement Test Apparatus

Gas/Llquld Mixing Measurement Capability

Assessment Test Apparatus

Computational Mesh

Velocity Profiles

Computed Mass Flux Properties

Injected Species Mass Fraction Contours

Injected Species Mass Fraction Contours

Gas Motion - With and Without Curtain Flow

Hlgh-Pressure Atomization Test Vessel

Test Vessel Reclrculatlon Control Devices

Intrusive Probe

Droplet Sizing Interferometry

Droplet Visibility

DSI Measurement lest Facility Configuration

lypical Data Sheet

DSI Output Plots

DSI labular Output

Typical Data Compilation Program Output

Spray Mapping coordlnate System

Alr Motion Tests, Hlgh Liquid Flow Rate Slze Plot (1)

Alr Motion Tests, High Liquid Flow Rate Slze Plot (2)

Alr Motion Tests, Hlgh Liquid Flow Rate Size Plot (3)

Wlth No Curtain Flow

With Curtain Flow

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Air Motion Tests, High Liquid Flow Rate Size Plot (3)

Alr Motion Tests, Low Liquid Flow Rate Size Plot (1)

Alr Motion Tests, Low Liquid Flow Rate Size Plot (2)

Alr Motion Tests, High Liquid Flow Rate Velocity Plots (1) .

Alr Motion Tests, Hlgh Liquid Flow Rate Velocity Plots (2)

Alr Motion Tests, Hlgh Liquid Flow Rate Velocity Plots (3)

Alr Motion Tests, High Liquid Flow Rate Velocity Plots (4)

Air Motion Tests, Low Liquid Flow Rate Velocity Plots (2)

Spray Droplet-Slze Distribution, Test l

Spray Droplet-Slze Distribution, Test 2

(0.508 cm point spacing)

Spray Droplet-Slze Distribution, Test 2

(l.Ol6 cm point spacing) .

Droplet Velocity Distributions, Test 2, Location (0, O, 23.5)

Droplet Velocity Distributions, Test 2, Location (0,-4.6, 23.5)

Spray Droplet-Size Distribution, Test 3 .

Inner Core Droplet-Slze Distribution, Test 3

Outer Annulus Drop-Size Distribution, Test 3

Spray Droplet-Slze Distribution, Test 4 . . .

Droplet Size vs Velocity at (0, O, 23.5), Test 4

Positive X-Axls Spray Droplets-Slze Distribution, Test 6

Positive Y-Axls Spray Droplet-Slze Distribution, Test 6

Posltlve X-Axls Spray Droplet-Size Distribution, Test 7

Positive Y-Axls Spray Droplet-Size Distribution, Test 7

Spray Droplet-Size Distribution, Test 9 .

Spray Droplet-Slze Distribution, Test lO

Spray Droplet-Size Distribution, Test 13

Spray Droplet-Size Distribution, Test 14

Inner Core Droplet-Size Distribution, Test 15 .

Outer Annulus Droplet-Size Distribution, Test 15

Spray Droplet-Size Distribution, Test 17

Spray Droplet-Size Distribution, Test 18

Spray Droplet-Size Distribution, Test 21

Spray Droplet-Size Distribution, Test 22

Spray Droplet-Size Distribution, Test 19

Spray Droplet-Size Distribution, Test 20

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TABLES

Baseline LOXlHydrocarbon Injectors

Objectives of the Liquld/Llquld-Trlplet Mixlng Study

Llquld/L1quld-Trlplet Test Matrix • •

Additional Tests Conducted to Assess Cavitation Effects . •

Results of Completed Liquld/Liquld-Trlplet Test Matrix

Summary of Variations in Orifice Contouring

Additional Tests Conducted to Test Fabrication Method

Results of Additional Fabrication Method Tests

Summary of Available Data from 3PL Triplet Studies

Design Specifications for Proposed LOX/RP-I Booster Injector

Design Specifications for Cold-Flow Models

Results of the Test Series for Multlple-Element

Llke-Doublet Models . •

Collection Efficiency Results . • •

Atomization Test Matrix • • •

Multlcomponent Liquids and Their Properties

Atomization Test Conditions - Gas/Llquld Injectors .

Llke Doublet (Wax Data Comparison) Test Conditions

Atomization Test Results

Injection Cond_tlons Investigated and SSME Main

Combustion Chamber Conditions • •

• II-2

III-6

• III-lO

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III-16

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• III-37

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D

I. INIRODUCIION AND SUMMARY

The primary objective of this Atomization and Mixing Study was the obtainment of

atomization and mlxing performance data for a variety of typical LOX/hydrocarbon

injector element designs. Such data are required to establish injector design

criteria for such elements, and to provide critical inputs to liquid rocket

engine combustor performance and stability analysis, and computational codes and

methods. For the most part, these results are sufficiently generic to allow

their application to similar injectors employing other propellant combinations.

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This program began in February 1982. During the first year of this effort, a

literature search and compilation of the liquid rocket injector atomization and

mixing data and correlations were performed. Two sets of mixing tests were also

performed during the first year: a set of gas/llquld element mixing characterizatlon tests and a smaller set of llquld/liquld triplet element mixing characterization tests. First-year work was reported in an interim technical report (Ref.

l-l). The summary of the liquid rocket injector atomization and mixing correlation and data contained in Ref. l-l is presented herein as Appendix A.

During this first year, deficiencies and problems with the atomization test

equipment were identified, and action was initiated to resolve them. While these

efforts were not a part of this contractual program, they did result in delays

and test plan modification in the atomization testing of this program. In addltion, the test results of the gas/llquld mixing tests Indlcated that an assessment of these test methods was required. Finally, the llquld/llquid triplet

testing performed Indlcated a need for a more extensive set of such tests. As a

result of these issues, an extension and several modifications of this program

were implemented. This work was accomplished during the period of July 1983

through October 1985, and is described in detail in this report.

From October 1983 through December 1984, the gas/llquld mixing assessment methods

were analytically and experimentally assessed. During the period of June 1984

through November 1985, several series of llquld/llquld element mixing tests were

performed. Finally, in December 1984, after several years of noncontract

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testing, problem resolution, and technique development, contract testing to

establish the atomization characteristics of selected elements began. This

effort culminated in a series of atomlzatlon tests durlng the summer and fall of

1985. All of these efforts and their results are described in detail in this

report.

v

This final report consists of five sections and two appendices:

Introduction and Summary

I/ Description of Injector Elements

Ill Liquld/Liquld Element Mixing Study

IV Gas/Liquld Element Mixing Testing

Atomization Study

Appendix A. Liquid Rocket Atomization and Mixing Technology

Appendix B. Basic Atomization Literature Review

V

Section I/ describes the injector elements constructed and tested as a part of

this program. The remaining sections are separately summarized in the remainder

of this section.

LIQUID/LIQUID ELEMENI MIXING STUDY

The objective of the liquld/llquld mixing study was the acquisition and correla--

tlon of cold-flow mixing data for LOX/hydrocarbon injectors. In the primary

phase of this study, a series of 71 liquld/llquld mixing tests were performed at

Rocketdyne durlng the period from June 1984 to October 1985. Ten slngle-element,

unlike-triplet injectors were used in a test matrix designed to determine the

effect on mixing efficiency of variations in liquid density, impingement

distance, element geometry, orifice characteristics, flow rates, and collection

RI/RDBS-312

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distance, for the overall purpose of evaluating existing triplet injector design

correlations and developing improved correlations where necessary.

A secondary phase of the llquld/llquld mixing study was conducted to evaluate the

mixing characteristics of a candidate llke-doublet injector pattern proposed for

use in large LOX/hydrocarbon boosters. A'serles of tests were performed in

October 1985, using four different injector models. In this case, the models

featured a multlp]e-element "unit cell" configuration representative of the overall llke-doublet pattern. The test matrix was designed to determine the effect

on mixing efficiency of propellant interchange, mass flow throttling, impingement

angle, and model scaling.

This program is the largest, and most thorough and comprehensive investigation of

the mixing performance of llquld/llquld-trlplet elements that has ever been per

formed. The findings and conclusions of this effort are, in some cases, contradictory to previous findings (based on much less comprehensive test programs) and

traditional design practice. Thus, these findings are of great importance and

should result in considerable change and improvement in triplet injector designs.

In accordance with these findings, the followlng design criteria and advice to

designers of liquld/liquld triplet elements are provided.

l , The use of an optimum value of the Elverum-Morey parameter as a design

criterion for llquld/llquld triplets is not Justified and should be discontinued. Mixing efficiency increases as the Elverum-Morey parameter

increases.

2. The more dense propellant should be injected from the outer orifices.

3. Small, outer to inner, orifice diameter ratios (e.g., 1 or less) are not

recommended.

no lhe changing of the injection velocity of both liquids by the same per

centage, has no effect on the mixing performance, over the range of

injection velocities tested.

RI/RD85 312

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5. An increase of the outer to inner veloclty ratio of the propellants

improves mixing performance.

6. The Impingement angle has no significant Influence on mixing and should

be chosen on the basis of other considerations.

7. As a guide to use in the tradeoff of various design considerations, it

Is recommended that the designer attempt to maximize the parameter.

t ot tvo/ 0.25

All of these findings, conclusions, design guidance, and especially the preceding

mixing efficiency design parameter, should not be extrapolated or applied beyond

the limits of the range of the variables tested in thls program without very

careful consideration. The range of the variables tested is:

Outer to inner orifice velocity ratio

Outer to inner orifice diameter ratio

Outer to inner orifice density ratio

Impingement angle (between outer

streams)

Orifice diameter

Injection velocities

0.37 to 1.69

0.92 to 1.58

0.76 and 1.32

30 to 90 degrees

O.ll7 to 0.236 cm (0.046 to 0.093 In.)

6.7 to 20.2 m/sec (22 to 66 ft/sec)

The llm!ted testlng performed on the llke-doublet, multlple-element injector supports the validity of the scaling methods employed. Throttling appeared to have

no effect on injector mixing performance. Poorer mixing efficiency was observed

for a smaller impingement angle/longer impingement distance variation on the

baseline injector.

v

GAS/LIQUID ELEMENT MIXING STUDY

Slngle-element, gas/11quld mixing measurements were to be, and have been, performed as a part of thls atomization and mixing study. Because of the very poor

RI/RD85-312

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V

collection efficiency measurements obtained on tests already performed, it became

necessary to assess and improve the methods by which such measurements are made.

Accordingly, a series of studies and experiments were performed, resulting in a

number of changes in the apparatus and measurement methods. These changes culminated in a demonstration of the ability to make moderately accurate gas/llquld

mixing measurements at low pressure, in lightly loaded (with liquid) flows, and

in relatively uniform, one-dlmenslonal flows. Whether these techniques can be

successfully applied to the much more heavily loaded and hlgh pressure Flows of

Interest, is not known. However, it is known that the flow must be relatively

straight, (i.e., reclrculatlon must be prevented).

E_

The technique commonly employed for the prevention of reclrculation is the use of

large quantities of purge gas flow (also referred to as base bleed or curtain

flow) circumferentlally about the injector. This technique has been applied in

all previous gas/llquld mixing measurements. However, while the purge gas

greatly reduces or prevents reclrculatlon, it also affects the dispersion of the

fuel gas (and perhaps, though certainly to a much lesser extent, the liquid disperslon). Thus, a situation exists where the purge gas is necessary to perform

the measurements. At the same time, however, it disturbs or changes the mass

flux distributions being measured.

F

_i _

While this problem was recognized in the past, means to assess the magnitude of

the effect of the purge gas on the fuel gas distribution did not previously

exist. Sufficiently complex, gas dynamlc/computer codes capable of estimating

the magnitude of this effect now exist. Such an analysis was performed, lhls

analysis indicated a large effect of purge gas on fuel gas distribution. As a

result of these studies, the following are concluded:

I. Single-element gas/llquld mixing measurements may be of limited value

for comparing the mixing performance of different elements or types of

elements, and also ma_a__b__ee of very limited value for assessing the relative effects of injector geometry or flow variables. Such measurements

serve only as a relative comparison of mixing efficiency.

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2. The use of such measurement data as input to the performance analysis

codes (e.g., SDER), or to establish design criteria (e.g., optimum

values of the Elverum-Morey parameter), is not Justified. Further testing of this type Is not recommended.

V

3. Effort should be directed toward the development of a means to assess

and measure injector mixing performance.

One very promising means by which thls may be accomplished is through the use of

multidimensional CFD codes such as the Advanced Rocket Injector Combustor Code

being developed under NASA/MSFC Contract NAS8-34928. Such codes can already

model cold-flow gas motion with good accuracy and can Include the effects of

droplets and combustion on gas dispersion. Modeling of the llquld phase (e.g.,

atomization, stream and droplet breakup, and droplet motion) is currently less

developed, but efforts to Improve such models are under way.

Another means for the measurement of the mixing performance of injectors Is the

utilization of advanced combustion diagnostic techniques such as Raman spectroscopy. These diagnostic techniques offer the potential capability to directly

and nonintruslvely measure combustor gas temperatures and compositions. Thus,

they could provide the first direct measurements of hot-fire mlxlng efficiency.

V

ATOMIZATION STUDY

The objective of thls task was the development of a body of information and

empirical correlation by which the atomization characteristics of typical LOX/

hydrocarbon injectors could be assessed or predicted. The survey of the state of

the art In thls area, presented in Appendix A, discusses the great need for such

information and the limited quality and, especially, the applicability of the

available data.

To obtain such data and improve upon the droplet-slze measurement techniques of

the past, the new and very promising technique of droplet sizing Interferometry

(DSI) was employed. Unfortunately, the application of thls powerful new

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V
i

technique was not as straightforward and simple as originally anticipated. Problems with the DSI, an inability to measure droplet sizes in hlgh-droplet-denslty

sprays, great difficulty in controlling spray reclrculatlon at high pressure,

continual difficulties with the monodlsperse droplet generator (a DSI calibration

tool), various shortcomings of the hlgh-pressure test vessel, and many other

problems (as described in Section V) occurred. All of these problems were

resolved at no expense to this contract prior to the start of these atomization

tests. However, in many cases, the solutions to these problems resulted in a

decrease in the scope, quality, or quantity of the atomization testing that could

be accomplished during this program. Nevertheless, this effort resulted in the

most detailed and complete measurements of the structure of the sprays produced

by several of the injector elements of interest. While this effort consisted

primarily of tests of coaxial elements, limited testing of gas/llquld-pentad and

-triplet elements, and a llke-doublet liquid element, was also performed. This

program demonstrated the capabilities of the DSI and resulted in the development

of procedures and methods for acquiring, compiling, and correlating the vast

quantities of data obtainable with the DSI. Specific findings and results from

this effort are:

I. Information is provided that indicates the hot-wax technique data, and

especially the correlations relating the effects of liquid properties on

droplet size, may be of questionable validity or applicability.

2. The Lorenzetto and Lefebvre droplet-slze correlating equation is not

recommended as a means to estimate droplet sizes for liquid rocket

coaxial injectors, especially at mixture ratios greater than I.

3. The Kim and Marshall droplet-size correlating equation is recommended

for the very rough estimation of droplet sizes for liquid rocket coaxial

injectors operating at high (main chamber) mixture ratios, lhis Is

being recommended only because there appears to be no alternative.

Based upon the results of these atomization tests, it is further recommended that the droplet sizes computed from the Kim and Marshall equation be increased by about 30 percent.

RI/RD85 312

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4. It is strongly recommended that future efforts to acquire atomization

data obtain such data at hlgher pressures and flow rates.

, The droplet-size measurement technique employed In thls program was the

visibility/Intensity (V/I) OSI technique. This second-generation DSI

technique is inferior to two new DSI techniques now commercially available. However, all of these DSI techniques, as well as all other nonintrusive droplet-size measurement techniques available, have deficiencies that limit their applicability to the study of liquid rocket injector atomization. These deficiencies are the lnablllty to measure droplet sizes In the very dense sprays typical of liquid rocket injectors at

nominal flow rates (especially the very fine and dense sprays of gas/

liquid injectors), and the inability to distinguish droplets by composltlon (required to assess unlike liquid injector atomization). Improved

capabilities In these areas are needed.

, The findings of this study demonstrate the need for considerable additional effort In the study of atomization. The validity, and especially

the applicability, of all available data and correlations are questionable. The technical challenges and problems are great, but the need for

thls information to support injector design efforts and the rapidly

developing and very promising field of spray combustion modeling Is also

great.

. Finally, it Is strongly recommended that users of any atomization data

and correlations become famlllar with the quality, validity, and applicability of them. Furthermore, reports of any analyses based upon such

data or correlations should clearly state the limitations and potential

errors assoclated wlth the utilization of such atomization data.

Additional results and findings are presented In Section V.

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V
APPENDIXA: LIQUID ROCKETATOMIZATIONANDMIXINGTECHNOLOGY

Thls appendix consists of the survey results of the liquid rocket atomization and

mixing data, correlations, and measurement techniques previously reported In

Ref. l-l. This appendix is an exact duplicate of that survey as published in

July 1983, in Ref. l-l, and is included hereln for the convenience of the reader.

It should be noted that certain findings of that survey, regarding mixing assessment techniques and correlations, have been found to be incorrect and have been

supplanted by the more recent findings of this program as described in Sections

III and IV of this report.

APPENDIX B: BASIC ATOMIZATION LITERATURE REVIEW

Available information on liquid atomization by rocket englne-type injector elements is presented In Ref. l-l and Appendix A, In which the literature on atomization by llke-doublet, triplet, pentad, and coaxial injector elements was summarized, discussed, and assessed. The general conclusions of that summary were

that reported atomization data are largely empirical and ad hoc, only qualitatively understood, and of little general validity or utility. This new review,

presented in Appendix B, covers the literature on the more basic or theoretical

aspects of liquid atomization. This effort Is primarily directed toward studies

related to droplet deformation, drag, and breakup, as these processes tend to

influence the ultimate size and motion of droplets and are of great importance in

efforts to model sprays. Certain more basic and general atomization studies for

alrblast atomizers are also included. The importance of the atomization process,

particularly In combustion application s , has resulted in the publication of hundreds of papers and reviews concerned with various aspects of these subjects. A

selection of these studies, representing classical and current procedures,

results, and theories, is summarized In this appendix. Thls summary, together

wlth that in Appendix A, provides a complete description of the state of the art

of atomization as it applies to liquid rocket engines. The summary should serve

as a useful reference to those familiar with this area and as a basic introduction for those entering this field of study.

h

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r _

V Y
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II. DESCRIPTION OF INJECTOR ELEMENIS

lhis section presents the details of the injector elements designed, constructed,

and tested as a part of this program. Early in this program, nine typical

LOX/hydrocarbon elements were selected for atomization and mixing testing. These

were all triplet, pentad, or coaxial elements. The primary characteristics of

these baseline elements are presented in fable 2-I. Each of the impinging elements was constructed in the form of a small disc (approximately 2.25 inches in

diameter by 0.75 inch thick). A high-pressure mounting assembly was constructed

to support and provide liquid and gas manifolding for the interchangeable element

discs. The coaxial elements could also be utilized with this mounting assembly.

Detail design drawings of the mounting assembly and these nine baseline elements

are presented in Ref. l-l.

lhe llquld/liquld m_xing testing originally planned to be accomplished during

thls program was very extensively, and successfully, increased, lhls testing

required the construction of additional injectors. The first set of these injectors consisted of eight new liquld/llquid-triplet, slngle-element discs. These

are elements lO through 17. Detail design drawings of these element discs are

presented in Fig. 2-!.

0

Io provide a basis for comparison of the results of these liquld/liquid-triplet

tests and earlier Jet Propulsion Laboratory (JPL) tests, a triplet element

employing the JPI_ design practices was constructed. To the greatest extent possible, every important feature of the JPL triplets was incorporated into this

element. A JPL triplet orifice design is presented in Fig. 2-2, and the

Rocketdyne "copy" of this design is presented in Fig. 2-3 and 2-4. Fig. 2-3

shows the overall assembly, and Fig. 2-4 presents the details of the injector

'orifice inserts.

The final injectors designed to support the expanded liquid/liquid mixing test

program consisted of a set of multiple element, like doublet injectors. This

injector pattern was identified under a Rocketdyne IR&D program as a likely candidate for a large I_OX/RP-I booster. This design was based on the concept of a

"box" of like doublets of one propellant surrounding each like doublet of the

e'RECF.I)iNG PAGf BL.kJNIK NOT

RI/RD85- 312

11-I
• • • _f
5

- I

I

"i

i _:_

i11.114

I

!

i

-- !

ORIGINAL PAGE IS

OF POOR _,,_'.,TY

PATTERN

1 TRIPLET

PB

(FOF)

2 TRIPLETI

PB

(FOF)

3 PENTAD

PB

(FOF)

4 TRIPLET

PB-EDNI

(FOF)

5 COAXIAL

PB

6 TRIPLET

PB

(FOF)

7 COAXIAL

MC

8 TRIPLET

MC

(OFO)

9 COAXIAL

MC

PROPELLANT

LOX/RP-1

(LIQUID/LIQUID)

LOX/CH 4

(GAS/LIQUID)

LOX/CH 4

(GAS/LIQUID)

LOX/CH 4

(GAS/LIQUID)

LOX/CH 4

(GAS/LIQUID)

LOXIC3H B

(LIQUID/LIQUID)

LOX/CH 4

(GAS/LIQUID)

LOX/RP-1

(LIQUID/LIQUID)

LOXIC3H B

(GAS/LIQUID)

PC Ic

(PSIA) (°R) MR

3500 2100 0.44

2-/28 2100 0.49

3500 2100 0.49

3500 2100 0.44

3500 2100 0.49

5250 1860 0.40

3000 6400 3.5

2000 5900 2.8

4000' 6400 3.0

TABLE 2-1. BASELINE LOX/HYDROCARBON IN,

AP0

(PSI)

700

700

700

600

700

505

I000

350

1000

INJECTOR ELEMENT SELECTI(

ORIFICE

DIAMETER

(INCHES)

APF

(PSI) OXIDIZER

700 0.0447

346 0.045

700 0.0712

850 0.016

700

905 0.05

400

700 0.065

8OO

FUEL

0.055

0.063

0.0587

0.027

0.08

0.050

OXII

0.:

'

O _

0.(

.

.

, ii

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l

0

IECTORS

ORIFICE

DIAMETER

(INCHES)

IZER FUEL

IMPINGEMENT

ANGLE

50

,00

74

165

_5

_5

0.239 60 °

0.276 50 °

0.248 60 °

0.062 60 °

0.40 60 °

0.35 60 °

,'Ri

Vo

FT/S

285

151

144

229

75

2O/

lO0

171

lO0

ORIGINAL PAGE IS

Of POOR QUALITY

VF

FT/S

290

402

554

595

634

188

500

288

600

FREE STREAM

LENGTH

(INCHES)

OXIDIZER

0.250

0.250

0.250

0.098

0.25

0.29

FUEL

O.289

0.275

0.289

0.112

0.29

0.25

REMARKS

REF: RI/RD81-129

(NAS8-33243)

REF: RI/RDBl-129

(NAS8-33243)

REF: RI/RDBl-129

(NAS8-33243)

REF: IR&D

REF: RI/RDBl-129

(NASB-33243)

ELVERUM-MOREY

= 0.66

REF: RI/RD79-278

(NAS8-33206)

ELVERUM-MOREY

= 0.66

SIMILAR TO

ELEMENT 7

OXIDIZER_

RECESS (Ii

0.12

O.l

0.23

FQU)O 

ii

COAXIAL ELEMENTS

ORIGINAL PAGE IS

POOR QUALITY

!

!

|

|

|

i

i

POST

WCHES)

INNER TUBE

ID (INCHES)

0.088

0.182

0.20

INNER TUBE

OD (INCHES)

0.1355

0.202

0.23

OUTER TUBE

ID (INCHES)

0.1735

0.247

INNER TUBE

LENGTH

(INCHES)

2.3

2.53

0.28 2.5

OUTER TUBE

UNDISTURBED

FLOW LENGTH

(INCHES)

0.525

0.5

0.5

INNER TUBE

ORIFICE

DIAMETER

(INCHES)

0.0455

0.086

0.091

.....

|

= _ |

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RI/RD85-312

II-2

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Figure 2-2. Schematic Drawing of Orifice Assembly Employed in JPL Studies

RI/RD85-312

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other propellant. These inJectbrs were constructed from two plexlglass plates

bonded together. This novel deslgn concept allowed relatively simple and

Inexpensive fabrication of these cold-flow Injectors. Detalls of the deslgns of

these Injectors are presented in Fig. 2-5 through 2-9. Fig. 2-5 presents the

baslc plate deslgn employed for all the Injectors, whlle Fig. 2-6 through 2-9

present the major detalls of each of the Injector deslgns. Rationale for these

deslgns and addltlonal descriptions of these element patterns can be found In

Section III.

v

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Figure 2-7. Small-Scale Element, Fuel Box Around Oxidizer Doublet, 60 ° Impingement, 1.25Scale
L

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Figure 2-9. Reversed Element, Oxidizer Box Around Fuel Doulet, 60 ° Impingement, 2.5 Scale

Ill. I_IQUID/LIQUID ELEMENI MIXING STUDY

The objective of the llquld/llquld mixing study was the acquisition and correlation of cold-flow mlxlng data for LOX/hydrocarbon injectors. In the primary

phase of thls study, a series of 71 llquld/llquid mixing tests were performed at

Rocketdyne during the period from 3une 1984 to October 1985. Ten slngle-element,

unl!ke-trlplet injectors were used In a test matrix designed to determine the

effect on mlxlng efficiency of variations In liquid denslty, impingement geom

etry, orifice characteristics, and collection distance, for the overall purpose

of evaluating existing triplet injector design correlations and developing

improved correlations where necessary.

A secondary phase of the liquld/llquld mlxlng study was conducted to evaluate the

mixing characteristics of a candidate llke-doublet injector pattern proposed for

use In large LOX/hydrocarbon boosters. A series of II tests were performed In

October 1985, using four dlfferent injector models. In this case, the models

featured a multlple-element "unit cell" configuration representative of the overal] llke-doublet pattern. The test matrix was designed to determine the effect

on mlxlng efficiency of propellant interchange, mass flow throttling, impingement

angle, and model scaling.

SINGLE-EI_EMENI, UNLIKE-1RIPLET STUDY

The efficiency of combustion attained in a rocket engine is highly dependent upon

the unlformlty of the mlxture ratio produced by the injector. Most engines are

designed to operate at or near overall mixture ratios corresponding to the maxl

mum of the theoretical performance curve. Since any devlatlon from the target

mixture ratio can result in lower performance, it Is evident that not only over

all mixture ratio, but also local mixture ratio must be accurately controlled in

order to achieve the highest possible performance from the englne. The objective

of successful injector design Is therefore to deliver the propellants to the combustlon zone in a manner such that local variations In mlxture ratio across the

InJectant spray field are mlnlmlzed wlth respect to the overall mixture ratio.

RI/RD85-312

Ill-1
The accepted parameter that defines the mixture ratio uniformity of an lnJectant

spray field was proposed several years ago by Rupe at JPL (Ref. 3-1). Known as

the overall mlxlng efficiency, or Em (E-sub-m) value, this parameter is computed

as a mass welghted summation of the mixture ratio variations exhibited by a number of samples collected over a surface in the spray fteld. The mixing efficiency

equation takes the form subsequently shown; each local sample having a mixture

ratio different from the overall value is treated as a decrement from perfect

mixing, and weighted by the fraction of the overall collected mass it represents.

Em = lO0 I l- i_MFsb R-Rsb R (i)

In Eq. l, Em is the overall mixing efficiency of the element, R represents mixture ratio expressed as oxidizer mass divided by total mass, and MF Is the fraction of the total collected mass contained In each individual sample. The subscript sa indicates a sample whose mixture ratio Is above the overall value,

while sb indicates one whose mixture ratio Is below the overall value. Mixing

efflclencles range from zero to lO0_, wlth I00% implying that all samples tested

have the same mixture ratio, and zero implying that all samples consist of either

one component or another.

v

The problem of optimizing mixing efficiency for various llquld-on-llquld impinging injector element types was Investlgated by Rupe and his colleagues at 3PL,

(Ref. 3-I), resulting In injector design correlations that have become widely

accepted over the years. These studies were among the first to document the

relationship between mixing uniformity and certain hydraulic and relative momentum conditions of the impinging streams. For the case of the coplanar liquid/

liquid triplet, the definitive 3PL study was performed by Elverum and Morey (Ref.

3-2), culminating In the correlating parameter bearing their names. All JPL

investigators cited the importance of liquid stream stability (i.e., stable, symmetrical velocity profile; similar centerllne to mean-stream pressure ratio; and

low degree of radial spreading) as a prerequisite for consistent Implngement and

subsequent mlxlng characteristics. Thls emphasis on producing perfectly impinging

turbulent streams led to thelr use of preclslon-machlned test orifices having

elllptlcally contoured entrances, roughened turbulence-lnduclng sections, and

overall lengths of at least 20 diameters.

RI/RD85-312

III-2
In the Elverum-Norey study, the propellant stmulant combinations employed were

water/kerosene and water/carbon tetrachlortde. In reporting the results, the

authors stated that by interchanging the two liquids tn each slmulant pair

between the inner and outer orifices of a particular element, they were able to

investigate the effects of density ratio (po/Pl) variation over a range

from 0.54 to 1.85. Based on densities of 0.82, 1.00, and 1.59 g/cm 3 for

kerosene, water, and carbon tetrachlorlde, respectively (Ref. 3-2), the four

discrete den- slty ratio possibilities for the reported slmulant combinations are

more accur- ately seen to be 0.63, 0.82, 1.22, and 1.59. Other variations

included in the OPL study were an orifice diameter ratio (Oo/Dl) range of

0.71 to 1.30 (no discrete values reported), and at least two included impingement

angles: 60 and 90 degrees. A curve fit of the ltmtted data generated In their

study produced a design correlation that has since become known as the

ElverumMorey criteria for triplet injectors:

75

= 0.66 (2)

where the subscript o implies an outer orifice and subscript I pertains to the

inner orifice. The A° term is the area of one outer orifice. The term on the

left is referred to as the Elverum-Morey parameter. Based upon the very limited

data of the JPL study, the mixing efficiency appeared to be greatest when this

parameter was equal to 0.66. These criteria were reportedly valid only for

60-degree included angles; the optimum value for the 90-degree case was found to

be 0.42.

Prior to the Elverum-Morey study, the only accepted injector design correlation

was that proposed by Rupe for unlike impinging doublet elements (Ref. 3-3):

Pl Vl 2d1

P2V22d2 =

1.O

RI/RD85-312

III-3

(3)
The Rupe criteria were hypothesized on the premise that the momenta and diameters

of the two streams must be equal to achieve optimum impingement and mixing characterlstics, a premise that was substantiated by extensive experimental results.

One objective of the Elverum-Morey study was to determine whether or not a correlation of the form proposed by Rupe was applicable to triplet elements. Substitution and rearrangement of Eq. 2 and 3 reveal the similarity between the Rupe

[Eq. (4)] and Elverum-Morey [Eq. (5)] forms:

V

%Vo2 (Ao O"5

=k

25

=k'

(4)

(s)

On fitting selected data to the Rupe form, Elverum and Horey reported an optimum

k value of 0.625 for two trlplet elements having dlameter ratios (Do/DI) of

1.00 and 1.29. They further stated that both correlations [Eq. (4) and (5)] gave

the same result for the case of Do/D 1 equal to 1.26, but diverged for diameter ratios significantly different from 1.26. Algebraic manipulation of Eq. (4)

and Eq. (5) reveals that the two forms actually converge at a Do/D i ratio of

1.41. Neither Eq. (4) nor Eq. (5) appears to have a theoretical basis for predlctlng optimum mixing characterlstlcs for triplet elements, and apparently the

empirical correlation shown In Eq. (2) was the best fit for the very limlted data

of the Elverum-Norey investigation.

Experimental Objectives

The objectives of the liquid/liquid trlplet-mlxlng study were to determine the

effect on mixing efficiency of several variations In element geometry and Injection conditions, In order to evaluate the validity of the Elverum-Morey crlteria,

and posslbly to develop an improved correlation. In addition to the Elverum-Rorey

parameter, other comparative stream dynamic parameters, such as momentum ratio

and velocity head ratio, were evaluated In thelr relation to mixing efficlency.

RI/RD85-.312

III-4

V
Table 3-I summarizes the parameter variations assessed in the test program.

Water and l, l, l-trIchloroethane (TRIC) were used as slmulants for RP-I and LOX,

respectively, giving two possible values for Stream density ratio: 0.76 and 1.32.

Ten triplet elements were employed ....a baseline element sized for an ElverumMorey parameter value of 0.66 at a nominal oxldlzer/fuel mass mixture ratio of

2.8, and nine elements incorporating geometric variations around the baseline.

At1 but two of the elements were fabricated by electrlcal-dlscharge-machlnlng

(EDM) the orifices. Orifice lengths were generally assigned a value of Five dlam

eters, and no entrance contouring was provided. The exceptions in terms of lab.

rlcation method were a drilled-orlfice element (orifice L/Ds = 5) whose orifice

entrances were rounded to a O.030-1nch radius, and a precision machined element

(orifice L/Ds _ 24) similar to those employed in JPL mixing studies. Both contoured elements were otherwise identical to the baseline EDMed element. Layout

drawings of the EDM elements (elements lO through 17) were presented in Fig. 2-I.

The drliled orifice element is the number 8 baseline element presented in

Table 2-I, and the details of the element designed to match the 3PL elements are

presented in Fig. 2-3 and 2-4.

Experimental Procedure

The laboratory apparatus employed in the llquld/llquld testing is shown schematlcally in Fig. 3-I. The injector elements were installed in a manifolding

fixture positioned over a rectangular collection grid comprising 260 (13 rows by

20 columns) 1/8-1nch square tubes (Fig. 3-2a), each draining into a corresponding

60-millillter graduated test tube (Fig. 3-2b). Between the collection grld and

the injector element was an alr-actlvated shutter for diverting flow to and from

the grid. Propellant slmulants were supplied to the injector by gaseous nitrogen

pressurization of the two 30-gallon liquid tanks. The procedure for performing

an individual run was as follows: target injection flow rates were achieved by

regulating the tank pressures with solenoid valves open and the shutter closed;

the shutter was opened for a timed .liquid collection interval (approximately 20

to 50 seconds); the shutter was closed, followed by the solenoid valves; and

flnally, the volumes of propellant slmulants in each test tube were recorded.

Because of the immiscibility and d_fferlng densities of the TRIC and water, two

separate phases were clearly distinguishable in the test tubes, and volume

RI/RD85 312

tit-5
TABLE 3-I. OBJECTIVES OF THE LIQUID/LIQUID-TRIPLET MIXING STUDY

PROPELLANT SIMULANTS

FUEL - WATER (DENSITY = 62.4 LB/FT3)

OXIDIZER - 1,I,1-TRICHLOROETHANE (DENSITY = 82.6 LB/FT3)

BASELINE ELEMENT

Do = 0.066 INCH

Di = 0.050 INCH

INCLUDED ANGLE = 60 DEGREES

(SIZED FOR AN ELVERUM-MOREY PARAMETER VALUE OF 0.66 AT

AN OXIDIZER/FUEL MASS MIXTURE RATIO OF 2.8)

EFFECT OF THE FOLLOWING VARIABLES ON MIXING EFFICIENCY TO BE ASSESSED

PARAMETER VARIATIONS

MIXTURE RATIO

DIAMETER RATIO (Do/DI)

IMPINGEMENT ANGLE

ORIFICE DIAMETER

MASS FLOW RATE THROTTLING

FABRICATION METHOD

COLLECTION DISTANCE

INJECTANT CONFIGURATION

2.8 _+25%

1.32 + 20%

60 DEGREES + 30 DEGREES

NOMINAL, v__X NOMINAL

NOMINAL, i/2 X NOMINAL

EDM, SHARP-EDGED; VS DRILLED,

CONTOURED ORIFICES

l, 2, AND 3 INCHES

O-F-O VS F-O-F

VALIDITY OF THE ELVERUM-MOREY CRITERIA TO BE ASSESSED

0.66

F

RI/RDB5-312

III-6
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ORIGINAL PAGE IS

OF POOR QUALITY

a. Square Tube Assembly

Figure 3-2.

b. Test Sample

Liquld/Liquld Mixing Test Apparatus

RI/RD85-312

III-8
measurements of each component were recorded to an accuracy of ±0.2 milliliter.

Distinction between the two liquids was facilitated by dissolving an inert red

dye into the TRIC each time the tank was filled. In many tests, it was necessary

to measure and record TRIC and water quantities from nearly all the graduated

test tubes. Thus, as many as 520 measurements of propellant distribution were

obtained in each test. All these data were input to a computer code that converted them to mixing efficiency data, plots of mass flux distributions, and

other data. Details of this code and plots of mass flux distributions from earlier llquld/llquld triplet tests are presented in Ref. l-l. The primary output

for this detailed study of the effects of injector design parameters on mixing

efficiency, is the mixing efficiency.

Test Plan and Prelimlnar_ Data Anal_Is

Prior to testing, all the elements were flow calibrated over a narrow pressure

drop range corresponding to the range projected for each element in subsequent

test runs. Based on the mass-flow-rate versus pressure-drop (discharge coefficient) data obtained, run conditions were formulated for a series of tests

designed to meet the parameter variation specifications outlined in Table 3-I.

This test plan is presented in Table 3-2. When the plan was formulated, it was

assumed that measurement of orifice pressure drops during test runs was sufficient for accurate computation of InJectant flow rates, given the discharge coefficient data for each element. Therefore, volumetric flowmeters were not included

in the test apparatus. During data analysis following the completion of the

testing, however, it became apparent that in some cases a single target pressure

drop across an orifice had produced widely disparate mass flow rates over the

course of several test runs. The evidence suggested that the inner orifice pressure drop of 30 psld specified for the majority of the test runs often resulted

in fluid cavitation, a phenomenon that went undetected during the narrow range

flow callbratlons.

To clarify the test results, further calibration testing was performed, wlth both

water and TRIC being separately flowed through first the inner and then the outer

orifice sides of each element. Calibration data for the element lO baseline are

shown in Fig. 3-3 and 3-4. Figure 3-3 shows data obtained by calibrating the

RI/RD85-312

III-g
TABLE3-2. LIQUID/LIQUID-TRIPLETTESTMATRIX

ELEMENT

TRIPLETlO

Do = 0.066 INCH

Di = 0.050 INCH

e = 60 DEGREES

(BASELINE

ELEMENT)

TRIPLET 11

Do = 0.055 INCH

Di = 0.050 INCH

e = 60 DEGREES

TRIPLET 12

Do = 0.079 INCH

D i = 0.050 INCH

e = 60 DEGREES

TRIPLET 13

DO = 0.046 INCH

D i = 0.050 INCH

e = 60 DEGREES

TEST

NUMBER

8

9

lO

ll

12

13

14

15

16

17

IB

19

20

21

22

23

24

25

26

27

28

29

COLLECTION DELTA DELTA

DISTANCE P OUT P IN

(INCHES) (PSl) (PSI)

2 5.7

2 lO.1

2 15.8

2 22.6

2 20.2

2 24.2

2 45.3

2

2

2

2

20.6

13.3

8.6

4.5

lO.l

lO.l

ll .3

20.0

30.0

2.8

5.1

7.9

12.8

21.3

12.8

8.5

5.9

II .7

30.0

29.3

17.6

II .7

30.0

30.0

30.0

30.0

60.0

60.0

60.0

30.0

30.0

30.0

30.0

30.0

30.0

30.0

30.0

28.8

30.0

30.0

30.0

30.0

30.0

3O.0

30.0

30.0

30.0

34.8

30.0

30.0

30.0

IN3ECTION

PATTERN

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

F-O-F

F-O-F

F-O-F

F-O-F

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

F-O-F

F-O-F

F-O-F

O-F-O

O-F-O

O-F-O

F-O-F

F-O-F

F-O-F

TARGET

OXIDIZER/FUEL

MASS MIXTURE

RATIO

2.10

2.80

3.50

4.20

2.80

3.50

4.20

0.33

0.41

0.51

0.70

2.80

2.80

2.10

2.80

3.50

2.10

2.80

3.50

4.50

0.23

0.30

0.36

l .00

l .40

2.10

0.59

0.77

0.94

V

RI/RD85-312

III-lO
IABLE3-2. (Concluded)

ELEMENT

TRIPI_EI 14

DO = 0.093 INCH

Di _ 0.071 INCH

e = 60 DEGREES

TRIPLET 15

Do = 0.066 INCH

D i = 0.050 INCH

e = 30 DEGREES

TRIPLET 16

DO = 0.066 INCH

D i = 0.050 INCH

e = 45 DEGREES

TRIPLET 17

Do = 0.066 INCH

D% = 0.050 INCH

e = 90 DEGREES

TRIPLE1 8

Do = 0.065 INCH

Di = 0.050 INCH

e = 60 DEGREES

TEST

NUMBER

30 2

31 2

32 2

33 2

34 2

35 2

36 1

37 3

38 2

39 2

40 2

41 2

42 2

43 2

44 2

45 2

46 2

47 2

48 2

49 2

COI.LECIION

DISTANCE

(INCHES)

DELTA

P OUT

(PSl)

9.0

16.0

25.0

32.6

21 .l

13.?

16.0

16.0

6.5

II .5

13.0

6.0

I0.8

16.8

7.I

12.6

19.7

6.9

12.2

19.0

DELTA

P IN

(PSl)

30.0

30.0

30.0

30.0

30.O

30.0

30.0

30.O

30.0

30.0

30.0

30.0

30.0

30.0

30.0

30.0

30.0

30.0

30.O

3O.0

INJECTION

PATTERN

O F-O

O-F-O

O F-O

F-OF

F-O-F

F-O-F

O F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-f -0

O-F-O

TARGET

OXIDIZER/FUEL

MASS MIXTURE

RATIO

2.10

2.80

3.50

0.33

0.41

0.51

2.80

2.80

2.10

2.80

3.50 '

2.10

2.80

3.50

2.10

2.80

3.50

2.10

2.80

3.50

DRILLED, CONTOURED ORIFICES IN TRIPLET 8 VERSUS EDM, SHARP-EDGED ORIFICES

IN ALL OTHERS

RI/RD85-312

III-II
50

40

0,.

3o

nLU

tT

W

iii

u.

_" 20

0

10

I

I

9

!

!

I

I

! 0

I

!

I /

I

,6 ,

I /

I

I 0 0

cu = o.s8,,,.,_r C/ I

I I

I

I

/

I

/ /

/

_'_ Cd = 0.8 3

/

/

d

o/'

/

/

I t i I i J

0,01 0.02 0.03 0,04 0.05

H20 MASS FLOW RATE (LB/SEC)

Figure 3-3. Water Flow Calibration Data for Element lO Inner Orifice

RI/RD85-312

III-12

V
J_d a_J_JO Ja_no OL _uamaL3 Jo$ e_eO uo_Jq_Le3 mOL_ 3I_1 "_-£ aJnSL_

co_s/8-_3±vu MoT_ssw o_u._

o_o s_'o 9_'o t,Lo _Lo ovo 80o 9oo

I I I I I I I o

f

f

f

f

&,'

i//

ZL'O = P:3 k f

\

f

.,'t_ /

f /

/8"

t /

1 /

t f

S"f

10,,_j _ *

I lj I

f

f

j --

O_

0

O_ -n

m

*0

;I1

m

c

o_:

0

0

Ot_

OS

r_

)

,_il,

)

......... J
inner orifice with water over a pressure drop range from lO to 50 psid. Data

points at pressure drops less than 25 psld follow the Cd = 0.83 llne, while those

above 35 psld follow the Cd = 0.6B llne. This shift in discharge coefficients

between 25 and 35 psld apparently signifies a transition from a noncavltatlng to

a cavltatlng flow regime. However, as evidenced by the data points near 30 psld,

the transition is characterized by a region wherein the exact pressure drop at

which the onset of cavitation occurs is uncertain. Unfortunately, the majority

of mixing test runs were conducted with inner orifice water pressure drops near

30 psid, resulting in flow-rate uncertainty. The calibration data for TRIC flow

through the element lO outer orifices, presented in Flg. 3-4, do not exhibit the

same pronounced Cd shift seen in Fig. 3-3. In this case, a smooth transition

occurs from Cd = 0.72 at 5 psld pressure drop to Cd = 0.66 at 50 psld, and

whether or not the decrease in Cd is caused by fluid cavitation Is uncertain. It

is alsoqulte likely that other effects, in addition to or instead of fluid cavitation, are responsible for the variations in discharge coefficient. Increased

turbulence, changes in the vena contracta et al, can affect flow rate.

..... =-

The calibration plots presented for element lO are typical of those generated by

the EDMed (sharp-edged orifice) elements. Element 8, however, wlth its radially

contoured orifice entrances, produced data for both inner and outer orifice sides

much llke those shown in Fig. 3-4, exhibiting only a gradual decrease in discharge coefficient as pressure drop was increased from lO to 50 psld (from Cd =

0.80 to Cd = 0.74 for the inner orlflce flowing water, and from Cd = 0.82 to Cd =

0.79 for the outer orifices flowing TRIC). Thus, the cavitation phenomenon resulting in flow-rate uncertainty appeared to be characteristic of only the inner orifices of the EDMed elements.

The issues remaining to be resolved, then, were two: an accurate confirmation of

mass flow rates in the tests already performed, and a determination of the effect

of fluid cavitation on mixing efficiency results. The first issue was resolved

by simply calculating the propellant slmulant flow rates for all runs directly

from the liquid masses collected in the grid and including a small correction for

the mass fluxes falling outside the grid as determined by an extrapolation technique. The accuracy of this method was verified by performing tests in which

injection flow rates were measured with turbine flowmeters, and compared with the

RI/RD85-312

Ill-14
flow rates determined subsequently by summation of the collected masses. The run

conditions for these 12 additional tests (tests 50 through 61) are presented in

Table 3-3, and In all cases, the calculated flow rates were within 5_ of the

metered rates. Thus, the issue of the validity of the flow-rate data was resolved.

TABLE 3-3. ADDITIONAL TESTS CONDUCTED TO ASSESS CAVITATION EFFECTS

ELEMENT

TRIPLET I0

DO = 0.066 INCH

DI = 0.050 INCH

e = 60 DEGREES

(BASELINE

ELEMENT)

TEST

NUMBER

50

51

52

53

54

55

56

57

58

59

60

61

COLLECTION

DISTANCE

(INCHES)

2

2

2

2

2

DELTA

P OUT

(PSI)

7.0

ll .5

18.8

30.0

50.0

7.2

II .8

20.0

31.8

50.0

29.7

31.7

DELTA

P IN

(PSl)

23.8

20.I

20.1

20.0

23.0

32.2

30.0

32.5

32.2

32.0

20.0

32.1

INJECTION

PATTERN

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

TARGET

OXIDIZER/FUEL

MASS MIXTURE

RATIO

l .94

2.71

3.46

4.29

4.90

1.99

2.63

3.33

4.08

4.87

4.26

4.08

The effect of fluid cavitation on mixing efficiency results was also addressed in

tests 50 through 61. In runs 50 through 54, inner orifice pressure drops were

held under 25 psld to ensure noncavltatlng flow. Runs 55 through 59 were then

conducted such that the InJectant flow rates corresponded as closely as possible

with the respective conditions of the previous five runs, except that In the latter five runs, the inner orifice pressure drops were increased to ensure cavltatIng flow. (The fact that identical mass flow rates can be achieved at two different pressure drops is clearly evident In Fig. 3-3, and producing the two conditions experlmentally was relatively easy where a turbine flowmeter was used.)

Summary of Test Results

Overall Comparisons. A summary of results for the 61 llquld/llquld m_xlng tests

Is presented In Table 3-4, including mixture ratios, propellant slmulant collectlon efflclencles, mixing efflclencles, and the comparative stream dynamic

RI/RD85-312

iii-15
1---i ::0 j CO

_.j

(._

...J

r_

alum

:1[o

-'10

r_

_mm

zrD-4_

--'4 m!m

-'4Qm

--t m

[,me

¢--. I

"--t

I.,,,4

r"m

"-I

--4 m

--'1

2Z

--'4

r--m

( ('(
k_1

parameters momentum ratio, velocity head ratio, Rupe parameter [defined in

Eq. (4)], and Elverum-Morey parameter. Collection efficiency is defined as the

ratio of collected liquid mass to injected liquid mass. Given the methods used

to determine injected mass flow rates, collection efficiency was calculated as

the ratio of collected mass to the sum of collected mass plus the overshoot estlmate. Collection effIclencles exceeded 90% in the majority of test runs. From a

qualitative standpoint, the raw volumetric data indicate that collection efflclencles had no appreciable effect on computed mixing efflclencles and performance

parameters (i.e., the minor portion of injected mass escaping collection probably

comprised approximately the same ratio of the two components as that exhibited by

the collected portion of the spray fan). The incidence of gradually decreasing

liquid collection efficiency with increasing stream dynamic head is evident, but

again, this effect was not expected to appreciably alter computed parameters.

Plots of mixing efflc_ency versus the comparative stream parameters of velocity

head ratio, momentum ratio, Rupe parameter, and Elverum-Morey parameter are presented in Fig. 3-5 through 3-8. Note that the convention adopted In the calculatlon of these parameters is the ratio of outer to inner streams--the fuel versus

oxidizer notation arises only in connection with mass mixture ratio (oxidizer/

fuel), and InJectant configuration, O-F-O and F-O-F. Because the reversal of

InJectant configuration from O-F-O to F-O-F consistently resulted in lower mixing

efflciencles over the entire range tested for each performance parameter, a distinction between the two configurations has been incorporated into Fig. 3-5

through 3 8. All four plots show a similar trend--mlxlng efficiency increases as

the comparative stream dynamic parameter in question increases, with no optimum

value evident. Both of these results run contrary to the findings of Elverum and

Morey, who reported no density effect over a wider range of variation, and an

optimum value of 0.66 for their correlating parameter, based on their very

limited data. While none ofthe plots exhibits a clear optimization of mixing

efficiency over the range tested, any of the four could be employed as a basis

for evaluating the effect on mixing efficiency of Variations in the individual

parameters specified in the test plan. Because the Elverum-Morey parameter is

the common design criterion, it was chosen as the basis for comparison of these

test results.

RI/RD85-312

III-l?
;0

P-4

M _0

! O0

..j ¢J_

GO I

C,J

r_o

1,0

0.9

0.8 --

-E

>-

Z

I11

_ 0.7 --

1.

W

Z

X

0.6 --

0.5 --

0.4

0.1

0

0

0

0.2

O0

0 000 0

°o

0

0

0

O0

0 0

0

[]

[]

0

OCb

O0

[]

0

0o%

0

[]

[]

0 0 o

0

0

0

[] []

[] rl

0

{Z]

[]

[]

[]

0 O--F--O CONFIGURATION

I'] F-O-F CONFIGURATION

I I I I I I I I

0.3 0.4 0.5 0.6 0.7 0,8 0.9 1.0

VELOCITY HEAD RATIO (O'UTER/INNERI

2.0 3.0

Figure 3-5. Overall Mixing Efficiency vs Velocity Head Ratio for All Test Runs

( ( (
ifi'

 ̧1"!"!

'i .................

II '_

i,,-i

. oo

U3,

,=.=1

0.9

0.8

Z

(.J

-- 0.7

w..

_J

Z

X

0.6

0.5 --

0.4

0.1

0

J I I__

02 0.3 0.4

O

O

O

O

O

0

O0 0 0

o oo

0 o

0

0

[]

0..5 0.6 0.7 0.8 0.9 1.0

MOMENTUM RATIO (OUTER/INNERI

0

O0 0

ccooo o

0 0 0

0 0

C_ 0 0 0

0 0

0

[]

no

[]

[]

[]

[]

[]

[]

[]

[]

o

O[] O-F-OCONFIGURATION F -O-F CONFIGURATION ]

2.0 30 4.0 50 6.0

F_gure 3-6. Overall M_x_ng Efficiency vs Momentum Ratlo for All Test Runs
I GO

0 l

C,3

r_J

1.0

0.9

0.8

--E

ILl

v

>-

0

Z

0

u. 0.7 --

u.

Z

X

0.6 --

0.5 --

0.4

0.1

0

0

0

0

000

CIZ)O 0 0

0 0 O0 0

0 00000

o o %00 o

0 []

[]

[]

d?_° °°o o

[]

[]

0 0 [] []

[]

[] []

[]

FI

[]

O O--F-O CONFIGURATION

[] F-O_F CONFIGURATION

I I I I I 1 I I I 1

0.2 0.3 0.4 0.5 0.6 07 0.8 0.9 1.0 2.0

RUPE PARAMETER

3.0

Figure 3-7. Overall Mixing Efficiency vs Rupe Parameter for all Test Runs

( ( (
1%3 ('_

C_

r_

1.0

0.9 -

0.8 --

Z

E 0.7 -

LI

Z

X

0.6 --

0.5 --

1;1.4

0.1

0

0

0

0

0 o

0

CCD 0 0

°o%

0

0

0

0 000CC_O

O0 O0

0 0 0

0

0

0 0

0

Fn

CO o

0

0

0

OF3 []

[] []

[]

F3

[]

[]

I

00-F--O CONFIGURATION I

I

[] F, O--F CONFIGURATION

I I I I I I I I 1 I

0.2 0.3 0,4 0.5 0.6 0,7 0.8 0.9 1.0 2.0

ELVERUM-MOREY PARAMETER

3.0

Figure 3-8. Overall Mixing Efficlency vs Elverum-Morey Parameter for all Test Runs
Cavitation Effects. Before further analysis of the data, it was necessary to

establish the effect on mixing efficiency of cavltatlng versus noncavitatlng flow

and collection distance, and to assess the repeatability of specific data points.

Foremost among these issues was the question of whether or not fluid cavitation

appreciably affected element mixing characterlstlcs. The cavitation issue Is

important for two reasons. First, because the incidence of cavitation in individual runs was fairly random throughout the test matrix, if mixing characteristics were shown to be appreciably altered by its occurrence, the validity of comparing results between cavitating and noncavltatlng tests would be suspect (i.e.,

the internal consistency of the data would be threatened). Second, in actual hotfiring of liquid rocket thrust chambers, propellants are injected into a highbackpressure environment, effectively precluding the incidence of cavitation in

the orifices. Therefore, if mixing characteristics were shown to be altered by

cavitation, the application of correlations obtained In cavitattng test runs to

conditions known to be noncavitatlng would be suspect (i.e., the external applicability of the data would be threatened).

Cavitation In fluids passing through an orifice wlth a sharp-edged entrance Is a

complex phenomenon. It Is most simply thought to occur when the static pressure

In the liquid falls below the vapor pressure of the liquid. Bubbles form that

can have large effects on flow characteristics such as discharge coefficients,

stream turbulence, and stream coherence, and, in some cases, can lead to detached

(from the orifice wall) and hydraulically flipping streams (detached and Jumping

from one side of the orifice to another side).

The flow and orifice conditions that affect or cause cavitation are not simple.

Vena-contracta effects at the entrance, upstream flow conditions, dissolved gases

In the liquid, and many other factors influence the initiation, degree, and

effects of cavitation. Furthermore, hysteresis effects cause the onset or cessation of cavitation to be dependent upon whether or not the flow rate Is increasing or decreasing. As shown in Ref. 3-4, different reglmes of cavitation can

occur which have differing, and sometimes little, effect on the measured mixing

efficiency of like doublets. If the stream is completely attached to the orifice

RI/RD85-312

III-22
wall at the orifice exit, upstream cavitation may be expected to have little

effect on mixing. Also, according to Ref. 4, if the stream is uniformly detached

(and not flipping), unllke, impinging doublet mixing will be 11ttle affected. To

resolve the issue of the effects of cavitation in the present test program, the

simplest approach was to perform tests that differed only by the presence of

cavitation, and to observe the effect on mixing efficiency.

Figure 3-9 shows the results of the test runs designed to address the cavitation

issue, lhe cavltatlng versus noncavltatlng test pairs at each of the five

Elverum-Morey parameter values (i.e., mixture ratio conditions) are seen to

closely coincide with respect to mixing efficiency -- the maximum spread in Em is

3% at an ElverumMorey parameter value of approximately 1.50. Presented in

Fig. 3-I0 is a plot of another four data pairs, showing the results of repeatability assessment tests. In this case, the maximum spread in mixing efficiency

observed between two members of a pair is 2%. In light of this 2% repeatability

error, the cavltatlng/noncavltatlng data pairs of Fig. 3-9 can be taken as essentlally coincident points, thus ensuring the validity of internal comparison and

external application of the data generated in the test matrix.

Collection Distance Effects. Figure 3-11 shows the limited number of data points

used to determine the effect of varying collection distance on mixing efficiency.

Choice of collection distance in mixing studies is essentially arbitrary. The

baseline value of 2 inches chosen for this test program is a fairly common standard for studies of this nature, and, in this case, a convenient distance for capturlng the triplet spray fans in a large portion of the I-5/8 by 2-I/2 inches

grid with a m%nlmum of overshoot. The four data points plotted in Fig. 3-12 show

that mixing efficiency increased from l to 2 inches, and again from 2 to 3

inches, by approximately 5% in both cases. Relative spatial distribution of propellant slmulants remained essentially the same in both test pairs. While the

magnitude of the mixing efficiency increase is greater than the limits of repeatability error, It is still not overly significant, and the trend Is in the

expected direction (i.e., mixing improves at greater collection distance). It

would be hard to rationalize a spray becoming less mixed with axial distance.

RI/RD85-312

III-23
I

1.0

0.9 -

0.8 --

i11

Z

- o.7 -

u.

u

u,I

Z

X

0.6 --

0.5 --

0.4

0,1

I I

0,2 - 0.3

[]

QO

0 0

[]

0

[]

O NONCAVITATING FLOW THROUGH INNER ORIFICE

[] CAVITATING FLOW THROUGH INNER ORIFICE

ELEMENT 10, TEST RUN NUMBERS 50.-54 AND 55-59

I I I I I i I

0.4 0.5 0:6 0.7 0.8 0.9 1.0

ELVERUM-MOREY PARAMETER

CO

I

2.0 3.0

Figure 3-9. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Fluid Ca4" _tlon (
co

t_j

cjm

P_

1.0

0.9

0.8

-E

uJ

u

z

w

-° 0.7

11

ul

o

z

x

0,6 B

0.5 _

0.4

0,1

I

0,2

0

0

0

0

0

0

TEST RUN NUMBERS 2 AND 56, 49 AND 70, 58 AND 61, AND 53 AND 60 ]

I

0.3

1 I I I I I I

0.4 0.5 0.6 0.7 0.8 0.9 1,0

ELVERUM--MOREY PARAMETER

2.0 3.0

Figure 3-I0. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Repeatability Tests
P-4

P--4

7

t_J

O_

_O

_O

GO

I

t_

--E

uJ

v

>-

_3

Z

Ill

U.

I*L

M.I

Z

X

r_

1.0'

0.9

0.8

0.7

0.6

0.5 --

0.4

0.1

I I

0.2 0.3

[]

O

O ELEMENT 10. COLLECTION DISTANCE = 1 INCH (TEST RUN 121

ELEMENT 10, COLLECTION DISTANCE = 2 INCHES (TEST RUN 2)

_;_ ELEMENT 14, COLLECTION DISTANCE = 2 INCHES (TEST RUN 30)

• ELEMENT 14, COLLECTION DISTANCE = 3 INCHES (TEST RUN 37)

I I I I I I I I

0.4 0,5 0:6 0.7 0.8 0,9 1.0 2.0

ELVERUM--MOREY PARAMETER

3.0

Figure 3-11. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Collection Distance

C (i
1.0

I

.-J

0.9

0.8 m

W

>-

o

z

0,7 -

u_

W

z

x

0.6--

0.4

0

I I I I I

0.5 1,0 1.5 2.0 2,5

O AEROJET OFO TRIPLET

(MR = 2.631

/_ AEROJET TRANSVERSE LIKE-ON-LIKE DOUBLET

(MR = 2.80)

O AEROJET PREATOMIZED OFO TRIPLET

(MR = 2.85)

I_1 AEROJET EDM'ed LIKE-ON-LIKE DOUBLET

IMR = 2,91 )

O ROCKETDYNE OFO TRIPLET 10

(MR = 2.58}

• ROCKETDYNE OFO TRIPLET 14

(MR = 2.25l

(AEROJET MR's FREON/H20}

(ROCKETDYNE MR's TRIC/H^O)

AEROJET DATA FROM REF.43-5

I I '

3,0

COLLECTION DISTANCE (IN,)

Figure 3-12. Overall Mixing Efficiency vs Collection Distance for Combined

Rocketdyne and Aerojet Data
Additional collection distance data are provlded In the report of an AeroJet

study (Ref. 3-5), In which four different injector, elements were tested at distances of 1/2, l, 1-1/2, and 2 inches each. The AeroJet data are presented _n

Fig. 3-12 alongside the data from this study. No conslstent trend Is evident In

these plots, as mixing efficiency varied anywhere from 2.4% to 6.7% for the elements as collection distance was varied. Furthermore, In some cases the mixlng

efficiency decreased with collection distance. The test performed as a part of

this program and the AeroJet tests indicate that the effect of varying collectlon

distance on mixing efficiency ls not substantial, under the conditions of these

tests.

Effect of Variations In Individual Parameters. Figures 3-13 through 3-15 show

the effect on mixing efficiency of relatlve stream diameter. Five elements havIng different outer orlflce to Inner oriflce diameter ratios were tested In the

O-F-O configuration to generate the data of Fig. 3-13. These data show no significant effect on mlxlng efficiency of variations In Oo/D I, except for the

element In which Do/D I was less than unity (element 13, Do/D I = 0.92).

Mixing efflclencles of this element fell significantly lower than those produced

by the other elements. Figure 3-14 shows the effect of dlameter ratio variation

for four elements flowed in the F-O-F configuration, and In this case, the data

exhibit no slmple trend that can be translated lnto a generalized diameter ratio

effect. The elements wlth both the highest and lowest diameter ratios experienced

poorer mixing. As mentioned prevlously, configuration reversal from O-F-O to

F-O-F consistently resulted in lower mixing efficiency. The data presented In

Flg. 3-15 were generated by using elements whose outer orlflce to lnner orifice

diameter ratios were equal, but whose total flow area differed by a factor of 2.

In thls case, the data lndlcate that orifice slze has no effect on mixing efficIency. This should not be Interpreted to Imply that larger elements w111 provide

mixing that Is as good as smaller elements. Mlxlng efficiency Is a measure of

mixing uniformity per unlt mass of propellant Injected. Larger elements wlll

have larger regions of "off-deslgn" mixture ratlo than smaller elements, even

though both may have the same mixing efflclency. These larger regions will

require longer chamber lengths to be mixed with surrounding gases. Figures 3-13

to 3-15 again illustrate that no definite Elverum'Morey optlmum was evident over

the range tested.

RI/RD85-312

111-28
1.0

i--,i

,';" co

r_ u"-

r,,,a

._..=

0.8

¥

u.I

v

>-

o

z

UJ

o

0.7

z

x

0.6

0.5

0.4

0.9 m

m

I

0.1

Figure

A

<>

I

0.2

3-13.

A

Z_ []

Oo oo

0%

0 ED AO,_,_ 0 0

oO

o 0

0

O ELEMENT10 Do= 0,0661N., Di = 0.0501N. (TEST RUN NUMBERS 1-4ANDS0--61)

[] ELEMENT 11 Do =0.0551N,,Di = 0.0501N. (TEST RUN NUMBERS 14-16)

/_ ELEMENT 12 Do = 0.079 IN., Di = 0.050 IN. (TEST RUN NUMBERS 17--20)

O ELEMENT 13 Do = 0.046 IN., Di = 0.050 IN. (TEST RUN NUMBERS 27-29)

(_) ELEMENT 14 Do = 0.093 IN., Di = 0.071 IN. (TEST RUN NUMBERS 30-32)

I

O3

I I I I I I I

0.4 0 5 0.6 0.7 0.8 09 1.0

E LVERUM-MOREY PARAMETER

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Diameter Ratio (0-F-0 Configuration)

!

2.0 3,0
1.0

H

I--I

'7

C,O

O

_O

:;O

CO

U'I

I

,.-J

r_

-E

UJ

C.)

Z

LU

0

u

i11

o

z

x

0,9 i

0.8 --

0.7 --

0.6 --

0,5 --

0.4

0.1

Figure

I

0.2

3-14.

0

0

0

A

0

0 o

CO

A

A

<>

0 ELEMENT 10 Oo = 0.066 IN., D i : 0.050 IN. (TEST RUN NUMBERS 8-11)

,A, ELEMENT 12 Do = 0.079 IN., D i = 0.050 IN. (TEST RUN NUMBERS 21-23)

O ELEMENT 13 D o = 0.046 IN., D i = 0.050 IN. (TEST RUN NUMBERS 27-29)

O ELEMENT 14 Do=0.093 IN., Di=0.071 IN. (TEST RUN NUMBERS 39--41)

i

I I 1 I I I I I

0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

ELVERUM-MOREY PARAMETER

I

2.0

Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Diameter Ratio (F-O-F Configuration)

3.0

( ( C
1.0

i,-,.I

y

t.4

.,,..a

;=o

co

I

,r.,,O

i.i

t"¢,

0.9

0.8

-E

I.I.I

>

z

LU

_ 0.7

ill

z

x

0.6

0.5

0.4

0

0 0

0 []

0

0

00C3 0 0

0

0

O0 QO

O ELEMENT 10 Do=0.0661N., D i = 0.0501N. (TEST RUN NUMBERS 1--4,8--11,AND50-61)

[] ELEMENT 14 D o = 0.093 IN., Di = 0.071 IN. (TEST RUN NUMBERS 30-32 AND 33-359

(FLOW AREA OF ELEMENT 14 IS TWICE THAT OF ELEMENT 10)

(OUTLINED SYMBOLS = O-F-O CONFIGURATION, SOLID SYMBOLS = F-O-F CONFIGURATION)

i

I I

0.2 0.3

I I I I I I I

0.4 0.5 0.6 0.7 0.8 0.9 1 0

ELVERUM-MOREY PARAMETER

20 3.0

Figure 3-15. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Orifice Scaling
Figure 3-16 shows data generated In tests where impingement angle was varied,

with orifice diameters held constant. The four values of total Included angle

between outer streams were 30, 45, 60, and 90 degrees. Although one of the data

points falls anomalously outside the scatter band of the other tests, the results

of this test series indicate that impingement angle has no significant effect on

mixing efficiency. By contrast, Elverum and Morey reported a pronounced effect:

for 60-degree Implngement angles, the optimum Elverum-Morey parameter value was

computed to be 0.66, while for gO-degree angles, 0.42 was the reported optlmum.

The effect on mixing efficiency of increased orifice pressure drop, with mixture

ratio held approximately constant, ls shown In Fig. 3-17. In four of these test

runs, mass flow rates were set at "nomlnal '° values for producing the desired mixture ratios. In the other four runs, elevated pressure drops (approximately twice

the nomlnal values) were achieved to produce approximately the same mixture

ratios, with approximately 40% increase in both TRIC and water flow rates. As is

evident from the plot, no significant effect on mlxlng efficiency was observed.

Data presented In Flg. 3-18 Illustrate the effect on mixing efficiency of different orlflce machinlng methods. The two elements employed in this test series were

dimensionally Identical, but In one case the orifices were drilled and provided

with rounded (O.030-1nch radius) entrances, and In the other, the orifices were

EDMed and left with sharp-edged entrances. Although only slx data points were

generated wlth element 8, the trend apparent from these points could posstbly be

construed to exhibit a mixing efficiency maximum near the optimum proposed by

Elverum and Horey (approximately 0.65). The element lO data followed the same

general trend observed In the bulk of the test results, wlth mixing efficiency

increasing wlth increasing Elverum-Morey parameter.

Investlqatlon of Orifice Contouring, Fabrication Method, and Flow Control Featur_

Effects. The question posed by the results of this fabrication method test series

was whether the apparent optimum In the element 8 data was In fact a reproducible

trend, or simply caused by data scatter. Furthermore, it was theorized that perhaps the differences between these tests and the limited JPL tests could be the

RI/RD85-312

III-32
1.0

I GO

1

-.J

I%)

0.9 m

0.8 B

uJ

v

>-

_J

z

uJ

rj 0.7

z

x

0.6 _

o.g _

0.4

0.1

0

0 o

ZX

0 0

O[3 Z_

o 0

o

0

0C30 0

0

0

[]ELEMENT 15_ = 30 DEGREES (TEST RUN NUMBERS 38-401

ELEMENT 16_ = 45 DEGREES (TEST RUN NUMBERS 41-43)

O ELEMENT 10(_ = 60 DEGREES (TEsT RUN NUMBERs 1-4 ANDS0--61}

OELEMENTI7_ = 90 DEGREES (TEST RUN NUMBERS 44--46)

((_= TOTAL INCLUDED ANGLE BETWEEN OUTER STREAMS)

(D O - 0.066 IN., Di = 0.050 IN. FOR ALL FOUR ELEMENTS)

I I I I I I I ! I I

0.2 0.3 0.4 0.5 0.6 0.7 0.8 0,9 1.0 2.0

ELVERUM-MOREY PARAMETER

3.0

Figure 3-16. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Impingement Angle
1.0

e-4

?

.b,

;o

_o

CO

l

P_

0.9 m

0.8

m

0.6_

0.5

0,4

0.1

t,M

>-

u

z

0.7

u

Z

x

Figure 3-17.

0

0

0

[]

[] 0

0 []

0 ELEMENT 10 NOMINAL FUEL AND OXIDIZER SIDE MASS FLOWS (TEST RUN NUMBERS

[] ELEMENT 10 40% INCREASE IN MASS FLOWS (TEST RUN NUMBERS 5-7 AND 53)

i

1--4)

I I I I I I I I I I

0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.91.0 2.0

ELVERUM-MOREY PARAMETER

0vera]] Mixing Efficiency vs Elverum-Morey Parameter - Effect of

Injection Ve]oclty at Constant Ve]oclty Ratio (Thrott]Ing)

3.0

( ( (
1.0

P-4

'7

;lO

co

u".

,,_J

t_

)-

(.3

z

W

u

C_

Z

X

0.9 m

0.8--

0.7 --

0.6 --

0.5 --

0.4

0.1

[]

O

O O

[]

0

0

0

00 0

o

(p [] o []

0 ELEMENT 10 SHARP-EDGED, EDM ORIFICES (TEST RUN NUMBERS, 1--4 AND 50-61)

[] ELEMENT 8 CONTOURED, DRILLED ORIFICES (TEST RUN NUMBERS 47-49 ANDREF 1-1 TESTS)

(Do = 0,066 IN,, D_ = 0.050 IN. FOR BOTH ELEMENTS)

I I I ! IIIII

0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

ELVERUM-MOREYPARAMETER

I

2,0 3.0

Figure 3-18. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Effect of Fabrication Method
result of the differences in orifice tnlet contouring, fabrication method, and/or

the JPL specifications for turbulence generation and long length to diameter orifices. Were the conclusions about the Elverum-Morey parameter optimum, indicated

by the JPL reports and perhaps by the slx tests with element 8, caused by these

differences, or were they simply the result of a very small collection of

scattered data? This is an important question, as the Elverum-Morey criterlum has

long been applied as an important triplet design consideration.

To resolve this issue, additional tests were proposed for the element 8, and for

an additional element, designated 20, which was fabricated to closely resemble

those employed by 3PL investigators. Triplet element 20 was formed by positioning three approach tube/orlflce insert assemblies on a mounting fixture In a

60-degree impingement configuration with impingement distances identical to those

of the one-piece elements 8 and lO. Element 20 orifice diameters matched those

of element iO, but lengthy orifice L/Ds, entrance contouring, and inner surface

roughening were included, as derived from the JPL standards. A summary of the

potentially important differences among the elements 8, lO, and 20 is presented

in Table 3-5, and detailed designs of these elements are presented in Section II.

A schematic drawing of an approach tube/orlflce insert assembly employed In the

JPL m}xlng studies (Ref. 3-I) Is shown In Fig. 2-2. Orifice inserts of thls type

were preclslon-machlned to include elllptlcally contoured entrances; a short,

roughened section downstream of the entrance for inducing turbulence (achieved by

tapping the inside diameter to provide a short length of threading); and overall

length-to-dlameter ratios exceeding 20.

Table 3-6A shows the additional tests performed with the elements 8 and 20, followed by the results of these tests. Data from the elements 8, lO, and 20 tests

are plotted In Flg. 3-19. Pronounced optimums In the elements 8 and 20 data are

not evident. The results presented in Table 3-6B and Flg. 3-1g appear to underscore the general trend observed for all elements tested In this study --- mixing

efficiency increased steadily as the Elverum-Morey parameter increased from

approximately 0.3 to 0.8, then leveled off at a high value with further increase

In the Elverum-Morey parameter. Thls result Is in marked contrast to the findings

of Elverum and Morey. Included In Fig. 3-19 are flve of the actual data points

RI/RD85-312

III-36
TABLE3-5. SUMMARYOFVARIATIONSIN ORIFICECONIOURING _J

ELEMENT 8:

ELEMENI I0:

ELEMENT 20:

OUTER ORIFICE DIAMETER 0.065 INCH

INNER ORIFICE DIAMETER 0.050 INCH

LENGTH/DIAMETER RATIO 5.0 (ALL ORIFICES)

ORIFICE ENTRANCE RADIUS 0.030 INCH (ALL ORIFICES

OUTER ORIFICE DIAMETER 0.066 INCH

INNER ORIFICE DIAMETER 0.050 INCH

LENGTH/DIAMEIER RATIO 5.0 (ALL ORIFICES)

ORIFICES ENTRANCES SHARP-EDGED

OUTER ORIFICE DIAMETER 0.066 INCH

ENTRANCE RADIUS 0.061 INCH

LENGTH/DIAMETER RATIO 24.0

TURBULENCE INDUCEMENT 2-56 THREADED SECTION

INNER ORIFICE DIAMETER = 0.050 INCH

ENTRANCE RADIUS 0.057 INCH

LENGTH/DIAMETER RATIO 24.0

TURBULENCE INDUCEMENT 0-80 THREADED SECTION

TABLE 3-6A. ADDITIONAL TESTS CONDUCTED TO TEST FABRICATION METHOD

COLLECTION DELLA

ELEMENT

TRIPLET 20

Do = 0.66 INCH

Di = 0.050 INCH

8 _ 60 DEGREES

TRIPLET8

Do = 0.065 INCH

Di = 0.050 INCH

8 = 60 DEGREES

TEST

NUMBER

62

63

64

65

66

67

68

69

70

71

DISTANCE P OUT

(INCHES) (PSI)

2 9.2

2 12.0

2 14.5

2 18.5 I

2 22.9

DELTA

P IN

(PSI)

26.9

26.5

26.-/

26.5

26.5

9.7 26.7

12.5 26.7

15.5 126.8

19.4 126.5

24.4 126.8

INJECIION

PATIERN

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

O-F-O

TARGET

OXIDIZER/

FUEL MASS

MIXTURE RAIIO

2.35

2.70

2.96

3.35

3.73

2.34

2.66

2.96

3.33

3.71

RI/RD85 312

III-37
TABLE3-6B. RESULTSOFADDITIONALFABRICATIONMETHODTESTS

I

O0 !

TEST ELEMENT

NUMBER NUMBER

62 20

63 20

64 20

65 20

66 20

67 8

68 8

69 8

70 8

71 8

OXIDIZER/FUEL

MASS

MIXTURE RATIO

2.35

2.70

2.96

3.35

3.73

2.34

2.66

2.96

3.33

3.71

OXIDIZER

(TRIC)

COLLECTION

EFFICIENCY

0.82

FUEL

(H20)

COLLECTION

EFFICIENCY

0.95

OVERALL

MIXING

EFFICIENCY

(Em)

0.88

OUT/IN

MOMENTUM

RATIO

l .20

OUT/IN

VELOCITY

HEAD

RATIO

0.34

RUPE

PARAMETER

0.45

0.83

0.83

0.83

0.83

.84

.88

.87

.87

.86

0.98

0.93

0.81

0.76

0.92

l .06

0.99

0.96

0.85

0.92

0.95

0.95

0.95

0.91

0.95

0.96

0.94

0.95

1.57

1.89

2.44

3.02

1.22

1.57

1.95

2.46

3.08

0.44

0.54

0.70

0.87

0.36

0.46

0.58

0.72

0.91

0.58

0.71

0.92

l .15

0.47

0.60

0.75

0.94

1.18

ELVERUMMOREY

PARAMETER

0.49

0.65 •

0.78

1.00

1.25

0.49

0.63

0.78

0.99

1.23

( ( (
*--4

I

_D

oo

{.P

t_

1.0

0.9

0.8

-e

LU

>.

z

0,7

LI_

LI.

.i

z

x

0.6

0,5

0.4

0

•

00 0

@ 0

@ _ O_ @@ O00 @ 0

00

0 •

A

0 ELEMENT10 SHARP-EDGED, EDM ORIFICES (TEST RUN NUMBERS 1-4 AND 50--61)

ELEMENT S CONTOURED, DRILLED ORIFICES (TEST RUN NUMBERS 47-49, 67-71, AND EARLIER TESTS [REF 1-1])

• ELEMENT 2OJPL MODEL, PRECISION ORIFICES (TEST RUN NUMBERS 62-661

(Do = 0.066, D i = 0,050 FOR ALL THREE OF THE ABOVE ROCKETDYNE ELEMENTS)

• JPL DATA POINTS FOR Do/D i = 129 AND 1.00 (O-F--O TRIPLETS)

I I I I I 1 I 1 I t

0.2 0.3 0.4 0.5 0.6 0.7 0,8 0.9 1.0 20

ELVERUM - MOREY PARAMETER

3.0

Figure 3-19. Overall Mixing Efficiency vs Elverum-Morey Parameter -

Rocketdyne vs JPL Data
employed In the Elverum-Morey correlation (Ref. 3-6 and 3-7). Four of these

points seem to follow the genera] trend observed In the present study, while the

fifth falls much lower than comparable points produced here. Efforts to uncover

all data generated in the 3PL trlplet studies culminating In the Elverum-Rorey

correlation proved largely fruitless. The only data available In referenced

documents are shown in Table 3-7. Valid comparison of Rocketdyne versus 3PL data

depends on the assumption that mixing results are independent of the sllghtly

different spray sampltng methods employed. While the valtdlty of this assumption

has not been directly addressed here, it was reported In a previous 3PL study

(Ref. 3-4) comparing Rocketdyne and 3PL data, that "none of the results appeared

to be significantly influenced In any way by differences In apparatus and experimental technique."

Data Correlation

Data from the 71 mixing tests were correlated wlth the ald of a regression program that computed the relation between the dependent variable (mixing efficiency) and a set of four independent variables (e.g., stream diameter ratio, density ratio, velocity ratio, and impingement angle). These four variables were

chosen as the most fundamental physical descriptors of a particular set of injection conditions. Parameters such as momentum ratio or mixture ratio could Just

as easily have been employed, but they represent combinations of the fundamental

quantities. The correlating equation computed was:

Em = 0.81 kdi/

.29 (Po _0.56 /Vo ?.25 03

V'l/ \Vll Cslne)-°" (6)

The applicability of this correlation Is subject to the following limits on the

variable ranges:

Outer/Inner stream density ratio

Outer/inner stream diameter ratio

Included impingement angle

Outer/Inner stream velocity ratio

0.76 to 1.32

0.92 to 1.58

30 to 90 degrees

0.37 to 1.69

RI/RD85-312

III-40

V
TABLE3-7. SUMMARYOFAVAILABLEDATAFROM3PLTRIPLETSTUDIES

TEST

NUMBER

REFERENCE Do

NUMBER (]NCH)

3-6 N/A

3-6 N/A

3-6 N/A

3-7 N/A

3- 7 N/A

3-2

3-2

3-2

OXIDIZER/FUEL

D_ PROPELLANTS

(INCH) Do/D i SIMULANTS

N/A l.O0 CCI4/H20

N/A 1.00 CC14/H20

N/A l.O0 CCI4/H20

N/A 1.29 CCI4/H20

N/A 1.29 CC14/H20

0.082 0.070 1.17 H20/KEROSENE

0.082 0.070 1.17 H20/KEROSENE

0.082 0.070 1.17 H20/KEROSENE

O-F-O O/F

DENSITY MASS

RATIO MIXTURE RAIIO

1.59 1.26

1.59 1.59

1.59 2.25

1.59 2.64

1.59 2.97

1.22 1.92

1.22 2.18

1.22 2.42

OUWIN

VELOCITY HEAD

RATIO

0.50

0.79

1.59

0.40

0.50

0.40

0.52

0.63

ELVERUM-MOREY

PARAMEIER

0.60

0.94

l.89

0.54

0.6/

0.52

0.66

0.82

MIXING

EFFICIENCY

(Em)

0.82

0.89

0.70

0.86

0.90

N/A

N/A

N/A

.,-,i
A plot of observed mixing efficiency values for the 71 tests versus the values

predicted using Eq. (6) Is presented in Fig. 3-20. Fifty-six of the 71 predicted

Em values fell within 6% of the corresponding observed values. It ls evident

that the majority of points whose predicted Em values deviated from the observed

values by more than 6% were those resulting from F-O-F configuration tests. The

greatest deviation from this correlation was observed for element 13, which was

the only element wlth a diameter ratio less than 1.

The empirical correlation of Eq. (6) mathematically defines, and agrees with, the

findings that were apparent from inspection of the data. For example, the great

benefit of havlng the denser fluid In the outer stream and the relative unimportance of impingement angle are readily apparent in both Eq. (6) and the graphical

results previously presented.

Conclusions

The inescapable conclusion to be drawn From the results presented here Is that

the Elverum-Morey criteria are an invalid correlation for the design of triplet

elements composed of sharp, edged entry, low L/D orifices. These types of elements

are highly representative of actual rocket engine injector elements. Furthermore,

limited data generated wlth elements composed of contoured entry, high L/D orlfices seem to support this conclusion more generally. The invalidity of the crlterla Is based primarily on the observation here that mixing efficiency does not

reach a maximum value at an Elverum-Morey parameter value of 0.66. Certain

earller Rocketdyne testing of F-O-F injectors wlth Do/D I _l (Ref. l-l)

indicates a peak at an Elverlm-Morey parameter value near 0.66. Those tests are

outside the range of conditions tested In this program. It Is possible that the

optimum Elverlum-Morey parameter Is valid under those test conditions.

V

Another conclusion demonstrated here is the importance of injecting the denser of

two liquids through the outer orifice pair of a triplet element, which is seen to

produce markedly better mixing efficiency than the reversed configuration (i.e.,

lighter fluid injected through the outer orifices). Thls effect may be the result

of the greater momentum of the outer streams wlth denser liquids. Since it Is

the impingement of the outer streams that causes the breakup and mixing of the

RI/RDB5-312

III-42
C

t--i

C_

r_

1.00

0.90

0.80

0.50

O4O

0.40

APPLICABLE RANGES OF PARAMETERS

0.92 < Do/D i <1.58

0.76<Po/p i < 1,32

0,37< Vo/vi< 1.69

30 ° < _ < 900

DO 0

0

ELEMENT OFO FOF

8 ®

10 O •

11 []

I I I I I

0,50 0,60 O. 70 0.80 0,90 1.00

Figure 3-20. Data Correlation for Liquld/Liquid Triplet Mixing Study
inner stream, increasing the momentum (by lncreaslng the density) of the outer

stream would be expected to enhance mixing. Regardless of the reason for this

effect, the results of these tests clearly demonstrate a very strong correlation

between triplet mixing efficiency and the selection of inner or outer orifices

for the more dense propellant. Slnce the most commonly employed propellant comblnatlons have fuels of lower density than the oxidizer, this finding would dictate the use of oxldlzer in the outer streams. In many instances, particu7arly

at preburner mixture ratios, other factors may preclude the use of oxidizer in

the outer streams. Llquld/llquid triplets may not be the optimum element choice

for such applications. The liquid/liquid triplet testing performed with preburner elements (elements 1 and 6) In an early part of this contractual program

(Ref. 1) demonstrated poor mixing efficiency. This may be caused by their F-O-F

configuration.

Another factor that appears to have an effect on mixing efficiency is the outer

to inner orifice diameter ratio. Thls Is apparent In the emplrica7 correlating

equation [Eq. (6)], but is not so readily apparent from the p7ots of the data

(Fig. 3-13 and 3-14). In the O-F-O configuration (Fig. 3-13), there Is no discernible difference In mixing efficiency between all the elements tested, except

for the element 73, which is the on7y element with an outer to inner diameter

ratio less than 1.0 (actually the only one less than 1.7). Figure 3-14 (F-O-F

configuration) seems to indicate that both e7ement 13 (Do/D I = 0.92) and element i2 (Oo/O 1 = 7.58) were relatively poor mixers, while those of intermediate diameter ratios provided better mixing. Thus, the following findings regardlng the effect of diameter ratio are indicated.

V

,

.

,

The empirical correlation indicates that larger values of diameter ratio

generally enhance mixing

Figure 3-13 indicates that mixing efficiency is unaffected by diameter

ratio except at low values of diameter ratio for O-F-O configurations

Figure 3-14 indicates that both large and small values of diameter ratio

inhibit mixing for F-O-F configurations.

RI/RDBS-312

III-44
J

While all three of these findings support the conclusion that lower values of

diameter ratio reduce mixing performance, they are contradictory regarding the

effects of higher diameter ratios. Even this conclusion is somewhat suspect, as

it is entirely dependent upon the results obtained with a single element 13.

However, this conclusion seems reasonable, since smaller outer streams might be

expected to less homogeneously break up and mix with a larger center stream and

even minor mlslmplngements could significantly degrade mixing performance. Addl

tlonal effort is recommended to confirm this conclusion and to investigate large

dlameter-ratlo elements mixing performance.

The other major parameter affecting the mixing performance of llquld/llquld triplets is the velocity ratio. This is demonstrated in Fig. 3-5, where mixing efficlency versus velocity head ratio (PoV_/piV_) is plotted. For

either the O-F-O or F-O-F configuration, the density ratio remains constant.

Thus it is apparent that increasing the velocity ratio (Vo/VI) increases mixing performance for either configuration. The empirical correlation [Eq. (6)]

also clearly demonstrates this effect. Limited testing (presented in Fig. 3-I?)

indicates that it is the velocity ratio rather than the individual velocities

which affect mixing. This is an important finding as it provides some basis to

support extrapolation of these data to higher injection velocities.

An extensive number of tests with four elements, whose only difference was

impingement angle, indicates that impingement angle has no apparent or significant effect on mixing. These results are presented in Fig. 3-16, and the correlating equation demonstrates no significant effect of impingement angle. This,

too, is in disagreement wlth the early SPL findings.

One additional flhdlng of this study is the insignificance of orifice size on

mixing efficiency. While the number of tests (presented in Fig. 3-15) is

relatively small, there appears to be no effect on mixing efficiency when the

orifice area is doubled. This finding appears to contradict common hot-fire test

experience on engines that indicates poorer performance for larger elements.

However, this poorer performance may be the result of atomization rather than

mixing. Also, even though the measured mixing efflclencles of large and small

elements may be the same, the regions rich in one or the other propellants are

RI/RD85-312

III-45
correspondingly larger for the larger elements. For example, both elements may

provide a I0% higher mixture ratio zone In the center of the distribution and a

surrounding lower mixture ratio region such that both elements have the same

cold-flow mixing efficiency. However, the slze of these regions may be expected

to scale wlth the element size. Thus, the larger element would have a correspondingly larger hlgh mixture ratio region, which would not be as readily

dlspersed and reacted as the smaller hlgh mixture ratio region of the smaller

element.

V

Thls program Is the largest, most thorough, and most comprehensive investigation

of the mixing performance of llquld/llquld-trlplet elements that has ever been

performed. The findings and conclusions of this effort can provide specific

guidance to designers of such elements. In accordance wlth these findings, the

following design criteria and advice to designers of llquld/llquld-triplet elements are provided.

l °

.

3.

.

.

,

•

The use of an optimum value of the Elverum-Morey parameter as design

criteria for llquld/llquld triplets is not Justified and should be dlscontinued. Mixing efficiency increases as the E1verum-Morey parameter

Increases.

The more dense propellant should be injected from the outer orifices.

Small, outer to inner, orifice diameter ratios (e.g., l or less) are not

recommended.

Changing the injection velocity of both elements by the same percentage,

has no effect on the mixing performance, over the range of injection

velocities tested.

Increasing the outer to Inner velocity ratlo of the propellants improves

mixing performance.

Impingement angle has no significant influence on mixing and should be

chosen on the basis of other considerations.

As a guide to use in trading off various design considerations, it is

recommended that the designer attempt to maximize the parameter:

RI/RD85-312

III-46

v
All these findings, conclusions, design guidance, and especlally the preceding

mixing efficiency design parameter, should not be extrapolated or applled beyond

the llmlts of the range of the variables tested in this program without very

careful cons%deratlon. The range of the variables tested are:

Outer to inner orifice veloclty ratio

Outer to inner orifice diameter ratio

Outer to inner orifice density ratio

Impingementangle (between outer streams)

Orifice diameter

Injection velocities

0.37 to 1.69

0.92 to 1.58

0.76 and 1.32

30 to 90 degrees

O.ll? to 0.236 cm

(0.046 to 0.093 inch)

6.7 to 20.2 m/sec

(22 to 66 ft/sec)

MULTIPLE-EI_EMEN1, LIKE-DOUBLETSTUDY

Background

Impinging liquid rocket engine injector patterns have traditionally been designed

around the concentric ring approach -- a series of concentric, annular propellant

manifolding passages that feed corresponding concentric rings of fuel and oxidizer orifices, with impingement occurring between orifices of the sameor adJacent rings. This approach has been used primarily because of manufacturing constraints. With the development of more sophisticated fabrication techniques

(e.g., electrical discharge machining and electrodepositlng processes), a greater

degree of flexibility is now possible in injector design. One of the recently

developed configurations identified by injector specialists at Rocketdyne is the

"repeating box" pattern. In this pattern, the injector face is divided into a

grid of square boxes. Impinging elements (usually llke-doublets or llke-trlplets)

of one propellant type (fuel or oxidizer) are positioned on the box perimeters,

while impinging elements of the other propellant type are positioned in the box

centers. The repeating box pattern has been likened _to an impinging analog of a

coaxial element array, with the spray fans of one propellant type surrounding the

fans of the other type.

_J

RI/RD85 312

III-47
Subject Injector for LOX/RP-I Booster

The subject injector for this series of mixing tests has been proposed as a candidate injector for a large LOX/RP-I booster. The proposed deslgn features a

llke-doublet version of the repeating box pattern, shown schematically in

Table 3-8. Each square box includes four llke-doublets on the perimeter (one per

side) surrounding one llke-doublet In the center. The "box" (perimeter) doublets

are each shared by two squares, making the ratio of "box" doublets to "surrounded" doublets four halves to one, or more simply, two to one.

The principal deslgn speclflcatlon5 for the LOX/RP-I booster injector are also

listed In Table 3-8 -- LOX flow rate = 3?70 Ib/sec, RP-I flow rate = 1440 Ib/sec,

injector pressure drop = 400 psld (both sides) -- corresponding to a chamber

pressure of 2000 psla. An additional 50% throttling case Is also of interest,

with a reduced injector pressure drop of lO0 psl corresponding to a chamber pressure of lO00 psla. Two InJectant configurations are possible -- the "fuel box

surrounding oxidizer doublet" configuration or the "oxidizer box surrounding fuel

doublet" configuration. Orifice sizes for the two configurations are given In

Table 3-8. The latter configuration (oxidizer surrounding fuel) is thought to

produce better mixing, while the former (fuel surrounding oxidizer) may be more

desirable for boundary layer cooling reasons.

Cold-Flow Testing

The objectives of cold-flow testing were to compare mixing efflciencles between

the two InJectant configurations, between 40-degree and 60-degree angles of

impingement, and between a number of different throttling cases. The experimental

apparatus employed was the same as that described in the first section of this

report, with water used as the fuel slmulant and TRIC as the oxidizer simulant.

Turbine flowmeters were used to monitor flow rates. Because of the large size of

the proposed hot-flre injector, obviously only a limited portion of the spray

pattern could be sampled with the 4.13 by 6.35 cm (I-5/8 by 2-I/2 inches) collection grid. The representative portion selected as the "unit mixing cell" was a

single box (four perimeter elements surrounding one central element) plus the two

adjacent elements whose spray fans are seen to interact wlth the fans of the box.

RI/RDBS-312

III-48
TABLE3-8. DESIGNSPECIFICATIONSFORPROPOSEDLOX/RP-I BOOSIERINJECTOR

%..J

O0 • • 0 •

0 ° ° • • 0 ° ° 0 i

O0 O0 0 •

• • • • • •

• • 0 °° O • • X

• 0 • O O O 0-----_ • • • •

O0 O0 • 0

NOMINAL PATTERN:

SCHEMATIC DRAWING OF

REPEATING BOX PATTERN

I ORIFICES OF "BOX" DOUBLETS

• ORIFICES OF "SURROUNDED DOUBLETS

OVERALL INJECTOR DIAMETER = 35.2 INCHES

CHARACTERISTIC LENGTH (X) = 0.30 INCH

"FUEL BOX SURROUNDING OXIDIZER DOUBLET"

NOMINAL PATIERN DESIGN PARAMETERS:

DO = 0.051 INCH NO = 22,500

Df = 0.024 INCH Nf = 45,000

FULL IN3ECTOR PRESSURE DROP = 400 PSI (BOTH SIDES)

mo = 3770 LB/SEC mo PER ORIFICE = 0.168 LB/SEC

mf = 1440 LB/SEC mf PER ORIFICE = 0.032 LB/SEC

50% THROTILE: IN3ECIOR PRESSURE DROP = lO0 PSI (BOTH SIDES)

re

_o = 1885 I_B/SEC mo PER ORIFICE = 0.084 LB/SEC

l_f = 720 LB/SEC l_f PER ORIFICE = 0.016 LB/SEC

REVERSE PATTERN: "OXIDIZER BOX SURROUNDING FUEL DOUBLET"

REVERSE PATFERN DESIGN PARAMETERS:

DO = 0.036 INCH NO = 45,000

Df = 0.034 INCH Nf = 22,000

FULL IN3ECTOR PRESSURE DROP = 400 PSI (BOTH SIDES)

mo = 3770 LB/SEC _o PER ORIFICE = 0.084 LB/SEC

_f = 1440 LB/SEC _f PER ORIFICE = 0.064 I_B/SEC

50% THROTILE: IN3ECTOR PRESSURE DROP = lO0 PSI (BOIH SIDES)

_o = 1885 LB/SEC _o PER ORIFICE = 0.042 LB/SEC

_f = 720 LB/SEC _f PER ORIFICE = 0.032 tB/SEC

RI/RD85-312

III--49
A schematic drawing of this representative unit mixing cell is shown In

Table 3-9. This mixing cell was selected for cold-flow modeling primarily for

low cost and ease of fabrlcatlon. However, the small size of a single box (0.30

by 0.30 inch In the hot-fire deslgn) was not particularly compatlble with the

collection grld dimensions (O.125-1nch-square tubes) if adequate mixing resolution was to be ensured. Thus, the cold-flow models designed for these tests were

photographically scaled-up versions of the design-scale unit mixing cell. The

baseline scaling factor chosen was 2.5, for this reason: an enlargement factor

of 2.5 produced a unit mixing cell with sides of 0.75 inch, which could be

aligned above the collection grid to coincide exactly wlth a 6 by 6 array of the

O.125-1nch tubes. In this manner, a 36-sample mixing efficiency measurement would

be determined for a representative square of the overall repeating box spray

pattern.

Four models were fabricated for cold-flow testing, as shown schematically In

Fig. 3-21. The baseline model was a 2.5-scale, "fuel box around oxidizer doublet" mixing cell with Implngement angles of 60 degrees. Two other 2.5-scale

models were fabricated: the first a 40-degree impingement angle version of the

baseline, and the second a reversed configuration ("oxidizer box surrounding fuel

doublet") model. To produce the 40-degree impingement angle modification of the

baseline model, the orifices of each doublet palr were moved slightly farther

apart, resulting in an increase In impingement distance from 0.325 Inch

(60-degree case) to 0.515 Inch (40-degree case). The fourth model fabricated was

a 1.25-scale version of the baseline model (aligned above a 3 by 3 array of collection tubes during testing). This model was tested primarily to assess the

validity of acquiring mixing data from scaled-up models of the subject injector.

The majority of mixing tests were conducted on the 2.5-scale models, to effect

mixing efficiency comparisons between the two injector configurations, the two

impingement angles, and a number of mass flow throttllng cases. The 1.25-scale

model was then tested under conditions congruous to those employed In a specific

baseline model run. A comparison between the 1.25-scale and the 2.5-scale test

results would then be an indication of whether photographically scaled-up models

can be used to accurately assess mixing characteristics of their small-scale

counterparts.

RI/RD85-312

III-50
TABLE3-9. DESIGNSPECIFICATIONSFORCOLD-FLOWMODELS

CASE l:

NOMINAL PATIERN:

DO _ 0.064 INCH

Df ; 0.030 INCH

• •

t SCHEMATIC DRAWING OF "UNIT CELL"

• • • • • Xi REPRESENIATIVE OF BOX PAITERN

(l BOX PLUS 2 ADJACENT DOUBLETS)

SCALING FACTOR = 1.25 (CHARACTERISllC LENGIH (X) = 0.3?5 INCH)

"FUEL BOX SURROUNDING OXIDIZER DOUBLET"

NO = 6

Nf = 8

_o PER ORIFICE = 0.26 LB/SEC TOTAL _o = 1.57 LB/SEC

_f PER ORIFICE = 0.05 LB/SEC TOTAL _f = 0.40 LB/SEC

CASE 2: SCARING FACTOR = 2.50 (CHARACTERISTIC LENGTH (X) = 0.?50 INCH)

NOMINAL PATTERN: "FUEL BOX SURROUNDING OXIDIZER DOUBLEI"

DO = 0.128 INCH NO = 6

Df = 0.060 INCH Nf = 8

mo PER ORIFICE = 1.05 LB/SEC TOTAL mo = 6.30 LB/SEC

mf PER ORIFICE = 0.20 I_B/SEC TOTAL _f 1.60 LB/SEC

CASE 3: SCALING FACTOR = 2.50/CHARACTERISTIC LENGTH (X) = 0.750 INCH)

REVERSE PATIERN: "OXIDIZER BOX SURROUNDING FUEL DOUBLEI"

REVERSE PAITERN DESIGN PARAMEIERS:

DO = 0.090 INCH NO = 8

Df = 0.085 INCH Nf = 6

mo PER ORIFICE = 0.52 LB/SEC

mf PEr ORIFICE = 0.40 I_B/SEC

TOTAL mo = 4.16 LB/SEC

IOTAL mf = 2.40 LB/SEC

"v

Rl/RD85-312

IIl-Sl
V

MODEL 1 (BASELINE)

o O

• o O0 0

0 o ° oO

o o

O DFUEL = 0.060 INCH

• Dox = 0.128 INCH

IMPINGEMENT ANGLE = 60 DEGREES

SCALING FACTOR = 2.50

MODEL 3

0 O • 0

0 0

0 • • 0

MODEL 2

O•

Oo5O eeo

ODFUEL = 0.085 INCH

• Dox = 0.090 INCH

IMPINGEMENT ANGLE = 60 DEGREES

SCALING FACTOR =2.50

O DFUEL = 0.030 INCH

• DOX = 0.064 INCH

IMPINGEMENT ANGLE = 60 DEGREES

SCALING FACTOR = 1.25

MODEL 4

ooO %o

• •

• Oo oOO

O DFUEL = 0.060 INCH

• Dox = 0.128 INCH

IMPINGEMENT ANGLE = 40 DEGREES

SCALING FACTOR = 2.50

V

Figure 3-21. Schematic Drawing of the Four

RI/RD85-312

III-52

Cold-Flow Models

V
|

T_w

-v

Mass flow specifications for the test models are presented in Table 3-9, based on

design specifications for the hot-flre injector. These conditions were calculated

in the following manner: first, it was noted that a 2.5-scale increase in orifice

diameter translates to a 6.25-scale (2.5 squared) increase in orifice area, and,

therefore, a 6.25-scale increase In mass flow per orifice. (In the 1.25-scale

case, the orifice diameter increase translates to a 1.56-scale mass flow

increase.) Second, the total hot-flre mass Flow rates were reduced to flow rates

per individual orifice (Ib/sec/oxldlzer orifice and Ib/sec/fuel orifice), then

these values were scaled up by a factor of 6.25 (or 1.56), in the translation to

cold-flow orifice sizing. Finally, the translated mass flow per orifice values

were multiplied by the number of respective orifices included In the cold-flow

models (six or eight, depending on the InJectant configuration). Thus, the mass

flow conditions specified In Table 3-9 represent constant mass flux enlargements

of the corresponding Table 3-8 conditions. However, because of pressure and

tankage constraints, these target cold-flow conditions could not be experimentally achieved, and the tests were instead conducted over a reduced mass flow

range of approximately lO to 15% of the Table 3-9 values.

VTest Results

A series of II test runs was performed with the four injector models, the results

of which are presented in Table 3-I0. For each run, the test conditions are presented first in terms of TRIC (oxidizer) and water (fuel) injection rates for the

model, with a corresponding injected mixture ratio; and additionally in terms of

the total mass flow rates of each simulant projected for an entire injector comprising model-slze orifices. For example, in test run l, the projected TRIC mass

flow for the overall injector was calculated by dividing the injected (model)

TRIC flow rate (0.375 Ib/sec) by the number of TRIC orifices (8), and then multiplying the result by the number of oxidizer orifices composing the overall

injection pattern -- 45,000.

The baseline collection distance employed in the test runs was 2 inches, with two

exceptions: In the one run conducted wlth the 1.25-scale model, the collection

distance was correspondingly reduced by half, to l inch, and in one run conducted

V

RI/RD85-312

III-53
TABLE3-10. RESULTSOFTHETESTSERIESFORMULTIPLE-ELEMENT LIKE-DOUBLETMODELS

=o

.,,.a

r'o

MODEL

COLLECTION

RUN DISTANCE

NUMBER (INCH)

2.50 SCALE - OXIDIZER BOX

SURROUNDING FUEL DOUBLE1

MODEL 3

60 DEGREE IMPINGEMENT ANGLE 4* 2

2.50 SCALE - FUEL BOX 5 2

SURROUNDING OXIDIZER

DOUBLET

MODEL 4

INJECTED

nITRIC

FOR

MODEL

(LB/SEC)

0.375

40 DEGREE IMPINGEMENT ANGLE 8*

2.50 SCALE - FUEL BOX 9*

SURROUNDING OXIDIZER

DOUBLET

MODEL 1

60 DEGREE IMPINGEMEN1 ANGLE lO

1.25 SCALE - FUEL BOX

SURROUNDING OXIDITER

DOUBLET

0.443

0.495

O. 559

0.572

6 2 0.668

7* 2 0.832

1.5 0.851

2 0.835

2 0.741

MODEL 2 II l 0.210

60 DEGREE IMPINGMEN1 ANGLE

INJECIED

mH20

FOR

MODEL

(LB/SEC)

0.215

0.255

0.287

0.325

0.146

0.170

0.216

0.216

0.217

0.190

0.055

INJECIED

MIXIURE RATIO

FOR MODEL

1.74

1.74

1.72

1.72

3.92

3.93

3.85

3.94

3.85

3.90

3.82

PROJECTED

nITRIC

FOR

INJECTOR

(LB/SEC)

2112

2486

2787

3145

2143

2505

3122

3195

788

PROJECTED

mH20

FOR

INJECTOR

(LB/SEC)

806

956

1076

1219

821

956

1215

1215

1219

I069

308

PROJECTED

MIXTURE RATIO

FOR INJECTOR

2.62

2.60

2.59

2.58

2.61

2.62

2.57

2.63

2.57

2.60

2.56

*DENOTES TESIS CONDUCTED A1 BASELINE CONDITIONS FOR 2.S-SCALE MODELS

MIXING

EFFICIENCY

0.86

0.90

0.89

0.90

0.72

0.78

0.82

0.70

0.90

0.89

0.92

( ( (:
on the 2.5-scale, 40-degree model, the collection distance was reduced to 1.5

inches with flow conditions held constant, in order to assess the effect of collection distance variation on mixing efficiency. Test runs performed at the

chosen baseline injection conditions are marked with an asterisk for comparison

purposes. The first four tests were conducted, using the 2.5-scale, oxidizeraround-fuel model, at four mass flow conditions: 67% of baseline, ?g% of basellne, 89% of baseline, and baseline. The resulting mixing efflclencles were not

slgnlflcantly affected by the range of throttling -- the increase from 67% base11ne to baseline was accompanied by a 4% increase in mixing efficiency, from 86

to 90%. Tests 5 through ? were conducted, using the 2.5-scale, 40-degree, fuelaround-oxldlzer model, at three mass flow conditions: 69% of baseline, 80% of

baseline, and baseline. The results of these three tests contrasted wlth the

initial four in two ways: first, the increase in flow rates from 69% baseline to

baseline resulted in a I0% increase in mixing efficiency (72 through 82%), and

second, the mixing efficiency at the baseline condition (82%) was 8% lower than

the corresponding oxidlzer-around-fuel baseline value (90%). Test 8 was performed

to assess the effect of collection distance variation on mixing efficiency, wlth

flow conditions held near the baseline values. The decrease in collection distance from 2 to 1.5 inches resulted in a mixing efficiency decrease from 82% to

70%. This result indicates that within a short distance downstream of the

impingement plane, mixing efficiency measurements are very sensitive to collection distance in multiple-element injector studies. Tests g and lO were conducted, using the baseline model - 2.5-scale, 60-degree impingement, fuelaround-oxldlzer. Two mass-flow conditions were tested: 89% of baseline and

baseline, wlth resulting mixing efficlencles insignificantly affected by the

throttling range.

_i ̧

v

Comparison of runs l through 4 wlth runs 9 and lO shows virtually the same m_xlng

characteristics for the nominal (fuel-around-oxldlzer) and reversed (oxidizeraround-fuel) versions of the 60-degree baseline modeil In both cases, mixing

effIclencles near 90% resulted, with Inslgnlflcant mixing variation resulting

from mass flow throttling. Comparison of runs 5 through ? with runs 9 and lO

shows s_gnlflcantly lower mixing efflclencles for the 40-degree model. This

effect may be partially explainable in terms of implngement dlstances --In the

40-degree model, a 2-1nch collection distance translated to 3.9 impingement

RI/RD85-312

III-55
lengths, while In the 60-degree model, 2 Inches translated to 6.15 Impingement

lengths. Thls value (6.15 lengths) corresponds to a collection distance of 3.17

lnches for the 40-degree model. Given the Increase In mlxlng efficiency from 70%

to 82% for an increase In collection distance from ].5 Inches (2.9 lengths) to 2

Inches (3.9 lengths) for the 40-degree model, It Is possible that an Increase In

collection dlstance from 2 to 3.17 inches would result In mixing efftclencles

exceeding 90%. Thus, the poorer mlxlng performance of the 40-degree Impingement

angle lnJector may be caused by the greater Impingement distance of that Injector.

Flnally, test run 11 was conducted on the 1.25-scale model at flow conditions

congruous to the baseline values of run 9. A value of 92% resulted from the

9-sample mlxlng efficiency measurement, as compared to the 36-sample value of 90%

for run g. This result supports the valldlty of baslng general conclusions on

the mlxlng characteristics of a small-scale Injector pattern on data obtained

from scaled-up models.

V

In attempting to assess the qualtty of the mixing provlded by these elements, tt

ls important to recognize the followlng:

,

2.

,

These results are for 2.5 scale models of the injectors.

The collection distance was not scaled. That Is, the traditional 2-1nch

collection dlstance was employed.

Measured mlxlng efficiency must lncrease wlth collection distance,

especially with unlike, multiple-element injectors. (Propellants would

certainly not become unmixed as they move downstream.) The one test

performed to assess collection distance effects on mixing (test 8) Indicates a considerable change _n mlxlng efflclency for only a one-halflnch variation In collection distance.

If the co]iectlon distance were scaled In proportion to the injector, the measurements wouid be made at S lnches. The collection grld employed In these tests

was too small to allow testing at this collection dlstance. Certainly the measured mixing efftciencies at the 5-Inch location would be considerably improved

over these test results. Thus, these injectors would have considerably better

mlxlng efflclencles than were measured In thls program.

RI/RDB5-312

III-56
Conclusions

L

The following set of conclusions was Inferred from data generated In the present

study of the proposed like-doublet repeating box Injector pattern:

•

o

.

°

.

The nominal (fuel-around-oxidizer) conflguratlon and the reversed

(oxidizer-around-fuel) configuration are virtually Interchangeable with

respect to mlxlng efficiency.

The mlxlng efficiency of the 40-degree Implngement angle Injector was

significantly poorer than the 60-degree Impingement angle Injectors.

Larger Impingement angles (and/or shorter Impingement distances) appear

to promote mlxlng of the fans of the elements In shorter dlstances.

In the 40-degree Impingement angle verslon of the baseline model, mixing

efficiency was significantly Increased (72% to 82%) by an Increase In

mass flow from 67% baseline to baseline conditions. Both the 60-degree

nominal basellne model and the 60-degree reversed baseline model produced mlxlng efftciencies insensitive to throttling over the ranges

tested.

General conclusions regarding mixing characteristics of the subject

injector pattern can be inferred from data obtained wlth scaled-up

models.

Mixing efflctencles of these elements would be considerably better If

the collection distance were scaled with the injector. The actual mixing

efftciencies of these injectors are presumed to be considerably better

than reported here.

RI/RD85-312

III-5/
V
IV. GAS/LIQUID MIXING TESTING

During the first portion of this program, several of the baseline gas/llquld

elements (elements 2, 3, and 4) were subjected to gas/llquld, cold-flow, mixing

efficiency measurement tests. Several flow conditions were tested for each element and the results were presented in the interim report (Ref. l-l). The

measurement technique employed in these tests was the standard "impact probe"

technique that has been in use for over 15 years. This technique employs a

large, total-pressure, pitot tube, which acts as a total pressure measurement

device in the standard fashion, and also serves to collect a portion of the

liquid spray. The total pressure is used to compute gas velocity and gas mass

flux, and the rate of liquid collection is used to compute liquid mass flux.

When such measurements are made at various locations throughout the spray field,

a map of mass flux and mixture ratio distributions can be constructed, and mixing

efflclencies computed. A schematic of the test apparatus is presented in

Fig. 4-I. Details of the test and data analysis methods, and the results, are

presented in Ref. l-l.

Another output of the data analysis codes is the collection efficiency. Collection efficiency is simply the ratio of the mass flow rate of each fluid computed

from the measured fluxes, to the actual mass flow rate of each fluid. As noted

in Ref. l-l, these collection efflclencles were sufficiently far from the optimum

value of l, to be of some concern. The collection efflclencles from these tests,

for both fluids, are presented in Table 4-I. Liquid collection efflciencles were

generally low and gas collection efflclencles were high. In some tests, only

about one fourth of the injected liquid was being accounted for, and in other

tests over twice as much gas was being measured as was being injected. Another

concern regarding these gas/llquld mixing test results was their applicability to

higher chamber pressure conditions. To investigate the effects of chamber pressure on mixing data, a series of hlgh-pressure tests were planned. These were to

be carried out in the hlgh pressure test vessel that was intended for use in the

atomization testing subsequently described. However, before proceeding to

higher-pressure testing, it was deemed necessary to resolve the problem(s)

responsible for the poor collection efflclencles. It was anticipated that these

problems would be even more severe at higher pressure.

RI/RD85- 312

IV-I
v

PROPELLANT

SIMULANT

PROPELLANT

SIMULANT

ELEMENT

FLOW

MIST

STATIC

INJECTOR

MODEL

TRANSPARENT

CHAMBER

PROBE

(PITOT)

)BE

POSITION

PROBE

VELOCITY

HEAD

LIQUID

SAMPLING

CONTAINER

Flgure 4-1. Cold-Flow Gas/Liquld Mixing Measurement System

RI/RD85-31 2

IV-2
TABLE 4-I. COLLECTION EFFICIENCY RESULTS

__J

TEST CASE nCOL LIQ nCOL GAS

TRIPLET 2

NOMINAL HOT-FIRE MOMENTUM RATIO (O/F)

÷20% NOMINAL MOMENTUM RATIO

-20% NOMINAL MOMENTUM RAIIO

*NOMINAL MOMENTUM RATIO

PENTAD 3

NOMINAL HOT-FIRE MOMENTUM RATIO

*20% NOMINAL MOMENTUM RATIO

L20% NOMINAL MOMENTUM RATIO

*NOMINAL MOMENTUM RATIO

NOMINAL MOMENTUM RATIO

TRIPLET 4

NOMINAL HOT-FIRE MOMENTUM RATIO

•20% NOMINAL MOMENTUM RATIO

-20% NOMINAL MOMENTUM RATIO

0.48

0.52

0.44

0.91

.95

.85

l .03

0.87

0.93

0.33

0.29

0.23

*50 PSIG BACKPRESSURES; ALL OTHERS 25 PSIG

1.52

1.59

1.69

1.94

1.49

1.19

2.42

0.64

1.15

2.07

1.26

1.78

To provide a hlgher-pressure test capability and eliminate the causes of the poor

collection efficiency, a number of potential problems were identified, analyses

and experimental studies were performed, and equipment modlfIcations implemented. These are briefly described:

l . High-Pressure Testing Requirement: Gas/llquld mixing testing must be

performed at high pressures to allow reasonable simulation of injected

gas density. This requirement was especially important For planned

testing of coaxial injectors. Thus the hlgh-pressure atomization test

vessel was modified to allow its use for gas/llquld mixing tests.

. Probe and Pressure Line Flooding: Past low-pressure test results

indicated that the test probe may have been occasionally fi111ng with

water. A new probe was designed and constructed to minimize this

r _

RI/RD85-312

IV-3
.

.

,

possibility. Water separators were Installed to capture any water In

the pressure gage lines.

Probe Overpressure: Droplets approaching and entering the probe during

total pressure measurements transfer momentum to the stagnated gas,

thereby increasing the measured pressure. Following study of the work

of Dussord and Shapiro (Ref. 4-I), an analytical technique was devised

to estimate this effect and incorporated into the data analysis code.

Isoklnetlc Liquid Sampling: If the gas velocity through the probe tlp

Is significantly different from that adjacent to the probe, then the

quantity of liquid collected wlll not be representative of that flowing

through that area when the probe Is not present. The droplets tend to

follow the gas streamlines, especially at high pressures, and go around,

rather than into, the probe. To provlde'Isoklnetlc (constant velocity)

liquid co]lectlon, the gas flow through the probe was measured and controlled so as to match the measured gas velocity (based upon previous

total to static pressure measurements). Thls was accomplished wlth a

calibrated orifice installed In a llne that vented the sample bottle as

shown In Flg. 4-2.

Liquid Accumulation _n Lines Prior to Start of Collectlon: Since the

collection probe must be In place prior to the start of flow through the

injector, and since sampling must occur some tlme after steady flow has

been established, liquid may accumulate in the sampling lines. To prevent this, a "dummy" sampling vessel was installed as shown In Flg. 4-2.

Prior to the obtaining of a liquid sample, the injector flow was stabilized and the vent valve Set to provide a predetermined gas velocity

through the probe tip. At the start of the liquid collection time, the

flow was diverted from the "dummy" collection vessel to the identical

sample collection vessel through the three-way ball valve. Thus steady

flow conditions were maintained within the probe and sample llne.

RI/RD85-312

IV-4
DUMMY

COLLECTION

VESSEL

PRESSURIZED

TANK

I NJ ECTO R

PROBE

LIQUID

COLLECTION

LINE

WALL

STATI C

TAP

3-WAY

GAUGE

SEPARATORS

LIQUID

COLLECTION

VESSEL

GAUGE

_._ DRILLED

ORIFICE

VENT VALVE

Figure 4-2. New Gas/Liquid Mixlng Measurement Test Apparatus

RI/RD85-312

IV-5
6. Static Pressure Measurement Location: In earlier testing, the static

pressure measurement (required for gas mass flux mapping) was made at

the test vessel wall. This simplified measurement hardware and was

believed to produce only minlmai error. Measurementsof static pressure

across the chamber indicated that only minimal variation of static pressure occurred. However, it was anticipated that this error could be

substantial at low velocity or high chamber pressure conditions.

Accordingly, the gas/llquld mixing assessment procedure was modified to

include a mapping of local static pressure (relative to the wall static

tap). Thus the local total pressure (also referenced to wall static)

could be related to the local static pressure to obtain more accurate

local gas mass fluxes.

7. Data Reduction Programs: To ensure the accuracy of the computer code

used to convert the gas/llquld mixing measurement data to mass flux and

mixture ratio distributions and plots, and to compute mixing and collection efflclencles, a set of data was analyzed by hand and compared to

code output. The sensitivity of the code to the quantity of data

(number of measurements made) was also assessed. Results indicate that

the code performed accurately, and the results were relatively insensitive to any reasonable quantity of data provided.

8. Reclrculatlon: This appears to be the major problem with gas/llquld

mixing measurements. The entrainment and reclrculatlon of gases from

outside the spray add considerable mass (several times the fuel gas

flow) to the fuel gas. The traditional means by which this problem is

circumvented is the addition of a curtain of flowing gas around the

injector. This flow is referred to as curtain flow, purge gas, or bleed

gas, and it serves to prevent reclrculatlon into the measurement

region. The oxygen, content of the curtain flow gas is made different

from the fuel gas, thus allowing (via oxygen concentration measurements)

the curtain gas flow to be "subtracted out" of the measured total gas

l r

flow. Thls approach requires large curtain gas flow rates (in excess of

ten times the injector gas flow). For hlgh-pressure testing, the

quantity of curtain gas flow wou_d far exceed the existing hlgh-pressure

RI/RD85- 312

IV-6
nitrogen flow capabilities of the atomization and mixing test cell.

Furthermore, there was considerable concern that the curtain flow

significantly retarded the spreadlng of the fuel gas, thus affecting the

fuel gas mass flux distribution.

Prior to the installation of an expensive hlgh-flow-rate, hlgh-pressure, nitrogen

supply system to provide adequate curtain flow, it was deemed advisable to experImentally evaluate the changes made in the measurement procedures and apparatus

(items l through 6). Accordingly, a hlgh-backpressure (800 pslg) mixing test was

performed with the triplet element (element 2) at high flow rates. Since sufflclent curtain gas flow was not avallable, a small narrow-angle, glass cone was

placed about the spray in an effort to minimize reclrculatlon. The results of

this test were disappointing. Liquid collection efficiency was 223%, gas collectlon efficiency was 719%, the total pressure was observed to fluctuate and vary

over a wide range, and the results were not generally satisfactorily repeatable.

The major problem is believed to be reclrculatlon, which the glass shroud did not

prevent. The injector flow was visually observed to fluctuate wildly, and despite all preventive efforts, water in the pressure lines is suspected to have

affected the measurements. Pitot-statlc pressure measurements in heavily loaded,

two-phase flows are very difficult, and there is little experience upon which to

draw.

As a result of these findings, the following actions were undertaken:

.

Experiments were performed to demonstrate/determlne the ability to make

accurate pitot-statlc pressure measurements in known, well-deflned, nonreclrculatlng gas flows with liquid loading.

2. Analyses were performed to assess the effect of curtain flow on the fuel

gas distribution.

z

F  ̧

Pending satisfactory results from these efforts, it was intended that a highpressure nitrogen system would be "plumbed into" the test facility to provide the

necessary curtain flow. These two efforts and their results are presented in the

following.

RI/RD85-312

IV-7
MEASUREMENT CAPABILITY ASSESSMENT TESTS

The objective of these tests was to determine whether the basic measurement technique, with all incorporated refinements, was capable of accurately characterlzing mass flux profiles in an ideal two-phase flow field, in the absence of

reclrculatlon. The test apparatus used to evaluate the optimum mixing measurement techniques developed under this contract is shown schematically in Fig. 4-3.

A 40-horsepower blower was employed to produce a uniform, metered air flow

through an 8-1nch duct. At the exit section of the duct, a conical spray fan of

water droplets was introduced by using a Delavan nozzle. The traversing assembly

necessary to position the sampling probe was mounted on the duct outlet.

The sampling matrix comprised 128 data points -- 16 points spaced 0.25 inch apart

from duct centerllne to a radius of 3.75 inches, along 8 rays spaced by 45-degree

increments. Four separate flow-fleld mappings were performed to obtain the ultlmate two-phase, flow-field characterization. Initially, a local flow-fleld

static pressure versus atmospheric pressure mapping was performed, using the

static port of a conventional traversing pltot-statlc tube in gas-only flow.

Since the two phase sampling probe has no local static pressure sensing capabilIty, its total pressure measurements must be referenced to the atmosphere. Thus

the local static pressure mapping provides the correction necessary to compute

true total pressure versus static pressure measurements at each point in the flow

field. Also, when the pitot-statlc probe was used in gas-only flow, a total versus static pressure mapping was obtained. This was then used to compute the duct

velocity profiles. These data indicated the flow was relatively axial (the

effects of the upstream tube elbow were apparent, however), and no reclrculatlon

was present. These velocity data were subsequently used as the reference for

Isoklnetlc probe inlet velocities during liquid collection.

Following pltot-statlc probe mappings, a gas-only total pressure versus atmospherlc pressure mapping was performed with the two-phase sampling probe. With

the application of the local static versus atmospheric pressure correction at

each point, the total pressure sensing capability of the larger two-phase probe

was compared with the pltot-statlc gas/llquld flow mapping. Gas total pressures

were recorded at each matrix point and liquid samples collected over 5-mlnute

RI/RD85-312

IV-8

W=¢
THERMOCOUPLE

PR ESSU R E

REGULATOR

PRESSURE

(200 PSI)

H20 SOURCE

TLE

VALVE

DELAVAN

NOZZLE

PROBE

TRAVERSING

ASSEMI

LIQUID

COLLECTION LINE

COAXIAL

ELBOW FITTINGS

TOTAL

PRESSURE

TAKEOFF

_PGAUGE

(101N. H20)

= i

3-WAY VALVE

LIQUID

3OLLECTIOI_

VESSELS

NEEDLE

VALVES

GAS

VENTING

LINE

8 IN. ID DUCT

PRESSURE

GAUGE

50 PSll

r"

m

Z_k P GAUGE

(40 IN, H20)

VENTURI

THROTTLING

VALVE

NEEDLE

VALVE

VACUUM

PUMP

GAS

FLOWMETER

(1500 STD IN.3/MIN)

BYPASS

VALVE

BLOWER

Figure 4-3. Gas/L1quld M1xlng Measurement Capability Assessment Test Apparatus
time intervals. Near Isoklnetlc liquid collection was produced at each sampling

point by venting the appropriate air flow rate through the gas flowmeter shown in

the schematic diagram, under the assumption of steady, incompressible air flow

through the probe and liquid collection lines. Results of these tests follow.

Known duct air flow rate

Known liquid flow rate

Pitot-statlc probe, gas only flow

Two-phase probe, gas only flow

Two-phase probe, gas/llquld flow

2.568 Ib/sec

0.089 Ib/sec

Gas collection efficiency = 1.094

Gas collection efficiency = 1.150

(No overpressure correction)

Gas collection efficiency = 1.160

Liquid collection efficiency = 0.926

(With overpressure correction)

Gas collection efficiency = 1.140

Liquid collection efficiency = 0.926

These measurements indicate that the gas/llquld mixing measurement techniques can

produce marginally acceptable results. (Fourteen percent more gas and 7.4% less

liquid were measured than were actually present.) However, these were very optimum test conditions (very lightly loaded with liquid - 3.3% water by weight, low

pressure, and well-deflned, nearly one-dlmenslonal flow). In addition, this

spray field was mapped far more extenslvely than would normally be done.

The next step in the verification of the measurement techniques employed was to

be a similar measurement at higher pressure. However, before this effort could

begin, results from the analysls of curtain flow effects on the In_ected gas were

completed.

V

CURTAIN FLOW EFFECTS ANALYSIS

While the preceding results indicated that two-phase mass flux measurement might

be possible in the absence of reclrculatlon, the effect of the reclrculatlon

preventing bleed gas (curtain flow) on these mass flux measurements had to be

ascertained. Reclrculatlon had to be prevented and the bleed gas technique is

RI/RD85-312

IV-lO
.y

the only known means by which this can be accomplished. However, if the bleed

gas significantly affected the fuel gas distribution, then the gas mass flux

distribution measurements would not be representative of those produced by the

injector in the combustor.

To examine this aspect of the problem, an analytical approach was employed. A

sophisticated and complex computational fluid dynamics code was used to model the

flow of fuel gas in the region between an injector and the normal measurement

plane (2 inches below the injector). The code selected for the task was the

Advanced Rocket Injector Combustor Code (ARICC) developed by Rocketdyne for NASA

under contract NAS8-34929, Turbine Drive Combustor Ignition and Durability Program, and described in Ref. 4-2 and 4-3. ARICC Is a tlme-marchlng, 2-D (axlsymmetric) computational fluid dynamics (CFD) code with provisions for spray mixing

and combustion. ARICC was developed from the Los Alamos CONCHAS - SPRAY code

(Ref. 4-4). Features of ARICC include:

Multlspecles mixing and diffusion

Lagranglan droplet dynamics

Droplet heatup and evaporation

Coupled gas - droplet mass, momentum, and energy transfer

Equilibrium and kinetic chemistry

Subgrld scale turbulence model

Choice of wall boundary conditions

The code is modular in format and modules can be added or deactivated, depending

on the problem requirements. For this case of gas Jet mixing, the chemistry,

11quld jet, and droplet modules were not used. Modeling the droplet dynamics

would have significantly increased the cost of the analysis.

The problem selected was an F-O-F triplet injector element flowing water and

nitrogen into a 900-psla environment. The chamber was 2 inches in diameter by 2

inches long. Injector operating conditions were:

RI/RD85 312

IV-ll
M/R = 0.49

M(H20) = 0.069 Ibm/s

M(N2) = O.141 Ibm/s

The orifice diameters were:

Dfuel = 0.063 inch

D = 0.045 inch

OX

(oxidizer orifice)

(fuel orifices)

The impingement angle was 50° and the impingement distance was 0.25 inch.

To address the problem in a timely manner using ARICC, several simplifications

were made in the representation of the injected flow. The major simplification

was to consider only the effect of purge gas on injected gas. (No liquid was

assumed to be present.) The water flow was replaced by a stream of water vapor

having the same axial mass and momentum flow as the liquid. The water vapor was

mixed with the nitrogen flow. Based on conservation of mass and axial momentum,

the velocity and diameter of a well-mlxed Jet of nitrogen and water vapor were

calcu- lated. For simplicity, the injected water vapor and nitrogen flow was

assumed to be a single specie at the average molecular weight for a flow at a

mixture ratio of 0.49. The mass fraction contours of injected fluid could in

this manner be easily tracked as a function of time. Three-dlmenslonal details

of the flow upstream of the impingement point could not be modeled with a

2-D/axlsymmetrlc flow analysis. In view of this, an axlsymmetrlc injection spud

was used to introduce the injected stream to the impingement point. The spud

diameter matched the diameter of the actual injector element disc that would be

used in the experiment. A spreading angle that is a function of radial distance

was imposed on the injected flow at the impingement point. The maximum angle was

45 degrees and corresponds to experimental observations. The mesh at the point

of injection was set up with boundaries approximately parallel to the injected

fluid streamlines, which served to minimize numerical diffusion. This

computational mesh is presented in Figure 4-4.

RI/RD85- 312

IV-12
:z

i l!i I:

\

i.

I

!

!

I

I

/jj iJ

iJ! l q

. I

|

!

!

i l

: I

II

!

!

I

fj

iI! r.

I l,

I

°

: r

" I

!.I

!i/

i

i

i

[

//

/

i

i

}

!l I l

"l\\\\\

!

Figure 4-4. Computational Mesh

RI/RD85- 312

IV-13

II

,//

i

!

i

I

i

iI

!

i!

II

/i

!

i

i

!

!

i
For the case with purge gas, a nitrogen gas flow of 2 lbm/sec was assumed to flow

uniformly over the 2-Inch-diameter chamber lnlet except for the area taken up by

the Injection spud. According to Becker et al (Ref. 4-5), the value of the

Curtet parameter that Is sufficient to preclude reclrculatlon, is 0.78. For the

selected triplet conditions and the indicated purge flow, Ct = 0.66. Thus the

purge gas flow used In this analysis was somewhat less than what would be

employed In a test.

The computations were run until near steady conditions prevailed. Results were

then plotted and comparisons of the injected gas distribution In each case were

made. Figure 4-5 illustrates the radial velocity profiles at 9.5 stream diameters downstream of the impingement point for both cases. An artificiality In the

injected velocity distribution shows up as an annular Jet In the velocity profile

for the case without curtain flow. An upstream reclrculatlon zone wake Is also

seen in this profile. The results with purge flow suggest a developing Jet profile. Fully developed coaxial Jet data scaled up from Abramovlch (Ref. 4-6) are

shown for comparison wlth the purge flow case prediction. The close agreement

wlth experimental data provides some degree of confirmation of these computations. Since the computed profile at the collection plane Is only 9.5 injected

stream diameters beyond the point of impingement, It Is not surprising that the

veloclty profile shows some evidence of a potential core.

Computed injected mass flux profiles are plotted versus radial distance In

Flg. 4-6 for the same axial location. Figures 4-7 and 4 8 show injected fluid

mass fraction contours for the two cases. Contour lines occur at steps of O.l In

mass fraction. From these two figures, It can be seen that the Jet without curtain flow diffuses much more rapidly In a radial direction than In the case wlth

curtain flow.

Velocity information plotted output from the ARICC code, at the time of these

computations, was inadequate for presentation here. However, the velocity information has been used to construct the plots shown In Fig. 4-9. These two

unscaled plots present the steady-state flow fields for the cases wlth curtain

flow (left side) and without curtain flow (right side). While the curtain flow

does not entirely eliminate the small reclrculatlon zone near the injector, the

RI/RD85-312

IV-14
L_

>-

FO

/

LU

>

150C --

1000 --

500 --

\

0

0

I

0 .5

[]

X = 4.13 CM

= 1.63 IN.

Y/Dj = 9.5

\

\

\

[]

\

\

q

\

\

[]\

\

\

\

[]

-- -- -SCALED DATA

FROM

ABRAMOVICH

FOR COF LOWING

JET

[] 2 LB/SEC

CURTAIN FLOW

O NO CURTAIN

FLOW

[]

O

O O

O

O

O O

I !

O

1.0 1.5

RADIAL DISTANCE (CM)

[]

O

I

2.0

[]

O

CHAMBER

LL

2.5

Figure 4-5. Velocity Proftles

RI/RD85-312

IV-IS
A

(o

LU

_0

X

(/}

<

rr

X

LL

(/3

_3

<

4O

30

2O

10

[]

[]

D

[]

O O

O

O

[]

O

[]

O

X = 4.13 CM

= 1.63 IN.

X/Dj = 9.5

[]

B

O O

D

1.0

RADIAL DISTANCE (CM)

O NO CURTAIN

F LOW

[] 2 LB/SEC

CURTAIN

F LOW

CHAMBER_

WALL

O

•l R

2.0

Flgure 4-6. Computed Mass Flux Properties

RI/RD85-312

IV-16

-- =

V
=

\

\

:\. .//'\

_LL_'_ _" A

m

\

V

I!

l

(

i t

YS 0.9 MASS FRACTION

.._ 0.8 /

% • ,.,-f •

!

I 7/_ I.,____oo,_,_o-r,o_

PLANE

Figure 4 7. Injected Species Mass Fraction Contours - With No Curtain Flow

Figure 4-8.

/_'X, 0.1 MASS

FRACTION

COLLECTION

PLANE

Injected Species Mass Fraction Contours - With Curtain Flow

RI/RD85-312

IV-17
COLLECTION

PLANE

F

WITH CURTAIN FLOW

J

NO CURTAIN FLOW

/

f

V

Figure 4-9. Gas Motion - With and Without Curtain Flow

RI/RD85-31 2

IV-18

V
=

flow at the collection plane is essentially axial. However, without curtain

flow, the recirculatlon zone near the injector is much larger, and a second

reclrculation zone draws gas from below the collection plane upward and along the

periphery of the chamber. Obviously the much greater reclrculatlon of the noncurtain flow case would promote additional mixing. This is the cause of the

greater spreading of the injected gas when no curtain flow is provided, as shown

in Fig. 4-7.

As a result of these computations, it can be concluded that the purge flow does

have a significant impact on the injected mass flux profile at steady state. The

purge flow causes a change in the reclrculatlon zone, resulting in a change in

the radial flux of injected material. In a flow where a reclrculatlon zone is

dominant, the radial flow of mass and momentum is convection dominated. For the

case where a purge flow suppresses the recirculatlon zone, less efficient turbulent diffusion controls the radial spreading of injected gas. Hence, the case

with purge flow has a much higher centerllne velocity and mass flux of injected

gas.

E

CONCLUSIONS

As a result of these efforts, it is concluded that:

1. The modifications to the gas/llquld mlxlng measurement procedure and

apparatus provide approximate and correct measurements at low

llquld-loadlng levels and low pressure, and in relatively one-dimenslonal flows with no reclrculatlon. Their ability to obtain correct

data at high llquld-loadlng levels and pressure is not known. Reclrculation must be prevented to obtain correct results.

2. lhe use of large quantities of curtain flow to prevent reclrculatlon

results in a very great modification of the fuel gas distribution,

thereby invalidating the measured results.

m_

RI/RD85-312

IV-19
Therefore, wlth regard to the utlllty and applicability of gas/liquid mixing

measurements, the following are concluded:

1. Single-element gas/llquld mixing measurements _ of llmlted value

for comparing the mixing performance of different elements or types of

elements, and also _ of very limited value for assessing the relative effects of injector geometry or flow varlables. Such measurements

serve only as a relative comparison of mixing efficiency.

2. The use of such measurement data as Input to the performance analysis

codes (e.g., SDER), or to establish design criteria (e.g., optimum

values of Elverum-Morey parameter) is not Justified. Further testing of

this type ls not recommended.

3. Effort should be directed toward the development of a means to assess

and measure lnJector mixing performance.

One very promising means by which thls may be accomplished ls the use of

multidimensional CFD codes such as ARICC. Such codes can already model cold-flow

gas motion wlth good accuracy, and can include the effects of droplets and combustion on gas dispersion. Modeling of the liquid phase (atomization, stream and

droplet breakup, and droplet motion) Is currently less developed, but efforts to

improve such models are under way.

Another means for the measurement of the mixing performance of injectors Is the

utilization of advanced combustion diagnostic techniques such as Raman spectroscopy. These diagnostic techniques offer the potentlal capability to directly

and nonlntruslvely measure combustor gas temperatures and compositions. Thus,

they could provide the first direct measurements of hot-flre mixing efficiency.

RI/RD85-312

IV-20
V. ATOMIZAIION

V

INTRODUCTION

Atomization data have long been recognized as one of the most Important inputs to

any spray combustion model. Both the droplet-slze and droplet-slze distribution

have been shown (Ref. 5-1 and 5-2) to have a large effect on the computed performance of a liquid rocket combustor.

Reference l-l contains a detailed assessment and description of available liquid

rocket combustor atomization correlations. In summary, the state of the art of

atomization modeling is generally inadequate to meet present needs. The physics

is only qualitatively understood at best. No quantitative theories exist. The

available data and correlations generally are of questionable validity and/or

utility. Many of the most critical parameters are unknown (e.g., combustion gas

veloclty field and multiple element effects) and/or are not simulated in tests

(e.g., gas densities, real propellant fluid properties, and combustion gas

motion). In addition, the measurement techniques generally used employ questionable assumptions or are incapable of sufficiently detailed or appropriate

measurements.

V

Recognlz_ng the importance of atomization data, and the poor quality of the presently available data, modelers have often used the initial droplet sizes as an

adjustable parameter to calibrate their codes. Thus, when the code computations

did not predict observed rocket combustor performance, the initial droplet s_zes

were modified to force agreement. This Is the case with the three major performance assessment codes: SDER, CICM, and TPP.

Obviously this degree of uncertainty about one of the most critical inputs to a

spray combustion code is not a satisfactory or acceptable situation. And yet, it

has been accepted for some time, and little effort has been directed toward

improvement. The last extensive program of liquid rocket atomizatlon research

was the "hot wax technique" study performed by Rocketdyne and completed in the

very early 19?Os.

RI/RD85 312

V-l
The neglect of such an important area of study may, In part, be the result of a

reduction In more basic liquid rocket research funding tn order to support engine

development efforts. It may also be caused by a lack of knowledge of (1) the

Importance of atomization data and (2) the inadequate level of quallty of available atomization data for combustor analysis codes. And flnally, neglect of this

Important research area Is caused, at least Inpart, by the lack of suitable

techniques for measuring droplet sizes accurately and rapidly. (Problems with

these techniques are discussed in detall In Ref. 1-1 and 5-3.)

V

Over the past ten years, the state of the art of droplet-size measurement techniques has improved considerably. Lasers have been employed In a variety of ways

to obtaln hitherto unavailable lnformatlon, and small computers have been

employed to rapidly convert this information to droplet-size data and to compile

and correlate these data. Droplet-size measurement instruments have been developed that utlllze such methods as photography, pulsed laser holography, X-ray

laser shadowgraphy, Fraunhofer diffraction, pulsed laser TV imaging, pulsed laser

photography, high-speed cinematography, and laser lnterferometry. Many of these

instruments and methods are "custom designed" by researchers for their own appllcatlons, but several have been developed and are sold commercially. These techniques have been employed In other atomization studies, primarily to evaluate the

atomization of dlesel, gas turbine, and larger commercial combustor fuel Injectors, as well as other spray devices (e.g., agricultural sprayers and flue gas

scrubbers).

v

In 1981, Rocketdyne began an In-house examination of these various drop-slze

measurement techniques and instruments, wlth the intent of applylng one of these

to the study of liquid rocket combustor injectors. These techniques were evaluated wlth regard to their capabilities in the following areas:

I. Nonlntruslve (no flow disturbance)

2. Large sample size. (Several thousand droplets must be measured to obtaln

accurate distributions)

RI/RD85-312

V-2
3. Speed of data acquisition, compilation, and analysis. (Technique must be

automated as manual methods are too expensive)

4. I0 to 500 micron droplet-slze measurement capability

5. No restrictions on test liquid properties

6. Vacuum to hlgh-pressure test capability. [(Requires that technique be

usable in pressure/vacuum vessel (i.e., window and access conslderatlons)]

?. Specific and small measurement location.

tlon of the spray)

(Will provide spatial resolu--

8. Temporal (flux-based) rather than spatial (concentratlon-based)

droplet-slze data. (An important and often overlooked consideration.

For a detailed description of this consideration, refer to Ref. 5-I

through 5-3, and ASTM standard E?g9)

g. Droplet velocity measurement capability

lO. Applicability to reacting flows

ll. Applicability to thlck/dense sprays

12. Commercial availability. (Rocketdyne did not wish to become developer

of such techniques, If possible)

Most of the techniques considered were relatively new, with little user experience or proven capability. Moreover, no standard spray exists by which to

measure the effectiveness, accuracy, et al, of a drop-slze measurement technlque. Thus it is necessary to rely upon drop generators that produce single

drop-slze (monodlsperse) and very dilute sprays, and upon the "reasonableness" of

the measurements of sprays, to assess the validity of an instrument. This

inability to verify the droplet-slze measurements of new techniques and

RI/RD85-312

V-3
Instruments Is a serious problem that continues to lnhlblt atomization studies,

as it did the Rocketdyne efforts.

By mid-1981, Rocketdyne had selected the technique of droplet sizing lnterferometry as the most promising technique to employ In the study of 11quid rocket

atomizers. Arrangements had been made to lease and subsequently purchase a droplet sizing lnterferometer (DSI), manufactured by Spectron Development Laboratories, and supporting equipment for Rocketdyne IR&D studies of injector

atomization characteristics. Also, a windowed pressure vessel was located at

Rocketdyne, which appeared satisfactory for high-pressure injector atomization

characterization. In October 1981, open-air tests of dllute sprays began. These

were part of an IR&O project to develop familiarity with, prove, and Implement

thls droplet sizing technique at Rocketdyne. Based upon the DSI manufacturer's

reports and the fact that over a dozen of these systems were then In use In the

U.S. and Canada, it was anticipated that this high-pressure droplet sizing capablllty would be easily and rapidly Implemented at Rocketdyne. Accordingly, when

this atomization and mixing study contract was announced by NASAMSFC, Rocketdyne

proposed to employ this advanced droplet sizing capability.

By the spring of 1982, the IR&O dllute spray studies had established that the DSI

dld not perform satisfactorily. Whlle the DSI was easily and repeatably able to

accurately determine the droplet size of a narrow stream of monodlsperse droplets, the Rocketdyne testing clearly demonstrated that spray test results

obtained with different instrument settings were quite inconsistent.

Ourlng the next year, Rocketdyne worked wlth the DSI developers at Spectron

Development Laboratories in the development and testing of a second-generation

DSI. The Rocketdyne primary function In this Joint enterprise was to evaluate

the new DSI and assist in the development of operating procedures. In

February 1983, It was determined that this second-generation DSI (using the visibility/Intensity technique later described) was performing satisfactorily.

Addltlonal effort was then begun to develop and test the high-pressure atomlzatlon test capability. The high-pressure, wlndowed test vessel, DSI and

associated equipment, and propellant slmulant tankage, plumbing, and control

RI/RD85- 312

V-4

V

V
hardware were installed in a test cell In the Rocketdyne Engineering Development

Laboratories. The test vessel (Fig. 5-1) Is a 6.2 RPa (900 psig), man-rated,

ASTR cylindrical pressure vessel with an inside diameter of 15.2 cm. The windowed section of the vessel is near the top. A single injector element could be

mounted in the center of the top of the vessel. Gaseous nitrogen flow (bleed

gas) was provided ctrcumferenttally about the element to reduce rectrculatton of

the droplets. The injected propellant stmulant and the bleed gas were exhausted

from the bottom of the tank. The monodtsperse droplet generator, which was

required for alignment of the DSI, was modified to mount in the center of the top

of the tank. A hlgh-pressure feed system, with a very low and precisely controlled flow rate, was designed and constructed for the monodtsperse droplet generator. The DSI alignment procedures were modified to account for the high-pressure

windows, and the ability Of the DSI to correctly measure the sizes of monodisperse droplet streams within the tank was confirmed.

i

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The first attempt to make hlgh-pressure, droplet-slze measurements quickly demonstrated the inadequacy of the bleed gas flow. Even at the maximum facility flow

rate, the bleed gas was incapable of preventing large quantities of liquid from

being reclrculated. Liquid was being drawn up along or near the chamber walis

and pulled back into the main body of the spray. This reclrculatlng liquid,

combined with the main body of spray, was far too dense to allow DSI measurements. The velocity of the" injected propellants Was creating a low-pressure

region near the injector face (in a manner similar to a Jet pump). Also, the

lower part of the tankwas _eXtremely turbulent and agltated. The reclrculatlng

gas was entraining some of this agitated liquid and flowing upward, along the

cylinder wall, toward this low-pressure region. A variety of baffles and open

celled, plastic foam wall liners were built and tested in order to minimize this

reclrculatlon. These were only partially successful. A more successful design

(presented in Flg. 5-2) consisted Of a combination of these baffles and foam,

together with a set of tubes' aligned axlaily along the periphery of the cylinder. This arrangement allowed the gases to flow through the tubes from the bottom of the tank to the t:op....... Foam _Was placed over the tube ends to reduce water

ingestion. This technique of controlled and filtered reclrculatlng gases greatly

reduced reclrcul ating (] ........... ....

RI/RD85-31 2

V-5
Flgure 5-1. High-Pressure Atomization Test Vessel

ORIGINAL PAGE ;_

RI/RDB5-312 _ POOR QUALMS(

V-6

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The DSI was realigned and callbrated, and attempts were again made to measure

droplet sizes in the center of the sprays produced by the gas/llquld injectors at

high pressure. Once more, it was found that the sprays were too dense. As will

be described in the following section, the DSI establishes a very small measurement region within the spray -- the brobe volume. As a droplet passes through

this region, its size and velocity are measured. However, only a single droplet

can traverse this probe volume at one time if the data are to be valid. The DSI

recognizes and discards invalid data. If the droplets are too closely packed

(i.e., the spray is too dense), then only rarely will a single droplet traverse

the probe volume. The gas/llquld injectors of interest produce such dense sprays

of very tiny droplets, that the DSI could not measure droplet sizes except near

the edge of the spray. Even the lowest flow rate element (triplet element 4)

produced sprays that were too dense. The droplets produced were smaller but Just

as densely packed. Three approaches were formulated to resolve this problem:

, Reduce probe volume size: The sizes of the laser beams forming the

probe volume were reduced to their minimum diameters. While this reduction improved the situation, it was not sufficient. V

2. Develop and use an intrusive probe: A complex cone-shaped probe was

designed, built, and tested to accomplish this. This probe is presented

in Fig. 5-3. The function of this probe was to intercept a small portion of the spray (the part entering the small end of the cone), separate it from the main body of the spray, and spread the droplets over a

larger area. All this must be accomplished without significant droplet

breakup or collision with the probe wall. Preliminary testing was

disappointing, and it appeared that the development of this technique

would be a long-term research project. Therefore, since schedules and

budgeting constraints were pressing, this approach was abandoned.

o Make measurements further downstream of the injector: As sprays move

downstream from the injector, they spread and the droplet density

decreases. Open-alr tests were performed that demonstrated which

RI/RD85-312

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measurements could be made far (as much as 1 meter) from a fully flowing

injector. Unfortunately the test vessel was designed for spray mappings

only a few (up to 10) centimeters below (downstream) of the injector.

Also, the test vessel had an inside diameter of only 15.2 cm, which was

insufficient to allow adequate spreading of the spray. Thus this

approach could not be applied In this test vessel. No other suitable

pressure vessel was available.

As a result of this unresolved problem, a decision was made to delete the

hlgh-pressure atomization testing and to perform all such tests unconfined (i.e.,

at atmospheric pressure). An open-alr test apparatus was constructed. The test

procedures, apparatus, rationale, plan, and results are later described In

detail. All this effort to establish this atomization test capability, as

described, was performed with Rocketdyne funding, and was not a part of the contractual program.

Thls decision to test at atmospheric pressure was driven by the prevlously discussed hardware and instrumentation (DSI) limitations. Thls approach was

employed only as a last resort, as it was recognized that the extrapolation of

the test results to the hlgh-pressure conditions of interest would be difficult.

Nearly all other atomization research programs have been similarly forced to

perform their tests at atmospheric pressure. As reported In Ref. l-l, essentially no hlgh-pressure atomization data exist for liquid rocket injectors except

for some very limited data obtained in support of the SSME development (i.e.,

coaxial injectors). Thus, a considerable need exists for hlgh-pressure

atomization data (one of the original goals of this program). Also, a lack of

data to support extrapolation of atmospheric pressure data to high pressure

remains (now one of the needs of this program).

V

DROPLET SIZING INIERFEROMETRY

Thls subsection describes the droplet sizing Interferometry technique employed in

this program. While a number of Interferometrlc techniques have been developed

to measure droplet sizes, thls discussion Is 11mlted to the combined

vlslbility/Intenslty (V/I) DSI technique employed on this program.

RI/RD85-312

V-lO
V

As with all DSI techniques, this technique employs a laser, beamsplltter, and

appropriate optics to cause the two beams to intersect at a point within the

spray where the drop sizes are to be determined (see Fig. 5-4). The intersection

of these laser beams produces an InterferenCe reglon-a series of fringes covering the region in space where the beams overlap. This region is referred to as

the probe volume. This probe volume may beTqulte small, perhaps as small as lO0

microns in diameter and a few hundred microns in length, and is of an ellipsoldal

shape. Collectlon optics are provided that image this tiny probe volume on a

photodetector. Now, as a droplet passes through the probe volume, it scatters

light (by refraction and reflection) onto this photodetector. This produces a

signal similar to the plot of intensity vs. time shown in Fig. 5-4. This signal

is referred to as a Doppler burst or signal, and consists of two components: an

ac component superimposed on a gausslan-shaped signal (the "pedestal"). The ac

componentresults from the passage of the droplet over the alternating bright and

dark fringes, lhe pedestal is caused by the gaussian Inten_slty profile of the

laser beams, lhe fringe spacing, which is determined by the optics of the system, and the measured frequency of the ac component can be used to compute the

velocity component of the droplet normal to the fringes, lhe velocity is simply

the fringe spacing (the distance the drop travels between fringes) divided by the

period of the ac slgnal frequency (the time required to travel that distance).

This technique is co_w_only referred to as laser doppler veloclmetry.

This Doppler signal can also be used to determine the size of the droplet through

the concept of visibility. Visibility is a measure of the size of the ac component relative to the pedestal. This relationship and its derivation are

described %n detail in Ref. 5-4 and 5-5. However, by referring to Fig. 5-5 it

may be possible for the reader to obtain a qualitative understanding of how vlsl

billty can be related to droplet size. Figure 5-5 presents typical signals that

may be obtained for a small droplet (upper plot) and a large droplet (lower plot)

passing through the probe volume. The small droplet produces a weaker overall

signal than the large droplet, simply because of its smaller size and consequent

lesser llght-scatterlng ability. In addition, the small droplet produces a sighal in which the ratio of the ac component to the pedestal (i.e., the visibility)

is high. Because the droplet diameter is much less than the fringe spacing,

RI/RD85-312

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smaller droplets have more fluctuation In their Doppler signal. When the droplet

Is in a bright fringe, the entire droplet is illuminated, and a relatively strong

signal is detected. When the droplet passes through a dark fringe, It scatters

almost no light, and the signal drops nearly to zero. Conversely, the large

droplet overlaps fringes so that it is never totally In darkness or completely,

brightly illuminated. Hence, lts signal does not "fluctuate as much. Thus it can

be seen that small droplets produce signals having hlgh ac components relative to

the pedestal (i.e., high visibility), and large droplets produce signals having

low ac components relative to the pedestal (i.e., low visibility). It Is important to recognize that visibility ls a relative measurement. Thus, If droplets

only pass through the edge of the probe volume, where the fringes are less

brlght, the strength of both the ac and the pedestal signals Is reduced. However, their respective magnitude relative to each other (I.e., the visibility)

remains constant.

It is also important to note that thls technique requires that only one droplet

at a tlme be present within the probe volume. Also, this technique is based upon

the assumption that perfect interference exists between the laser beams. The

beams must be of equal intensity and phasing so as to produce fringes of very

hlgh contrast (e.g., extremely black, dark fringes and bright, light fringes).

If the fringes become smeared (lower fringe contrast), so that the intensity of

the bright fringes Is reduced and the brightness of the dark fringes is

increased, then the ac component of all Doppler signals is decreased. However,

the pedestal Is relatively unaffected, and therefore, the vlslbillty Is also

decreased. Thls condition then makes all the droplets appear larger than they

really are.

v

The original DSI that Rocketdyne planned to utilize was based upon this visibility technique. While the DSI performed very well on a narrow stream of monodlsperse droplets, Rocketdyne testing demonstrated its inability to make accurate,

or even approximate, measurements In a distributed spray of modest density. The

problem was determined to be a result of poor fringe contrast. The major cause

of this poor fringe contrast was the passage of droplets through the beams prior

to their intersection. Thls action had the effect of selectively reducing the

intensity (or perhaps altering the phasing) of one or the other of the laser

RI/RD85-31 2

V-14

V
l

beams. Thus, the fringes were of poor contrast, and the droplet sizes measured

were often incorrect.

It was impossible to prevent the droplets from passing through the beams without

somehow shleldlng the beams from the spray (i.e., placing something intrusive

into the spray). This solution was not acceptable. Instead, a modification of

the visibility technique was employed to circumvent the problem. This modified

technique employs the absolute amplitude of the signal (i.e., the intensity of

the light scattered by the droplet) as an additional measure of droplet size.

The droplet visibility Is measured and the droplet size inferred from visibility

as previously described. In addition, the peak value of the signal (peak intensIty) is measured. This peak intensity is also a measure of droplet size, as

larger drops scatter more light. Thus two separate measures of droplet size are

obtained, which are compared by the DSI instrumentation and, if they are not

within tolerable agreement, that measurement Is rejected. Thus, the intensity

droplet-slze measurement is used to confirm the visibility droplet-size measurement. This combined technique is referred to as the V/I DSI technique. The

preceding discussion of the V/I DSI technique is considerably simplified and

neglects a number of important issues and features. A more detailed description

of this technique is presented in Ref. 5-6.

i

i

TEST APPARATUS AND PROCEDURES

Test Configuration

A pictorial view of the test facility is shown as Fig. 5-6. The DSI He/Neon

laser is shown in the lower left of this photograph. A single laser beam Is

split within the laser assembly to form two coherent beams that converge to form

the measurement probe volume. In the vlew shown, the Berglund-Liu droplet generator (upper left) is operating to form a monodlsperse droplet stream for calibration of the DSI system. Light scattered by droplets traversing the probe volume

Is collected by a photodetector cell in the DSI receiver (mld-rlght). Electrical

signals are generated from the reflected/refracted light, and the signals are

transmitted to a data processor (not _hoWn) for visibility and intensity evaluatlon. Valid data from the processor are then stored in the computer for

RI/RDBS-312

V-15
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'_ DSI MEASUREMENT

TEST FACILITY

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TABLE

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immediate visual presentation. These data may be transferred to a floppy disc

for subsequent retrieval and presentation in various forms.

Prior to each test, pretest inspection and alignment are conducted on the DSI

system to verify that the system Isoperatlng properly. Measurements are conducted to confirm that the laser beam focus, spacing, and beam intensity are satisfactory. The DSI signal processing equipment settings are appropriately set, and

the DSI data collection and processing computer software is initialized and

appropriate inputs provided. An oscilloscope with a signal storage capability is

set up to monitor the Doppler burst signals from the DSI receiver photomultlpller

tube.

A Berglund-Liu monodlsperse droplet generator is used to calibrate the DSI

measurement system prior to each test. The Berglund-Liu droplet generator is

designed to generate a stream of precise, constant diameter droplets by dlspersIng a constant liquid flow rate at a specified injector vibration frequency. The

injector vibration is produced by a plezoelectrlc crystal powered by appropriate

signal generation equipment. The DSI optics are adjusted to obtain maximum data

rate output and to obtain close agreement (within 5 microns) between the measured

DSI droplet size and the known droplet size produced by the Berglund-Liu droplet

generator. The Berglund-Liu droplet generating system is very sensitive to con

tamlnatlon within the fluid system, irregularities of the pulsing orifice

diaphragm, variations of the fluid upstream pressure, and other unknown factors.

Frequently, calibration of the DSI system was prolonged (occasionally for several

days) by problems with the Berglund Liu droplet generator. Many of the problems

caused by contamination were eliminated by placing a O.5-m_cron filter in the

llquld supply system immediately upstream of the injector. Also, immersion of

the injector orifice diaphragm in a detergent solution whenever the diaphragm was

not in use, reduced calibration delays. At other times, when no detectable cause

for problems with the droplet generator was apparent, the replacement of the

orifice diaphragm with a new diaphragm improved the performance of the droplet

generator.

RI/RD85- 312

V-17
During the calibration procedure and the subsequent injector testing, a

strobe-llght was used to visually observe the monodlsperse stream or the spray

produced by the injector. Scanning of the frequency range of the strobe-llght

allowed the "periodic frequency" of a specific area of the 'injector flow field to

be determined and "frozen" for visual study over a period of time. This capability aided in the determination of the spray characteristics.

DSI callbratlon was considered complete when adjustment of the DSI optics

resulted In (1) a maximum data rate output at the processor, (2) a proper Doppler

signal display on the oscilloscope (i.e., correct number of fringes and symmetrical and nondistorted signal), and (3) the agreement within 5 microns of the predicted Berglund-Liu generated droplet size with the DSI measured droplet size.

After completlon of the DSI calibration, the x-y positioning table shown In

Flg. 5-5 was adjusted by a thread/screw mechanism to locate the centerllne of

the injector at the laser probe volume spatial location. No adjustments were

made to the DSI system after the calibration.

The desired liquid flow rate for the test was then obtained by adjustment of a

small precision value in the injector liquid inlet llne Immedlately upstream of

the injector. The injector liquid effluent was collected in a calibrated container for a preset period of tlme and the valve adjusted to obtain the desired

flow rate. The GN2 flow was determined by measuring the upstream pressure and

temperature of a calibrated sonic venturl in the facility gas supply system. To

ensure sonic flow conditions, pressure measurements were also observed immediately downstream of the venturl. The GN2 pressure was also measured immediately upstream of the injector.

All flow and pressure measurements for each injector test series were documented,

together with other important test information on test data sheets. A typlcal

test data sheet is presented as Flg. 5-7.

RI/RDBS-31 2

V-18
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RI/RD85- 312

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After veriflcation that the target test flow rate and mixture ratio were satisfactory, an exhaust fan was turned on to prevent reclrculatlon of the injector

effluent. The exhaust fan system consists of a hlgh-volume blower with a large

flexible duct c_nnected to the injector spray catch pan (Fig. 5-6). The exhaust

system is effective in eliminating spray reclrculatlon without affecting test

results. After the start of the b}ower exhaust system, DSI measurements of the

injector spray were initiated.

ATOMIZAIlON CHARACTERIZATION TESIING

A_

The intent of these tests was to obtain a detailed mapping of the spray produced

by the nine baseline elements previously described and presented in Table 2-I.

Such mappings consist of droplet-size distribution plots at vari6us points within

m

the spray. These may then be combined to produce droplet-slze distributions

characterizing different portions of the spray and the overall spray. In additlon, various representative droplet sizes characterizing the spray, such as mass

median, Sauter mean, and volume mean, can be computed from these mappings. These

results would then be available for use in the various liquid rocket engine tombustor codes.

_3

Droplet Discrimination by Composition

All nine baseline injectors employ both propellants in each individual element.

Three of these are unlike impinging llquld/llquld elements, and the remaining six

are gas/llquid elements. The DSI technique cannot discriminate between droplets

of different composition and can only correctly size droplets of one composition

(one index of refraction) at a time. Thus, if two different types of droplets

are present, the DSI will only correctly measure those of one composition, and

those of the other composition wlll be incorrectly sized (uniess they happen to

have the same index of refraction).

Furthermore, it will not be possible to distinguish one type of droplet from the

other. The two different drop-slze distributions would be simply added

together. Almost every droplet-size measurement technique has this deficiency,

but it is only of importance to researchers interested in multiple liquid

RI/RD85-312

V-21
atomizers, such as those In liquid rocket engines. With the exception of some

small amount of "hot wax" test data (where water and wax were employed), no data

exist on the droplet sizes of each type of liquid produced by unlike liquid

injectors.

To circumvent this difficulty, tests were planned to investigate the feasibility

of making one of the two injected liquids invisible to the DSI. These tests were

to be accomplished, using dyes to absorb the incident laser light rather than

refract It to the DSI receiver. If only one liquid were so dyed, Its droplets

would be invisible to the DSI. By an alternative dylng of each liquid, the

obtaining of separate droplet size data for each llquld was thought to be possible. However, prior to these tests being conducted, this technlque was determined unworkable. Rocketdyne IR&D studies indicated that large quantities of

mixed-composition droplets would occur. Even immiscible liquids would be

expected to form such mixed-composition droplets. These mixed-composition droplets would be only partially visible to the DSI and would be incorrectly slzed.

As a result, this particular approach was abandoned, and the three ltquid/llquld

impinging elements were deleted from the testing.

Test Flow-Rate Considerations

= __

Another problem area previously discussed was the chamber and DSI limitations

that prevented higher-pressure test!ng. Thus, the open-air test apparatus

described in the preceding section was constructed. This inability to test at

higher pressure resulted in additional test difficulties as described below.

Once the pressure at the downstream end of the injectors (the backpressure) was

llmlted to atmospheric pressure, flowing these injectors at any significant percentage of their design flow rates would result In pressure drops across the gas

orifices of the injectors of at least several thousand kllo Pascals (several

hundred psi). Th_s result would, of course, requlre that these orifices be

choked. If these orifices were operated In such a highly choked condition, the

resulting gas flow would experience a rapid expansion upon exiting the Injector.

Such a flow would be characterized by a gas Jet wlth complex expansion waves, a

much larger diameter than the orifice, and a very high velocity. This result was

RI/RD85-312

V-22
I

i

considered to be unacceptable. As wlll be subsequently discussed, where the test

rationale and conditions are described, the ga s stream diameter and veloclty are

important variables. Thus, all tests were performed wlth unchoked orifices.

To maintain the injectors unchoked, the gas mass flow rates employed must be very

small. Also, since the ratio of gas flow rate (the atomizing force) to liquid

flow rate (the quantity on which this force acts) was considered important, It

was desirable to maintain mixture ratios comparable to those for which the injectors were designed. Thus, the liquid flow rates were also quite small. Thls

generally resulted in liquid injection veloclties Of less than 0.5 m/sec. At

these low liquid flow rates, the spray behaved In a_most peculiar fashion. The

two triplet injectors exhibited drastic changes In the spray pattern that were

not repeatable. In some cases, streams of larger droplets were observed moving

almost horizontally from the impingement point, or at a very large angle from the

main body of the spray. Thls was especially apparent for the smaller triplet

(element 4). Measurements of the sizes of these droplets were not obtained as

this would have required 'a major modification of both the optical equipment (to

measure these much larger droplets) and the test apparatus (to permit measurements so far from the main body of the spray). Furthermore, these strange spray

patterns were not repeatable. At higher liquid flow rates (mixture ratios much

higher than reasonable for testing), these effects disappeared and the spray

appeared morenormal and well behaved. It is believed that these peculiar spray

patterns observed wlth the triplets result from the effects of surface tension

(and perhaps small contaminant particles). The triplets, especially the smaller

triplet, have much smaller liquid orifices than the other gas/liquid injector

elements. At higher liquid injectlon velocities, surface tension forces would be

negligible, but at these low flow rates, surface tension forces may be a significant factor influencing the atomization process. Such effects are, of course,

unrepresentative of actual injector performance. Therefore, after several mappings of the main bodies of these sprays and after many attempts to obtain

repeatable flows with the two triplet injectors, these two elements were deleted

from the testing.

The sprays of the remaining four elements (elements 3, 5, 7, and 9 of lable 2-I,

a pentad, and three coaxial elements) were found to be repeatable and relatively

RI/RD85- 312

V-23
well defined at the desired basellne test condltlons. The sprays of the two

preburners elements (the pentad and coaxial element 5) exhibited a few large

droplets that are probably not typical of realistic sprays. Droplets were

observed to form and grow on the face of the pentad and move about unttl they

became large and contacted one of the gas Jets. They were then blown off and

atomized, producing a brtef disturbance of that gas Jet. This process occurred

at such a low frequency, and had such little apparent effect on the spray, that

_t was not considered to Jeopardize the valld_ty of the droplet-size _ measurements. The preburner coaxial element produced a spray containing a very few

larger droplets (I.e., greater than 100 microns In diameter), which were not

Included In the measured droplet-size dlstrlbutlon. At higher flow rates, these

disappeared.

Thus, the elements tested In this program were the pentad and the three coaxial

elements. Certain preliminary checkout tests, the rationale for the test plan,

and the final test matrix follow.

Measurement Repeatability

Prior to beginning these tests, It was necessary to perform several checkout

tests to ensure the validity and ut_11ty of the data. As a part of the test

procedure previously described, the OSI was aligned and Its droplet slzlng capability confirmed vla measurements of a monodlsperse stream of droplets. Thls was

done prior to each mapping. In addition, a number of measurements and spray

mappings were made to confirm the repeatability of the measurements. Even with

DSI realignment and recallbratlon, the repeated mappings never varied significantly (i.e., dlstrlbutlon peaks and representative droplet slzes never varied by

more than about 5 microns) from the original mapping.

Measurement Plane

Another concern was the effect of the axial location of the measurement plane.

All mapping measurements were made in a plane located 0.235 meter below the

Injector. This distance was chosen for several reasons. First, as previously

discussed, the DSI performs better In dilute sprays. Making measurements at thls

RI/RD85-312

V-24
distance allows the spray to spread and become more dilute. Secondly, close to

the injector, the spray may contain nonspherlcal liquid particles (ligaments and

oscillating, deformed droplets). The DSI technique requires that the droplets be

spherical. At greater distances from the injector, all the ligaments should be

broken and the droplets stabilized to a spheri'cal shape. To confirm that the

spray was stable at the measurement plane, mappings were made at axial locations

of 0.222 and 0.235 meter. The results were essentially identical.

Sr_S_S__Mapplng Orientation, Data Comg_j.latlo___En,__and Dta Presentation

Once the DSI is aligned and calibrated, and spray droplet-slze measurements

begin, vast quantities of data can be obtained in very short times. The DSI

measures the sizes and velocitles of droplets at a particular probe volume locatlon within the spray and at rates as high as several thousand droplets per second. Droplet-size counting rates in these mappings were generally much lower --

on the order of a few hundred counts per second or less. Droplets are counted

for a period of up to 120 seconds, resulting in droplet counts of about tO00 to

lO,O00. Furthermore, the DSI only counts the droplets passing through the probe

volume that lle within a particular velocity range. The velocity range can be

modified, but is rarely large enough to encompass the total range of the droplet

velocities within the spray. "Thus,'It is necessary to make two or three dropletsize measurements at each location and to subsequently add the droplets of

similar sizes from each of these samples together.

The total number of droplets counted at a partlcular locatlon within the spray,

therefore, could exceed 20,000. Furthermore, to obtain accurate and representative droplet-slze data for the entire spray, it is necessary to make measurements

at a large number of locatlo_s. (The determination of the necessary number of

locations is subsequently discussed.) Thus, it can be seen that the quantity of

data obtained is immense, and data handling, compilation, and presentation prove

a challenging task.

The droplet-slze data are stored in a microcomputer that is a part of the DSI

system. This computer has the capability to compile and output data from a

single run. (All the drop sizes and velocities measured at a single point over

RI/RD85-312

V-25
a slngle veloclty range.) Such output consists of the number of droplet counts

for each of 53 size groups or blns encompassing the measurement droplet-size

range. Droplet veloclttes are similarly separated Into bins. Fig. 5-8 and 5-9

are examples of the DSI standard output. In Fig. 5-8, the left plot Is the droplet-size distribution and the right plot, the droplet-velocity distribution. The

abslclssa values for these plots are printed below the plots, together with a

variety of information regarding DS! parameters. Also shown is the total number

of counts comprising these dlstrlbutlons (7271 In thls case) and the time over

which this measurement was made (120.2 seconds). Figure 5-9 presents the plotted

data In tabular form. The number of droplets In each size bin ls defined in the

first two columns and the number of droplets in each velocity bln ls presented In

the last four columns. Bins with less than five counts are not printed.

V

Note that the velocity distribution Is incomplete. While there are apparently no

droplets with higher velocities than were measured here (up to 22.6 m/sec), there

are quite obviously many droplets having velocities below the minimum velocity of

thls measurement (7.25 m/sec). Since It cannot be assumed that slower droplets

would have the same size distribution as these droplets, It would be necessary to

obtain similar data at this spray location over at least the next Iower-veloclty

range. The droplet counts per second of collection time In each of the nonoverlapping velocity ranges can then be added to obtain the complete distribution for

that location. The DSI software can also produce special plots of droplet slze

vs. velocity.

The standard DSI software does not have the capability to add droplet distribution plots. Furthermore, to obtain representative droplet-slze distributions

characterizing selected regions of the spray or the entire spray, the dlstrlbutlon at each location must be suitably weighted by the area of the spray represented by that measurement location. The standard DSI software also does not

have thls capability. Accordingly, an additional data compilation code was constructed to perform these computations. An example of the output of this code Is

presented In Flg. 5-I0, which is a compilatlon of all the droplet-slze data for

the coaxial element (element 7) at the flow condltlons of test 21. Twenty-one

separate droplet-size measurements, taken at various locations within the

measurement plane of this spray, and over different velocity ranges, were

RI/RD85-312

V-26
Z

©

DIAMETER (MICRONS)

PLOT LABELS

MIN DIAMETER:

MID DIAMETER:

MAX DIAMETER:

MIN VELOCITY:

MID VELOCITY:

MAX VELOCITY:

9 MICRONS

54 MICRONS

98 MICRONS

7.25 M/S

14.88 M/S

22.6 M/S

TIME PERIOD: 120.2 SECONDS

VALID VALOCIIY SAMPLES: 7271

VALID SIZE SAMPLES: 7271

NORMAL ACQUISITION

LASER WAVE LENGTH: 0.6238 MICRON

COLLECTION FACTOR: 8

FRINGESPACE: 14.5 MICRONS

BEAM SPACE: 40.8MM

XMIT LENS FOCAL LENGTH: 935MM

VELOCITY (METERS/S)

VP-1001 STATUS

RANGE: 3

FRINGE COUNT: 13

% ERROR: 3

COMPARE: 12/6

HIGH VOLTAGE: 329

THRESHOLD V (MV): 0.48

Figure 5-8. 0SI Output Plots

RI/RD85-312

V-27
R A W D A T A L i S T ! N_C

SERIES: J RUN: 14

COMMENTS: INTRUSIVE DSI TESTS

DIA BIN

M!C COUNT

84 5

73 12

71 11

$g !2

67 ii

66 21

64 12

$2 24

61 22

59 40

57 91

55 55

54 63

52 59

50 81

48 _S

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45 112

44 121

42 12_

_0 138

38 16!

37 161

35 191

33 201

32 199

30 209

28 251

25 330

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21 416

20 497

IS 480

15 579

15 614

13 492

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rVEL IN BIN

MISEC COUNT

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VEL IN BIN

M/SEC COUNT

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9. I 189

9 153

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7.S 207

7.7 206

7.7 196

7.6 208

7,5 IB7

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7,3 226

7,3 13_-

7.2 IS8

Figure 5-9. OSI Tabular Output

RIIRD85-312

V-28

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plotted. In addition, several representative droplet sizes are computed from

these data. These sizes are the mass medium diameter (_), the volume mean diameters, (030), the Sauter mean diameter (D32), and the linear mean diameter

(010). It should also be noted that the droplet-slze range of Flg. 5-10 is

less than that measured by the OSI and presented In Fig. 5-8. The Rocketdyne

data compilation program has been constructed to exclude the data in the five

bins containing the counts of the smallest droplets. Based upon detailed examlnation of the methods by which the V/I OSI technique functions, It has been

determined by Rocketdyne that these data are of highly questionable valldlty.

Hence, they have been deleted.

V

Figure 5-10 contains the primary droplet-size information required by spray combustion modeling codes and provides a complete, overall assessment of the atomlzatlon characteristics of this spray. If more detailed information is desired,

similar plots can be obtalned for selected locations within the spray and for

particular regions of the spray, as will be subsequently shown.

In many instances during these atomization tests, a substantial portion of the

droplets was smaller than could be reliably measured wlth the DSI. In such

cases, only the mass median droplet diameter is presented, as it would be least

affected by the presence of the tiny drops. However, It must noted that In such

cases, the reported mass median drop size Is In error by some unknown amount.

To systematically record and specify the location of the measurements within the

spray, a standard three-dlmenslonal Cartesian coordinate system was employed.

This system is presented In Fig. 5-11. The center of the injector element is at

the x=o, y=o, z=o point, and any location within the spray can be readily identified by it (x,y,z) location. As previously noted, the element is mounted on a

framework resting on an x-y table that can be easily and precisely transiated in

the x or y directions. Thus, during testlng, the coordinate system and spray are

traversed In the x and y directions, and the DSI equipment and the probe volume

remain stationary.

RI/RD85-312

V-30

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Figure 5-11. Spray Mapping Coordinate System

RI/RD85- 312

V-31
Element orientation within this coordinate system was as follows. For coaxlal

elements, the LOX post is centered about the negative z-axis, with the injector

face at the z=o locatlon. Thus, the center of the LOX post at Its downstream end

is (o, o, -R), where R is the LOX post recess of the particular element. The

pentad element was oriented so that a particular palr of outer orifices were

always aligned along the y axis. Pentad orientation is presented in Fig. 5-11.

Effects of Ambient Air Motion on the Spray

One final issue of concern is the effect of the ambient air motion on the characterization of the spray. As discussed in the section of this report describing

the gas/llquld mixing test effort, the spray causes motion and reclrculatlon of

the surrounding air. Considerable evidence exists that demonstrates the motion

of the surroundlng gases can greatly affect the measured droplet size (Ingebo -

Ref. 5-7, ZaJac - Ref. 5-8 and 5-9, and Falk - Ref. 5-I0). To prevent the reclrculatlon of the spray in these tests, a large and deep spray catch pan was

installed slightly below the measurement plane. This catch pan was equipped wlth

a 51-cm-dlameter duct connected to a blower. This pan provided a constant flow

of alr downward into the catch pan, thereby containing and preventing the reclrculatlon of the spray. While such a spray removal device was necessary, some

concern was expressed that the suction produced by the blower would alter the gas

motion sufficiently to change the droplet sizes.

In addition, there was concern that anything disturbing or modifying the manner

in which the surrounding gas was ingested into the spray, might affect droplet

sizes in the spray. If this effect is strong, It has serious consequences

regarding the ability to extrapolate from any cold-flow atomization data to real

injectors operating in engines. The presence of the injector face and the adjacent elements, and the expansion of the fuel gas caused by heating and combustion, will most certainly and drastically alter the motion of the surrounding

gas. These effects cannot be simulated in cold flow.

To partially investigate thls, a series of tests were conducted to determine the

effect of the surrounding air motion on the atomization characteristics of the

RI/RD85-312

V-32

V
_t

water spray. If such an effect was observed, and if the effect was significant,

then the applicability of this or any atomization data to any combustor would be

very dlfficult and/or of doubtful validity.

For the investigation of this effect, two different configurations were tested.

The first set of tests was conducted with the catch pan exhaust fan operating to

promote removal of the spray droplets from the measurement region. The exhaust

fan inhibits reclrculation of the air and water droplets. The data from the

tests wlth the fan operating were compared to tests at the same condition without

the fan operating.

The second set of tests was performed with and without a shield (a cone) around

the injector spray pattern. The function of the shield (cone) was to mlnlmlze or

distort the natu'ral influx of the surrounding air into the spray. The objective

of these tests was to determine if such a change in air movement would significantly change the spray droplet size and/or velocity characteristics. These

tests were also conducted with and without the exhaust fan in operation.

These tests were performed very early in the program (before the installation of

flow-measurement devices). The gas and liquid side injector pressure drops were

measured, and flow rates estimated from these. However, subsequent to these

tests, some leakage was discovered between the gas and liquid sides of the element because of a faulty 0 ring. Thus, the actual flow rate computed in this

manner could be somewhat in error. Since the intent of these tests was to determine the relative effects of air ingestion variations, the spray produced need

only be constant (which it was). These tests employed the large gas/llquld preburner triplet (element 2), flowing at very high liquid flow rates so as to produce a constant and repeatable spray that was free from the unrepresentative

spray distortions effects observed at low liquid flow rates.

The cone used in certain of these tests fit tightly about the injector element

and was approximately 6 cm in diameter at that end. The cone concentrically

enclosed the spray (without touching it) for a distance of 15 cm and was 30 cm in

diameter at the open end. The cone ended 7.9 cm above the measurement plane (z

22.9 cm for these tests).

RI/RD85 312

V-33
Measurements were made at a number of points within the spray, and the

droplet-size distributions were found to be comparable (i.e., within about lO

microns) at all points measured. The droplet sizes measured were generally very

small, and it was apparent that many droplets had diameters less than what could

be reliably measured with the DSI (i.e., about 18 microns). It is believed that

the upstream leakage between orifices was responslble for thls fine atomization.

Since large portions of the spray consisted of drops having diameters smaller

than could be measured, any computed representative drop sizes would be of little

validity. Consequently, these were not computed. Thus, the comparison between

conditions must be based upon a direct comparison of the droplet-size and velocity distributions.

Typical droplet-slze distributions at the center of this spray are presented In

Fig. 5-12 through 5-17. In each figure, the injection pressure drops, and the

fan and cone conditions are noted. Figures 5-12 to 5-14 present centerllne

drop-slze distributions for various Fan and cone conditions. Figure 5-15

presents the droplet-slze distribution near the periphery of the spray, for the

fan-off, cone-off condition (comparable to Flg. 5-12); there appears to be little

difference In these distributions. Figures 5-16 and 5-17 present data at the

same location In a spray wlth a lower liquid flow rate and a higher gas flow

rate. These Flow conditions produce an even finer spray than that of

Figures 5-12 through 5-15.

Corresponding velocity profiles are presented in Flg. 5-1B through 5-22.

Figures 5-18 through 5-20 present the measurements of the velocity distributions

of the droplets In the center of the spray at the higher water flow rate. Each

of these figures contains two or three veloclty-distrlbutlon plots corresponding

to the slze-dlstributlon plots of Fig. 5-12 through 5-14, respectively Each

slze-distribution plot was a combined result from two or three runs (note run

numbers). Each of the veloclty-distrlbutlon plots is normalized by the DSI software so that the maximum number of counts always reaches the top of each plot.

This must be kept In mind when viewing these plots. For example, in Fig. 5-18,

the actual numbers of counts per second at 3.8 m/sec in run lO0 is approximately

equal to the peak in the distribution of run lOl, If these plots had not been

normalized then it would be apparent that the high-velocity end of run lOl fits

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o

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V

V

RIIRDB5-31 2

V-40
i_ 41

j

_ii,ir ̧ ̧ ̧ !ip

m C_

._b GO

r

INJECTOR #2

FAN OFF : CONE OFF

PINJ : H20 = 15 PSI N2 = 2 PSI

LOCATION (0, O, 22.9)

SERIES II RUNS lO0 & lOl

RUN lOl

1.85 3.82

DROPLET VELOCITY

'_I I ,

3.70 7.65 ll.60

METERS/SEC

Figure 5-18. Air Motion Tests, High Liquid Flow Rate Velocity Plots (1)
4=, OO

mO _m

T

mO

INJECTOR #2

FAN ON : CONE OFF

PINJ : H20 : 15 PSI N2 = 2 PSl

LOCATION (0, O, 22.9)

SERIES II RUNS 124 & 125

RUN 125 RUN 124

i

m •

Mq

I

1.85 3.82 3.70 7.65 II.60

DROPLET VELOCITY METERS/SEC

Flgure 5-19. Alr Motlon Tests, High L1quld Flow Rate Velocity Plots (2)
._, O0

I

INJECTOR #2

FAN OFF : CONE ON

PINJ : H20 = 15 PSI N2 = 2 PSI

LOCATION (0, O, 22.9)

SERIES II, RUN 150 & 151

RUN 151 RUN 150

l !

I.85 3.82 3.70 7.65 II.60

DROPLET VELOCITY METERS/SEC

Figure 5-20. Air Motion Tests, High Liquid Flow Rate Velocity Plots (3)
I

INJECTOR #2

FAN OFF : CONE OFF

PINJ : H20 = ]5PSI

LOCATION (2.5, O. 22.9)

SERIES II

RUN 105

N2 = 2 PSI

RUNS I04 & I05

RUN I04

n - mlll

'm

'N

m

e

1.85 3.82 3.70 7.( .60

DROPLET VELOCITY METERS/SEC

Figure 5-21. Air Motion Tests, High Liquid Flow Rate Velocity Plots (4)

( ( (
(I!_

{._ (jmi

P_

RUN 5

1.85

!

3.82

INJECTOR #2

FAN ON : CONE OFF

PINJ : H20 = 5 PSI N 2 = 2 PSI

LOCATION (2.5, O, 22.9)

SERIES HH RUNS 3, 4 & 5

3.70

RUN 4 RUN 3

14.89 22.60

I

Figure 5-22. Air Motion Tests, Low Liquld Flow Rate Velocity Plots (2)
smoothly with the low-velocity end of run I00. Figure 5-18 demonstrates that

nearly all the droplets in the spray at this location (0, O, 22.9) were traveling

at velocities between about 2 and II m/sec, and that the peak in the velocity

distribution (mode) occurs at approximately 5 m/sec. Comparison of Fig. 5-18 and

5-20 indicates that the cone appears to cause the droplets to move at slightly

lower velocities. Comparison of Fig. 5-18 and 5-21 indicates that little

difference exists in the droplet velocities at different locations within the

spray. Figure 5-22 is the velocity distribution plots corresponding to the

slze-distribution plot of Fig. 5-17. These mappings were obtained at a lower

liquid flow rate and higher gas flow rate. These injection conditions resulted

In a large number of droplets moving at higher velocities than in the other

(higher liquid and lower gas flow rates) injection condition. The effect of the

higher gas injection velocity has resulted in an increase in some of the droplets

velocities, as might be expected.

These results are typical of many such measurements performed to assess the

effects of the cone and exhaust fan on the spray. The results indicate that the

drop-slze distribution is not slgn%flcantly affected by the presence of the cone

or the suction of the exhaust fan. The droplet velocities appear to be slightly

reduced by the presence of the cone--a result that seems reasonable. The cone

would impede the ingestion of air into the upper portion of the spray, creating a

lower-pressure reglon near the injector. This causes air to move upward, along

the inner surface of the cone, and toward that lower pressure region. This air

flow would oppose and, hence, retard, the spray motion.

The suction of the exhaust fan had no obvious or significant effect on droplet

velocities. Thus, it was concluded that the droplet size and velocity results of

the planned atomization tests would be unaffected by the suction of the fan (no

experimental apparatus effect). Furthermore, at these relatively low flow rates

at atmospheric pressure, the effect of a major distortion of the natural ingestion of air into the spray was minimal. Only the droplet velocities were

affected, and that effect was small.

Before leaving this subject, it is instructive to consider how much air has been

ingested into the spray. Based upon the injector pressure drops for the high

liquid flow rate condition, assuming a discharge coefficient of 0.8 for these

RI/RD85-312

V-46
orifices and incompressible gas flow, and using the principle of conservation of

axial momentum, the resultant (after impingement) axial velocity of the injected

fluids should be approximately 18.5 m/sec. Since the measured velocities ranged

from only about 2 to 7 m/see (see Fig. 5-18 through 5-20), it was obvious that

significant quantities of surroundlng air were being ingested into the spray and

thereby slowing it. A second, simple conservation of momentum computation,

assuming an average spray velocity of 5 m/sec, indicates that the quantity of

ingested air required is on the order of three times the total mass flow rdte of

the element. This simple computation provides some indication of the potential

importance of the ambient air motion on the spray. While this is still an area

of concern, the testing performed in this program to assess the effects of slight

modifications to this air motion, indicated that such effects were minimal.

Test Plan and RAtionale

_V

The primary intent of these atom%zatlon characterization tests was to develop an

empirical correlation defining the spray droplet sizes in terms of the flow,

geometry, and other governing parameters. Overall spray droplet sizes are commonly characterized by a single representative droplet size, such as the mass

median diameter (D). The choice of the governing parameters varies from the most

basic parameters, such as injector velocity, propellant density, stream diameter,

and impingement angle, to parameters that combine these factors (e.g., mixture

ratio, penetration parameter, and Elverum-Morey parameter) or are directly relatable to engines (e.g., thrust per element). Correlations based upon the most

fundamental parameters can be combined to form the less fundamental parameters.

Also, the more fundamental parameters offer more general applicability and util

Ity. Therefore, the philosophy of this test program was to employ these more

fundamental parameters as the independent test variables. Based upon the extensive study of the atomization literature of Ref. l-l, the following variables are

deemed to be of greatest importance:

Impinging Elements

• Injector velocities, VL and Vg

• Stream diameters (generally assumed to be equal to orifice diameters),

d and d

L g

RI/RD85-312

V-47
Chamber gas density, Pcg

Injector gas denslty, pg

Impingement angle, =

Liquid density, PL

Liquid viscosity, _L

Liquid surface tension, o

Coaxial Elements

Injection velocity, VL and Vg

Inner (liquid) stream diameter, dL

Outer (gas) stream annulus dimension, Y

Inner tube recess, R

Inner tube thickness, t

Chamber gas density, Pcg

. Injected gas density, pg

Liquid density, PL

Liquid viscosity, _L

Llquld surface tension, o

Wh_le other parameters have been considered (e.g., gas viscosity, and stream

turbulence levels and ve]oclty profiles), the listed parameters are those usually

previously studied and/or the variables currently considered to be of greatest

importance. One nonfundamental parameter also considered _mportant _s mixture

ratlo. Mixture ratio Is Important to these alrblast atomizers as It relates the

ratio of the injected liquid (the quantity to be atomized) to the injected gas

(which provides the force of atomlzatlon).

Since there are no atomization theories sufficiently developed or valid for the

types of alrblast atomizers employed In 11quld rocket engines, the approach generally employed Is a straightforward parametr_ c evaluation of the effects of as

many of these Independent variables as possible. Results are then compiled and

an emp_rical equation developed of the form:

or D32 et a] = f(V L, Vg, pg, d L, dg, etc.)

RI/RD85-312

V-48

V
L --

Obviously, wlth so many dependent variables, the size of any test matrix that

would comprehensively evaluate each variable would be immense. Accordingly, the

approach employed in this program is to use a baseline test condition and to

independently evaluate the effects of variations of each of these variables on

this baseline condition. Thus, each parameter would be separately varied from

the baseline cond_tlon, while all other parameters remain constant. While this

approach mlnlmlzes the size of the test matrix, it necessarily suffers from an

inability to identify any synergistic or combined parameter effects. For

example, if the gas density effect on atomization varies wlth liquid viscosity,

thls would not be observed with such an approach.

lhe previously described problems that limited the testing to atmospheric back

pressure, low flow rates, and only one impinging element (the pentad), precludes

large variations of certain of these parameters. Also, for some parameters

(e.g., chamber gas density), no variations are possible. Furthermore, in some

instances, It is difficult to vary one parameter independently. Despite these

difficulties, the approach employed here was, whenever possible, to Independently

vary one parameter at a tlme from a baseline condition. When this approach Is

not feasible, more complex data analysis techniques, such as the regression anal

ysis technique employed in the llquld/llquld mixing study, are required to ascertaln the effects of each independent variable.

The general test matrix employed for these tests is presented In _able 5-I. lhe

four basic gas/llquld elements (element 3, a pentad, and elements 5, 7, and 9,

coaxial elements, all from Table 2 l) were used. lhe first series of tests or

mappings establlshed the atomization characteristics of these elements under

baseline flow conditions. The second series of tests were intended to establish

the effects of injector velocity variations at constant mixture ratio.

lhe next three sets of tests were intended to investigate the effects of l_quld

viscosity, density, and surface tension, respectively. To vary each of these

independently, it was necessary to find liquids having the same properties as

water, except for the one property whose effects were to be studied, lhls proved

to be quite difficult. A study of pure liquid properties quickly established

_w../?

RI/RD85-312

V-49
Table 5-1. Atomlzatlon Test Ratrlx

C_ I

Go

PO

I

OBJECTIVE ELEMENTS MIXTURE RATIO FLOW RATE , FLUIDS NUMBER OF MAPPINGS

BASELINE

FLOW RATE EFFECTS

VISCOSITY EFFECT

DENSITY EFFECT

SURF TENSION EFFECT

COMPARE WAX DATA

MIXTURE RATIO EFFECTS

GAS DENSITY EFFECT

ALL

ALL

I COAXIAL

3 PENTAD

7 COAXIAL

3 PENTAD

7 COAXIAL

3 PENTAD

LIKE DOUBLET

ALL

7 COAXIAL

3 PENTAD

BASELINE

BASELINE

BASELINE a

BASELINE a

BASELINE a

2 PER ELEMENT

BASELINE a

BASELINE

2 PER ELEMENT

BASELINE a

BASELINE a

BASELINE a

b

BASELINE GAS

VARY LIQUID

BASELINE a

H20 AND N2

H20 AND N2

GLYCEROL/H20 AND N2

NaCl/H20 AND N2

C2H3C13 AND N2

H20 OR C2H3CI 3

H20 AND N2

H20 AND CO2

H20 AND He

aTHESE VARIABLES ARE BASED ON VOLUMETRIC RATHER THAN MASS FLOW RATES

bHIGH FLOW RATES COMPARABLE TO DICKERSON WAX TECHNIQUE TESTS (REF. 5-11)

( (
Attention was then given to mixtures of

amm_

that no such pure liquids existed.

liquids, and solutions.

An infinite set of such liquids was available; however, there was often little

properties test data to support the selection. This search was quickly narrowed

by the requirement that the liquids be relatively nonflammable, nontoxic, of low

volatility, relatively clear, relatively inexpensive, and compatible with the

materials in the feed system and injector assemblies.

In the course of this search for materials, a particular problem was discovered

regarding the use of surfactant-type materials to reduce surface tension. Such

materials were originally Judged to be very attractive for the purpose of this

study, as very small amounts appeared quite effective in reducing surface tension. Thus surface tension could be varied without appreciably changing the

density or viscosity of the liquid. However, after additional study, it became

apparent that thls approach was questionable. These materials function by

selectively migrating to the surface of the liquid. Thus, their concentration at

the surface is much greater than In the bulk of the liquid. Surface tension

measurements are based on a determination of the surface tension at the surface

of a relatively large quantity of quiescent liquid. Hence, adequate time exists

'for the surfactant-type materials to migrate to the surface, and the surface area

Is small relative to the quantity of liquid present. However, the atomization

process is characterized by the rapid creation of large amounts of surface area.

Thus, It is necessary to question whether these relatlvely small amounts of surfactant-type materials will have sufficient time to migrate to the surface, and

if sufficient material is present In the liquid to effectively reduce surface

tension on the vast surface area of the spray. Because of thls concern, only

liquids containing relatively large amounts of each of its constituents were

considered for use in thls study.

Chemical handbooks and other references were employed to identify potential

liquid m_xtures and solutions. Several promising candidates were identified and

materials were procured, mixed, and sent to Truesdall Laboratories of Tustin,

California, for properties determination. Some of these materials contained very

small quantities of butyl acetate (a surface acting-type of material as

RI/RD85-31 2

V-51
previously discussed), as this was prior to the concerns about such materials.

Results of these properties determinations were generally in agreement wlth whatever literature data existed on these multlcomponent liquids. The liquids given

In Table 5-2 were chosen for testing. (Their pertinent properties and those of

water and LOX are also presented.)

TABLE 5-2. MULTICOMPONENT LIQUIDS AND THEIR PROPERTIES

LIQUID

16% GLYCEROL AQUEOUS SOLUTION

17.4% SODIUM CHLORIDE AQUEOUS SOLUTION

l, l, 1 - TRICHLORDETHANE

WATER

LOX (@ NBP)

SURFACE TENSION

(dynes/cm)

69

77

22

72

13.2

VISCOSITY

(cP)

1.49

1.38

1.2

0.94

0.196

DENSITY

(kg/m 3)

I041

1135

1316

997

1137

Except for the LOX, all properties are at room temperature. All concentrations

are mass percents.

The glycerol solution had propertles similar to water (within about 5%), except

for a viscosity increase of about 50%. The salt solution was intended to provide

a variation in density; however, the viscosity varlatlon from pure water was

significant. No suitable surface-tenslon variation liquid could be identified,

so TRIC was employed. Thls offered a large variation In surface tension wlth

only about 20% and 30% increases In viscosity and density, respectively. No

suitable, room-temperature liquids could be found with surface tensions and vlscosities comparable to those of LOX. Although they dld not provide precisely the

desired properties variations, these liquids were the best available for the

liquid property effects tests of thls atomization study.

Additional tests were planned to investigate mixture ratio variations and changes

in the density of the injector gas (vla variations In the gas composition).

However, slower than anticipated testing (primarily because of difficulties with

RI/RD85-312

V-52
the droplet generator used to calibrate the DSI) prevented the accomplishment of

the gas density effects tests and most of the mixture ratio effects tests. The

planned test conditions for the tests performed are presented in Table 5-3.

One additional, important pair of tests was planned to provide a comparison with

the large body of "hot wax", llke doublet, 'atomization test data. As discussed

in Ref. l-l, certain evidence and supporting arguments cast doubt upon the valldity of this technique. For the purpose of this comparison, the tests and

results of Dickerson (Ref. 5-11) were employed. Dickerson found excellent correlation of his results with the following equation:

= 7.84Xi04 dL0.57/V L 0.85

where d L is the orifice diameter employed, in inches, VL is the wax injection

velocity in feet per second, and _ is in microns. The intent of these tests was

to match the test conditions of Dlckerson as closely as possible and to compare

these results to his wax results. One of the llke doublets extensively tested by

Dickerson had orifice diameters of 0.066 cm (0.026 inch) and a 60-degree impingement angle. The baseline triplet element fabricated for this program has 0.69 cm

(0.027 inch) outer orifices and a 60-degree impingement angle. By flowing only

the outer orifices of this element, at flow rates within the range tested by

Dickerson, it was possible to closely simulate his tests.

However, one major difference (other than the dropletslze measurement technique)

exists between these tests and those of Dickerson the fluids used. As d_scussed

in Ref. I-I, the extrapolation of atomization test results from one liquid to

another is a major difficulty. Reference I-I presents two liquid property correlations commonly applied, and Dickerson recommends one of these (the WolfeAnderson correlation), for extrapolation of his results to real propellants.

This correlation was employed in this current program to correlate the results of

these tests wlth those of Olckerson. The test conditions for these wax comparison tests are presented in Table 5-4.

RI/RD85-312

V-53
TABLE5-3. ATOMIZATONTESTCONDITIONS- GAS/LIQUIDIN3ECTORS

I

...a

lEST INJECTOR MASS MIXIURE RATIO FLOW RATE INJECTION VELOCITY (M/SEC) &AS

ELEMEN1 (OXIDIZER/FUEL) GAS (KG/MIN) LIQUID CR3/MIN) GAS LIOUID MACH NUMBER

BASELINE TESTS

1 PENTAD 3 0.49 0.123 60.0 242 0.40 0.70

2 COAXIAL 5 0.49 0.113 55.0 259 0.21 0.75

3 COAXIAL 7 3.50 0.116 408 152 0.40 0.44

4 COAXIAL 9 3.00 0.171 5ll 183 0.43 0.53

FLOW RATE EFFECTS TEST (MIXTURE RATIO = CONSTANT)

(Percentage of baseline flow rates in parenthesis after element)

5 PENTAD 3 (120) 0.49 0.147 72.0 291 0.49 0.84

7 COAXIAL 5 (120) 0.49 0.135 66.0 311 0.24 0.90

9 COAXIAL 7 (80) 3.50 0.093 326 122 0.30 0.35

ll COAXIAL g (80) 3.00 0.137 409 146 0.34 0.42

6 PENTAD 3 (130) 0.49 0.159 78,0 315 0.52 O.gl

B COAXIAL 5 (130) 0.49 0.147 71.S 337 0.27 0.98

l0 COAXIAL 7 (130) 3.50 0.151 530 198 0.52 0.57

12 COAXIAL g (130) 3.00 0.222 664 238 0.55 0.69

LIQUID PROPERTIES EFFECTS TESTS (VOLUME FLOW RATE OF LIQUID AND GAS REMAINS THE SAME AS BASELINE.

BECAUSE OF LIQUID DENSIIY CHANGES.)

• VISCOSITY EFFECTS (16% GLYCEROL SOLUTION))

13 PENIAD 3 O.Sl 0.123 60 795 0,40

14 COAXIAL 7 3.64 O.ll& 408 500 0.40

DENSITY EFFECTS (17.4% NaCI SOLUTION)

15 PENTAD 3 0.56 0.123 60 795 0.40

16 COAXIAL 7 3.97 0.I16 408 500 0.40

SURFACE TENSION EFFECTS (l,lol TRICHLOROETHANE)

17 PENTAD 3 0.64 0.123 60 795 0.40

18 COAXIAL 7 4.61 0.116 408 500 0,40

(TESTS Ig AND 20 ARE PRESENTED IN TABLE 5-4)

MIXTURE RATIO EFFECTS TESTS

(Percentage of baseline liquid flow rates In parenthesls after liquid flow rate)

21 COAXIAL 7 2.63 0.116 306 (75%) 152 0.30

22 COAXIAL 7 5.25 O.116 612 (150%) 152 0.59

ITIXTURE RATIO CHANGES

0.7

0.44

0.7

0.44

0.7

0.44

0.44

0.44

( (
TABLE 5-4. LIKE DOUBLET (WAX DATA COMPARISON) TEST CONDITIONS

I

TEST

.w

Ig

20

INJECTOR

ELEMEN1

TRIPLET 4

(OUTER ORIFICES ONLY)

TRIPLET 4

(OUTER ORIFICES ONLY)

LIQUID

TRICHLOROETHANE

FLOW RATE

(LITERS/MIN)

1.46

H20 2.75

INJECTION VELOCIIY

(M/SEC)

34

62

TEST RESULTS AND FINDINGS

The results of these atomization characterization tests are presented in this

section. Twenty-two complete spray mappings are presented. Several mappings of

the triplet elements were also performed prior to the discovery of the flow

irregularities previously discussed. These results have been discarded. The

intended flow rates and mixture ratios for all of these tests (as shown in

Tables 5-3 and 5-4) were very closely obtained in the testing. Spray mappings

were obtained along several rays extending radially outward in the measurement

plane from the (0, O, 23.5) point. Initially measurements were made along two to

four such rays in 0.508 cm (0.2 inch) increments. As will be shown, subsequent

testing established that very little accuracy was lost if measurements were made

along only two rays in l.Ol6 cm (0.4 inch) increments.

Each test consists of many (sometimes in excess of I00) droplet-slze distribution

and corresponding velocity distribution data sets. Various combinations of these

have been used to construct droplet-slze distribution plots for (1) selected

points within the spray, (2) along particular rays, (3) inner and outer regions

of the spray, and (4) the overall spray. In addition, velocity distribution data

and plots of droplet size vs. velocity are available. This vast quantity of data

is too extensive to be included in this final report. Accordingly, the results

of each test are discussed, observations based upon these data are presented, and

specific plots of data are presented either as examples or to support a discussion.

RI/RD85- 312

V-55
The overall spray droplet-slze distribution for test l 'is presented In

Flg. 5-23. Note that some portion of the actual droplet-slze distribution conslsted of droplets having diameters less than could be measured by the DSI. If

these droplets had been included, the measured representative droplet diameters,

especlally the linear mean drop size (Dlo) would be reduced from those shown in

Fig. 5-21. Thls was a common problem that occurred in many of the spray mappings. The mass median diameter would be least affected by these small droplets. Accordingly, it is the primary representative droplet size employed in

these tests.

Figures 5-24 and 5-25 present results obtained from a very extensive mapping of

the preburner coaxlal element (5) at baseline flow conditions (test 2). One hundred separate droplet-slze distribution measurements were made along four different rays (_x axis, -x axis, ÷y axis, and -y axis.). Measurements were made along

these rays to a distance of about 5 cm In 0.508 cm increments.

This series of mappings were then employed to assess the mapping resolution

(i.e., number of rays and measurement spacing) required to obtain accurate

results. Figure 5-24 is a total spray droplet-slze distribution constructed from

the data obtained at all measurement locations. Figure 5-25 Is a corresponding

plot that utilized only the data from every other mapping 1ocatlon. The similarity of these plots demonstrates the adequacy of employing l.Ol6 cm spacing

between measurement points. Data from all pairings of the four rays measured

were used to construct an overall spray droplet-slze distribution. These results

indicated that any two rays would adequately reproduce the more comprehensive,

high-resolutlon results of Fig. 5-24. Similar comparisons of measurement resolution in some of the subsequent tests further supported the findings of thls test

regarding measurement location requirements.

The droplet distributions measured at each location in thls test were surprisingly uniform throughout the spray. However, slightly larger drops appeared to

exist in the central part of the spray. Velocity profiles at two points within

the spray are presented in Fig. 5-26 and 5-27. These clearly demonstrate that

the droplets are generally traveling faster near the center of the spray. Also,

the number of droplets counted (samples) in the center of the spray is much

RI/RD85-31 2

V-56
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RI/RD85-312

V-57
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RI/RD85-312

V-58

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RI/RD85-312

V-59

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RI/RD85-312

V-60
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RI/RD85-312

V-61
greater than near the edge.

measurement time.

A]I data obtalned are based upon a maximum 2-minute

The overall spray droplet-size distribution for test 3 (coaxial element 7 at

baseline flow conditions) Is shown In Fig. 5-28. Also presented are droplet-size

distribution plots for the inner core region (Fig. 5-29), out to a radius of 2

cm, and the outer annular region'(Fig. 5-30), between 2 and 4 cm from the center

of the spray. These results indicate that slightly larger droplets exist In the

outer region. However, the difference Is quite small, and the spray is surprisIngly unlform.

Figure 5-31 presents the overall droplet-size distribution for the coaxial element (element 9) at baseline flow conditions (test 4). As In the prevlous test,

the droplets In the outer region of the spray were sllghtly larger than In the

central core of the spray. Figure 5-32 Is an example of another type of data

comonly obtained In all these tests. This Is a plot of droplet slze vs. velocity obtained at the (0, O, 23.5) locatlon within the spray. Thls plot suggests

only a weak functional relationship between droplet veloclty and diameter, a

finding substantiated by plots of similar data collected at other locations in

the measurement plane. Measurements of this type were obtained at different

measurement locations in many of these tests, wlth very slmllar results. The

larger droplets appear to be moving faster than the smaller droplets, but the

difference Is small. This Implies that the difference between spatial and temporal droplet-slze data (see Ref. 1-1) would be very similar at this locatlon

within the spray.

Tests 5 and 6 were lntended to establish the effect of the Injection velocities

of the flutds on the atomization characteristics of the spray. These tests were

performed at flow rates of 120% and 130% of the baseline conditions. Lower

llquld flow rates were Judged undesirable because of the previously discussed

issues of droplets forming on the injector face and fluctuating, distorted sprays

observed at lower liquid flow rates. Higher flow rates could not be attained

without choking the gas orifices. Both of these tests demonstrated a reduction

In droplet size from the baseline case (test l). Flgures 5-33 and 5-34 present

spray droplet-size distributions, based upon measurements along the positive

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these sprays. Unfortunately, no droplet-size measurements were obtained along

any diagonal rays. However, observations of the pentad spray (using a strobe

light source) indicated no apparent fans or lobes from this pentad. The spray

appeared similar to that of a coaxial element.

Tests 7 and 8 were performed to assess injector velocity effects (at constant

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preburner coaxial element (element 5). Tests were performed at 120% and 130% of

the baseline flow rate. Both test conditions produced slgniflcantly larger droplets than the baseline case. Also, both tests demonstrated a distinct difference

between results obtained along the two rays mapped. In both tests, the

droplet-size d_stributions measured along the +y axis had a mass median diameter

that was 10 microns lower than that of the +x axis. These results, for the 120%

flow rate case, are presented In Fig. 5-35 and 5-36. No such difference In drop

sizes along any of the four rays measured In the basellne flow-rate case were

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flow rates for this element allow some reduction In liquid flow rate without

producing unrealistic sprays. The overall spray droplet-size distributions for

this element, at 80% and 130% of basellne flow levels, are presented In Fig. 5-37

and 5-38. These results, together with the baseline test (test 3) appear to

lndlcate a slight increase In droplet size with flow rate. However, measurements

performed at the (0, O, 23.5) locatlon at 80%, 100%, 130%, and 150% of baseline

showed no obvious trend. In all cases, droplet velocities increased with flow

rate.

Tests ll and 12 consisted of spray mappings of the coaxlai element (element 9) at

80% and 130% of the baseline flow rate. The results of these tests, together

with the baseline test, are inconclusive. Change in the droplet-size distributions among tests 4, 11, and 12 are small and indicate no trends.

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chamber coaxial element (element 7), flowing different liquids. Baseline volumetric flow conditions were employed. C, The nitrogen flow rate was maintained at

the baseline value and the liquid volumetric flow rate was set equal to the baseline volumetric flow rate. Density variations in the liquids employed resulted

in corresponding variations in the mass mixture ratio from the baseline case.

lo investigate the effects of viscosity, a mixture of distilled water and glycerol was prepared. This mixture consisted of 16% glycerol by mass. The overall

spray droplet-slze distributions for the pentad (test 13) and the coaxial element

(test 14) are presented in Fig. 5-39 and 5-40, respectively. Comparing these

results with the corresponding baseline tests (Fig. 5-23 and 5-28, respectively)

demonstrates the following: the effect of viscosity on the pentad element is

large - the more viscous fluid produced a much finer spray. The effect of vlscoslty on the coaxial element spray is smaller and opposite -- the more viscous

fluid produced slightly larger droplets. It is generally believed, and confirmed

by others investigating fluld properties effects on atomization, that the effect

of higher viscosity is to increase droplet diameters. Viscosity is thought to

retard the shearing of the fluid and thereby allow larger fluid particles to

escape from the primary atomization region. The pentad element, operating at

this low flow rate, was already subject to flow abnormalities, as previously discussed. It may be that the increased fluid viscosity further aggravated this

problem and produced this surprising, and possibly unrealistic, result.

lests 15 and 16 utilized the same pentad and coaxial elements as the previous

pair of tests, lhe liquid employed was a salt (pure sodium chloride with no

additives) dissolved in distilled water (17.4% NaCl by mass). This resulted in a

liquid of greater density, but also of significantly greater viscosity, than pure

water. The pentad test (test 15) resulted in droplet sizes comparable to those

of the viscosity effects tests (test 13). Droplets near the center of the spray

were found to be considerably smaller than those near the periphery of the spray,

as shown in Fig. 5-41 and 5-42. This effect is also apparent, though to a lesser

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(test 1) with this pentad element.

The coaxial element density effects test (test 16) produced a spray with an overall droplet-slze distribution very similar to the baseline test (test 3) shown in

Fig. 5-28. Also, as wlth the baseline test, the droplets near the periphery of

the spray were larger than those near the center.

The final pair of tests In these properties effects tests employed pure l, l, l,

--trlchloroethane liquid. This liquid provided significant deviation of density

and viscosity from pure water, but the major difference was the change In surface

tension. The overall spray droplet-slze distribution for the pentad element is

presented in Flg. 5-43. Note the change in the droplet size (abscissa) scale.

This is caused by the substantial index of refraction difference between water

and trlchloroethane, which modifies the DSI droplet measurement range. The index

of refraction changes for the other llqulds used was insignificant. Thls test

again demonstrated larger droplets In the outer region of the spray.

The total spray droplet-slze dlstrlbution for the coaxial element flowlng trlchloroethane and nitrogen at baseline volumetric flow rates Is presented In

Flg. 5-44. Comparison of this case with the baseline (Fig. 5-28) shows how

strongly the computed mass median diameter is affected by a few larger droplets.

Test IB produced a droplet distribution wlth a peak (mode) at about 26 to 30

microns, as compared to the test 3 peak at 20 to 24 microns. Yet, test 18 has a

smaller mass median droplet size. Even though the majority of the measured droplets In test 18 are larger than those of test 3, the mass median diameter In test

18 is smaller than that of test 3. Thls is the result of the presence of a relatively small number of very large droplets in test 3. The value of D Is highly

sensitive to these larger droplets. This reduction In larger droplets may be

caused by the lower surface tension, which Is the cohesive force holding the

droplet together. Again, most of the very largest droplets were found to be in

the outer periphery of the spray.

The hot-wax test comparison results (tests 19 and 20) are discussed shortly.

Tests 21 and 22 were performed with coaxlal element 7. The nitrogen flow rate

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21) and 150% (test 22) of the baseline value. The overall spray droplet-size

distribution _n these tests Is presented In Fig. 5-45 and 5-46. The low liquid

flow-rate test (test 21) exhibited much larger droplets than the baseline case.

Some of these droplets were larger than the largest size that could be measured

wlth the DSI configuration (fringe spacing)utilized. Undoubtedly, the actual

of thts spray would be conslderably larger than that measured, if these addi

tlonal large droplets were Included. The h_gher llqu_d flow-rate test produced a

droplet-size d_str_butlon slm_lar to that observed In the basellne test.

Two additional tests (test 19 and 20) were performed to provide a comparison w_th

the hot-wax test results of Otckerson (Ref. 5-11). As previously described,

these tests used the outer or_flces of the small trlplet (element 4) flowing

water (test 20) and trtchloroethane (test lg). This like-doublet was quite s_m_-

lar to one of Olckerson's elements. The overall spray droplet-size dtstrlbutlons

for these tests are presented as Ftg. 5-47 and 5-48. Flow rates In these tests

were very much higher (1.49 and 2.75 11ter/mtn) than all other tests. This was

possible because there was no concern about choktng a gas orifice. The OSI performance was nomlnal durlng these much hlgher flow-rate (more dense spray) tests.

These tests provided 11quld Injection velocltles within the range tested by

Olckerson, almost exactly matched hls like-doublet geometry, and injected into

the open alr In a manner very similar to that of the hot-wax tests. In test 20,

droplet sizes at several different flow rates were examined. Lower injection

rates resulted in larger droplets, which Is in qualitative agreement wlth

D1ckerson's ftnd_ngs and those oF al_ other l_ke-doublet atomization Jnvestlgators (see Ref. 1-1). Both of these tests produced very homogeneous sprays w_th

little variation in droplet size from the Inner to the outer region. Also, a

dlst_nct "fan pattern" was apparent In the data, wtth the spray being conslderably wlder tn one dlmenslon than the other. (In test 19, the dtmenslons at the

maln body of the spray at the measurement plane were 5.0 by 8.1 cm.)

RI/RO85-312

V-83
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A summary of the actual test conditions and the measured, total spray, mass

median droplet diameters is presented In Table 5-5, As previously mentioned, in

some tests a significant portion of the droplet-slze distribution was below the

size measurement range employed In these tests (and In test 21). In the tests In

which this effect appears to be greatest, the mass median droplet diameters are

marked with an appropriate greater than (>) or less than (<) symbol. The test

variables noted In Table 5-5, together with appropriate geometric variables for

each injector, are the basic and most fundamental variables upon which

droplet-slze correlating equations are based.

V

ANALYSIS OF RESULTS

Analysis and discussion of the hot-wax test comparisons will be presented first.

The Dlckerson correlation for llke doublets Is

D = 7.84XI04 dL 0.57/V L 0.85

The correlation for the effect of liquid properties recommended by Dlckerson Is

the Wolfe-Anderson correlation:

V

_(fluld A) = _aA PB

(fluid8) \°B/ \"8/

I/6

Properties for wax, water, and trlchloroethane are

LIQUID

WAX (shell 270 at 95°C)

WATER

TRICHLOROETHANE

P PL

(dynes/cm) (kg/m 3)

25 764

72 997

22 1316

_L

(cP)

3

0.89

1.2

RI/RD85-312

V-88
i

TABLE 5-5. ATOMIZATION TEST RESULTS

i

co

po

po

Do

c,o

r_

TES1 ELEMENT GAS VELOCIIY LIQUID VELOCIIY LIQUID DENSIIY

(M/SEC) (M/SEC) (kg/m 3)

BASELINE TESTS

1 PENIAD

2 COAXIAL 5

3 COAXIAL ?

4 COAXIAL 9

242 0.39

259 0.21

152 0.40

175 0.43

FLOW-RATE TESTS

5 PENTAD

6 PENIAD

7 COAXIAL 5

8 COAXIAL 5

9 COAXIAL 7

10 COAXIAL 7

ll COAXIAL 9

12 COAXIAL 9

291 0.49

315 0.52

311 0.24

337 0.27

122 0.30

198 0.52

146 0.34

238 0.55

VISCOSITY TESTS

13 PENTAD

14 COAXIAL 7

SURFACE TENSION VISCOSIIY

(dyneslcm) cP

DENSITY TESTS

15 PENIAD

'16 COAXIAL 7

997 72 0.94

SURFACE TENSION TESTS

17 PENTAD

18 COAXIAL 7

233 0.39 TO41 69 1.49

152 0.39 I041 69 1.49

WAX DATA COMPARISON TESTS

19 LIKE DOUBLE1

20 LIKE DOUBLEI

242 0.39 1135 77 1.38

152 0.39 1135 77 1.38

MIXIORE RATIO TESTS

240 0.39 1316 22 1.2

149 0.38 1316 22 1.2

33.6 1316 22 l.?

62.0 997 72 0.94

62

<45

<06

67

50

49

56

59

69

61

74

69

<54

74

<50

<67

<56

62

72

71

21 COAXIAL 7 152 0.30 997 72 0.94 >74

22 COAXIAL 7 156 0.60 997 72 0.94 67
Through the use of these properties, the correlations for wax to water and wax to

trichloroethane are V

DWATE R = 1.08 DWA X

DTRIC = 0.63 DWA X.

Following the employment of Dickerson's droplet-slze correlating equation and the

preceding properties correlation for test Ig and 20 conditions, the predicted

mass median droplet size would be lit and ll8 microns, respectively. As shown In

Flg. 5-47 and 5-48, the measured mass median values were 72 and 71 microns,

respectively. Thus, considerable disagreement lles between these test results

and those of the earlier hot-wax technique test results.

Thls disagreement can be attributed to one or more of four factors:

1. A basic flaw In the hot-wax technique as employed by Dlckerson

2. An incorrect liquid properties correlating equation

3. A basic flaw in the DSI measurement technique as employed in thls program

4. Some unrecognized difference in the element geometry or flow conditions

between these sets of tests

Item 4 seems h%ghly unlikely. As previously discussed, all the geometric and

flow variables considered of importanc e were very nearly the same In both sets of

tests. The three other potential causes of thls difference are discussed In the

following paragraphs.

As discussed in Ref. l-l, a number of reasons and certain experimental evidence

indicate that some flaw may exist in the hot-wax technique. In using the wax

technique, it is necessary to be concerned with such potential error issues as

(1) the wax droplet density change upon freezing and Its relationship to droplet

size, (2) the proper collection and sampling of the wax particles -- particularly

the fines, (3) abrasion of wax particles during the sieving process, (4) dropsize resolution provided by the sieving process, (5) spherical uniformity of the

particles, and (6) agglomeration of freezing particles during and after the

atomization process. Many of these issues have been addressed by investigators

RI/RDB5-312

v-go
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employing this technique; however, in most cases, some questions remain.

One very basic question regarding this technique ls the effect of the droplet

cooling and freezing on the atomization process. When the 367 K liquid wax is

sprayed into the cooler atr (as was done In Oickerson's tests), the wax immediately begins to cool, and the properties of the wax undergo significant change

during this process. Furthermore, the outer portions of the droplets are cooled

and frozen first, thereby limiting additional atomization. These effects would

tend to limit the atomization, thus producing larger droplets. If significant

cooling of the surfaces of the wax droplets occurs prior to the completion of

atomization, then the validity of the wax data is highly suspect.

As discussed in Ref. I-I, the liquid properties correlating equations in common

use are of questionable applicability. Another properties correlating relationship that could be applied with as much Justification as Wolfe-Anderson's relatlonshlp, is that of Ingebo (Ref. 5-12):

I/4 I/4

(fluid A) OA UA PB UB

(fluid B) PA P B

This correlation yields wax to TRIC and wax to water droplet-size correction factors of 0.48 and O.gO, respectively. The use of these properties correction factors with Dickerson's droplet-slze equation results in predicted mass median

droplet sizes of 8g microns (test Ig) and 98 microns (test 20). These sizes are

closer to the measured values of 72 and 71 microns, respectively. Valid and proven properties correlating equations are an important requirement for comparing

data obtained wlth different fluids. This is especially important, since nearly

all cold-flow atomization data must be extrapolated to real propellants (e.g.,

LOX). Unfortunately, such correlating equations are not available.

The final potential cause of the differences between the hot-wax test results and

those of this program involve the validity of the DSI measurements. At the start

of each test, the DSI measuring capability is confirmed via measurement of a

stream of monodisperse droplets of known size. However, this does not guarantee

RI/RD85-312

v-g1
the validity of the measurements of the more distributed and dense sprays of

interest. The lack of a recognized, established, standard spray for confirmation

of the performance of droplet-size measurement instruments, ls a serlous problem

that ls being addressed by an ASTM subcommittee. At present, the only available

confirmation of the DSI consists of the monodlsperse droplet stream measurements

and the repeatablllty, consistency, and "reasonableness" of the data.

V

It ls impossible to determine which of the preceding factors ls responsible for

the observed disagreement between the wax data and the corresponding tests of

this program. Certainly, part of the problem is the lack of valid liquid propertles correlations. These results demonstrate the great need for such correlations, and cast some doubt on the valtdlty of the hot-wax technique test

results. Attention is now directed toward the results of the other atomization

tests.

Examination of Table 5-5 shows very little variation In mass median droplet slze

with any of the parameters tested. The DSI measurements are considered valid

within approximately lO microns, and, In most cases, the spread In the data lles

within these limits. This result Is somewhat surprising as considerable variations were employed for some of the important parameters (e.g., gas injection

velocity). The very low liquid flow rates necessary for these tests are a source

of concern. These low liquid flow rates were responsible for the large and

unrealistic deviations from expected spray appearance for the triplet elements,

and for lesser spray anomalies observed with the other two preburner elements

(the pentad and coaxial 5). The test using the two main chamber coaxial elements

is considered to be the most valid. However, no dlscernable trends In any of the

data are apparent. With so little spread In the data, many more tests would be

required to assess the effects of changes to the governing parameters. Attempts

were made to correlate the data In a variety of ways but they met wlth no success. The regression analysis technique that was so successfully applied to

assess the llquld/llquld triplet mixing test results, was used In this endeavor

wlth no success. There were too little data, too little spread In the data, and

too many variables for such an analysis.

RI/RD85-312

V-92
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However, a number of observations can be made from these data that are of some

importance _to future research in this area. Also, the DSI provides previously

unavailable information regarding the internal structure of the spray. These

findings regarding general spray characteristics are:

I. These sprays were surprisingly homogeneous. Droplet mass median diameter variations between the inner and outer regions of the spray were

generally less than lO microns.

2. In general, at any location within the spray, the average velocity of

the larger droplets was only slightly greater than the average velocity

of the smaller droplets.

3. The droplets in the center of the spray have a larger average velocity

than those near the periphery.

4. A slight but general tendency for the larger droplets to be nearer to

the outer edge of the sprays produced in these tests was shown. The

preburner coaxial element (element 5) may be an exception.

5. The effects of flow rate and liquid properties variations on the

measured droplet sizes were mixed or very small (within the estimated

measurement accuracy of the DSI). For example, higher flow rates

appeared to decrease droplet sizes for the pentad, increase droplet

sizes for the coaxial element (element 5), and have no consistent or

discernible effects on the other two coaxial elements.

These results provide the first detailed mapping of liquid rocket injector elements. Data on droplet sizes as a function of location within the spray, droplet

slze vs. velocity information, and detailed velocity mappings through the spray

are available. This information can be of value to efforts for the development

of spray combustion models. To model spray combustion, the spray must be

modeled. Detailed information of the spray structure, droplet sizes, and droplet

velocities provides the necessary data for comparison with the output of the

spray modeling codes. Such "anchoring" and improvement of the various

RI/RD85 312

V-93
su'bmodels of these codes ls necessary In the development of, and to ensure the

validity of the overall spray_combustlon codes for llquld rocket engines

In much the same way that the results of tests 19 and 20 were used to compare

wlth previous hot-wax test results, these coaxial test results can be used to

evaluate droplet-size correlations obtained by "other Investigators. A 11quld

rocket coaxial Injector would be more generally referred to as a plain-Jet (pneumatlc or alrblast) atomizer. However, a wide variety of plain-Jet alrblast

atomlzers exlsts and, many of thls type are quite different, In form, function,

and operating principle, from those employed In liquid rockets. When comparing

results, It ls necessary to ensure that the Injectors are slmllar.

An excellent summary of alrblast atomlzer research and findings has been prepared

by Lefebvre (Ref. 5-13). In this survey paper, two droplet-size correlating

equations are identified, which offer some potential applicability to llquld

rocket coaxlal Injectors. These equations are that developed by Lorenzetto and

Lefebvre (Ref. 5-14) and that developed by K1m and Marshall (Ref. 5-15). These

equations were developed through empirical correlation of the results of a large

number of spray droplet-size measurements. The atomizers employed In these

studles are presented In Fig. 5-49. Also shown In Fig. 5-49 are the range of the

varlatlons of the parameters tested by both of these investigators and the range

employed for the maln chamber coaxial elements (elements 7 and 9) tested here.

SSME maln Injector conditions are also presented as a reference polnt. Dls the

liquid Jet diameter and A Is the area of the gas Jet. All tests were performed

at atmospherlc backpressure. The correlating equations developed by Lorenzetto

and Lefebvre and K1m and Marshall are

= 5.36 x 10-3

0.41 0.32

_L

Ipg VR210.57 Ag0.36 eL 0.16

0.17

+ 3.44 x I0-3 (_L I WL

\PL _/ Wg

-0.57

VR

RI/RD85-312

V-94

(Klm and Marshall)
L

TABLE 5-6. INJECTION CONDITIONS INVESTIGATED AND SSME MAIN COMBUSTION CHAMBER CONDITIONS

L_ID

1

!

u'l

_0

p-i

t

,,-a

P_

Lorenzetto and Lefebvre Atomizer (Ref. 5-14) Kim and Marshall Atomizer (Ref. 5-15)

INJECTION CONDIIIONS

INVESTIGATOR LIQUID PROPERTIES BACKPRESSURE INJECTION VELOCIIY MIXTURE

SURFACE DENSIIY VISCOSITY GAS LIQUIO RAIIO

TENSION 3 D

(N/m) x TO (kg/m 3) (kg/m/sec) x TO 3 (atm) (m/sec) (m/see) (Liquld/Gas) (m) x 103

Lorenzetto

and Lefebvre 26-76 794-2180 1-76 l 60-180 <3 0.063-I 0.4-1.6

Kim and

Marshall 29.6-31.2 a 782-834 8.7-49.2 a l 75-393 Negligible 0.025-16 1.4-5.6

This Program

(coaxial 7

and 9) 22-77 g97-1316 0.94-1.49 1 122-23,8 0.3-0.6 2.6-5.3 4.6-5.1

SSNE MC 13 b 1137 b 0.2 b 210 322 32 3.3 4.78

a These values appear to be reported incorrectly in Ref. S-13. Values from Ref. 5-1S are shown here.

b LOX properties at normal bo_ling point.

A

(m 2) x TO 6

27-507

5.7-6

10-13

56.7

!!
I (a WL)0"33 l

D32 = 0.95 0.37 0.30

VR PL Pg

+ W__L] 1

wgj

1.7

.7

(Lorenzetto and Lefebvre)

= _

Note that both these equations contain gas density terms, even though all the

tests were performed at atmospheric backpressure. These terms were included to

preserve the dimenslonallty of the equations and, perhaps, with some understandlng of the effect of gas density on atomization from other investigations (generally with different types of atomizers). The reader should also consider the

injectors employed in the development of the correlations. Both these injectors

provide some degree of radial inward motion of the injector gas, which is not

realistic In comparison with liquid rocket coaxial injectors. The Klm and

Marshall atomlzer,has a smaller gas flow annulus than a typical coaxial element,

and no recess of the center tube. The Lorenzetto and Lefebvre atomizer has an

extremely large gas flow annulus and center tube recess, and a very small liquid

tube diameter, as compared to typical coaxial elements. Ftnally, the method by

which the droplets were measured (and the measurement location) must be considered. Lorenzetto and Lefebvre used a light scattering technique (presumably

based on Fraunhofer diffraction), and Klm and Marshall employed the hot-wax technique. Problems and issues related to these measurement techniques are discussed

In Ref. l-1 and 5-I.

Now, a combustion modeler or combustor designer requires droplet-slze information

to assess combustor performance, stability, et al. The proper choice of droplet

size has been shown (e.g., Ref. 5-2) to be of great importance In such computations. Despite a large array of questions regarding the appllcablllty of such

correlations, the analyst generally can use no better correlations than those

Just presented. One recent example of this is a study by Carroll et al (Ref.

5-16), which used the Lorenzetto and Lefebvre correlation to assess SSME performance. As Is apparent from Flg. 5-49, thls assessment requires a drastic extrapolation of the Lefebvre equation beyond the test conditions under which it

RI/RDBS-31 2

V-96

V
L

=

was developed (particularly the backpressure and liquid velocity). Even modest

extrapolation of such correlating relationships Is extremely dangerous, as the

relationships are strictly empirical (i.e., there Is no physical basis for their

form). Those who apply and use results and models employing these correlating

relationships must be aware of the very great lack of accuracy incorporated tnto

such results and models.

While the data obtained in this program can do little to extend these correlatlons to higher pressures, etc., these results can be used to assess the correlating equations at conditions somewhat closer to actual coaxial injector

designs. The coaxlal element (elements 7 and 9) baseline flow conditions are In

general agreement with the test conditions of Lorenzetto and Lefebvre and Klm and

Marshall with the following major exceptions:

, Klm and Marshall tests employed substantially lower surface tension and

higher viscosity liquids.

. Lorenzetto and Lefebvre utilized much lower mixture ratios, far lower

than typical maln chamber combustion conditions.

.

.

.

,

Lorenzetto and Lefebvre used much smaller liquid injector tlp diameters.

Klm and Marshall employed smaller gas annulus widths.

Lorenzetto and Lefebvre used very much larger gas annulus widths.

While both injectors differ from the coaxial injectors of thls program,

the Lorenzetto and Lefebvre design appears most different.

To assess the effects of these differences and compare correlations, the baseline

test conditions and geometries for the coaxial injectors (elements 7 and 9) were

input to both of the correlating equations, wlth the following results (D values

in microns)

F .....

RI/RD85-312

V-97
Coaxial ? Coaxial g

Lorenzetto and Lefebvre equation D*

Kim and Marshall equation

575 443

49 37

Measured D 66 67

*Assumes 5/932 = 1.2 as recommended by Lefebvre (Ref. 5-13)

Even this relatively modest extrapolation of these correlations demonstrates the

great errors introduced -- especially with the Lorenzetto and Lefebvre equation.

While the OSI-measured droplet sizes could conceivably be in error by 20 to 30

microns (no reason exists to suspect this, however), it ts inconceivable that

they could be in error by 400 to 500 microns. If these correlations do not provide reasonable estimates for these modest extrapolations, then their accuracy

must be of even greater concern when they are extrapolated to the much more different conditions of operating llquld engines.

Inspection of the correlating equations indicates one major factor influencing

droplet size in the Lorenzetto and Lefebvre equation that has a much smaller

effect in the Klm and Marshall equation. This is the effect of mixture ratio.

In the Lorenzetto and Lefebvre equation, thls factor increases droplet size by a

factor of lO. Since Lorenzetto and Lefebvre tested at much lower mlxture ratios,

the applicability of the equation at the higher mixture ratios of interest Is

especially doubtful. Additional attempts were made to correlate the coaxial

element test results of this program with the Klm and Marshall correlation, wlth

the following results:

RI/RD85-312

V-98

V
L

Test Measured D Klm-Marshal Equation

2 <45 11

7 56 9

8 59 9

9 69 57

10 61 41

11 74 42

12 69 30

14 74 69

16 <67 50

18 62 69

21 >74 40

22 67 67

With the exception of the preburner element tests (tests 2, 7, and 8), the Kim

and Marshall equation produces results not excessively different from these

tests. However, it continues to predict generally lower values of mass median

droplet size than were measured. The measured droplet sizes are, on the average,

36% larger than the Kim and Marshall computed sizes.

ATOMIZATION TESTING CONCLUSIONS AND RECOMMENDATIONS

The original intent of this study was to develop droplet-slze correlating equations and information that could be employed, with reasonable confidence, to

assess injector designs, and to provide important data to the combustor analysis

computer codes. The previously discussed problems encountered in the course of

this program resulted in the test conditions being limited to low flow rates and

atmospheric backpressure. It was anticipated that such tests would provide sufficient variations in the results to allow their extrapolation to conditions of

greater interest. However, this was not found to be the case. Accordingly,

except for the very general confirmation of the Kim and Marshall correlation, it

was not possible to extract specific droplet-slze correlating equations from

these results.

RI/RD85- 312

V-99
The following are the major conclusions, recommendations, and findings of these

atomization charactertzatlon tests. Supporting rationale for these has been

discussed tn the preceding sectlon.

1. These tests have provided the flrst detal]ed velocity and size distribution mapplngs for several gas/llquld injectors. In addition to the

insight now available regarding the structure of these sprays, these

data can be used to anchor and valldate the spray models being developed

as a part of the efforts to construct liquid rocket engine combustion

codes.

2. These results provlde information that indicates that the hot-wax technique data, and especlally the correlations relating the effects of

liquid properties on droplet size, may be of questionable valldity or

applicability.

3. These tests demonstrate the capability of droplet sizing lnterferometry

to provide extensive and detailed spray lnformatlon that was previously

unavailable, unfortunately, this effort also demonstrates the great

difficulty encountered In attempting to employ the DSI, and especially

the monodisperse droplet generator, under conditions of greatest Interest (e.g., high pressure and dense sprays).

4. The Lorenzetto and Lefebvre droplet-size correlating equation ls not

recommended as a means for estimating droplet sizes for llquld rocket

coaxlal injectors, especially at mixture ratios greater than 1.

5. The Klm and Marshall droplet-size correlating equation Is recommended

for the very rough estimation of droplet sizes for llquld rocket coaxial

injectors operating at high (main chamber) mlxture ratios. This equation Is recommended only because no alternative appears to exist. Based

upon the results of these atomization tests, It Is further recommended

that the droplet sizes computed from the Klm and Marshall equation be

increased by about 30%.

RI/RD85-31 2

V-lO0

V
"5

. It Is strongly recommended that future efforts to acquire atomization

data obtain such data at higher pressures and flow rates. It is universally agreed that the effects of gas density and flow rate are important but there is scarcely any data to assess these effects. Such

hlgh-pressure and high-flow-rate testing would be expected to provide

data that could be used to develop droplet-size correlating equations of

high accuracy and applicability.

f The DSI can satisfactorily obtain data at the flow rates and pressure

tested in thls program. Very much higher pressures and flow rates may

result In sprays too dense to allow such measurements by the DSI.

i

i

,

.

Higher-pressure test facilities must be designed so as to greatly reduce

the natural reclrculatlon of the spray, and must do so In a manner that

least affects the spray.

lO.

These tests have demonstrated that the sprays produced at these conditions are relatively homogeneous In terms of droplet size. Also, since

droplet velocity is relatively independent of size, there is little

difference between concentration and flux based results (spatial versus

temporal data). If this condition continues to be true at other test

conditions, It may be possible to reduce the spatial resolution of the

OSI measurements or to use a different (and stmpler) measurement technique, such as the Fraunhofer dlffractlon method. However, OSI measurements wlll remain a requirement to ensure that these conditions continue

to exist.

The droplet-slze measurement technique employed in this program was the

V/I DSI technique. This second-generatlon DSI technique is inferior to

two new DSI techniques now available. However, all of these DSI technlques, as well as all other nonlntruslve droplet-slze measurement techniques available, have deficiencies that limit their applicability to

the study of liquid rocket injector atomization. These are the inability to measure droplet sizes In the very dense sprays typical of liquid

rocket injectors (especially the very fine and dense sprays of gas/

RI/RD85-312

V-lOl
11quid _nJectors), and the Inability to distinguish droplets by

composition (required to assess unlike liquid injector atomization).

Improved capabilities In these areas are needed.

11. The flndlngs of this study demonstrate the need for considerable addltlonal effort In the study of atomlzatlon. The validity, and especially

the applicability, of all available data and correlations are questionable. The technical challenges and problems are great, but the need for

this Information to support Injector design efforts and the rapldly

developing and very promising field of spray combustion modeling Is also

great.

12. Flnally, It Is strongly recommended that users of any atomization data

and correlatlons become Familiar with their quality, validity, and

applicability. Furthermore, reports of any analyses based upon such

data or correlations should clearly state the limitations and potential

errors assoclated with the use of such atomization data.

RI/RD85-31 2

V-102

V
VI. REFERENCES

l-l Ferrenberg A., and V. Jaqua, "Atomlzatlonand Mixing Study Interim Report,"

NASACR170943, Rocketdyne Report No. RI/RD83-170, July 1980

3-I Rupe, J. H., "The Liquid Phase Mixing of a Pair of Impinging Streams,"

Progress Report No. 20-195, Jet Propulsion Laboratory, California Institute

of Technology, Pasadena, California, August 6, lg53

3-2 Elverum, G. W., and T. F. Morey, "Criteria for OptimumMixture Ratio Distribution Using Several Types of Impinging Stream Injector Elements," MemorandumNo. 30-5, Jet Propulsion Laboratory, California Institute of Technology,

Pasadena, California, February 25, 1959

3-3 Rupe, J. H., "A Correlation Between the Dynamic Properties of a Pair

Impinging Streams and the Uniformity of Mixture Ratio Distribution in the

Resulting Spray," Jet Propulsion Laboratory, California Institute of Technology, Pasadena, California, March 28, 1956

3-4 Hoehn, F.W., J.H. Rupe, and J.G. Sotter, "Liquid Phase Mixing of Bipropellant Doublets," Technical Report 32-1546, Jet Propulsion Laboratory, Callfornia Institute of Technology, Pasadena, California, February 15, 1972

3-5 Fang, J., "Final Report: LOX/Hydrocarbon Fuel Carbon Formation and Mixing

Data Analysis," Contract NAS3-22823, AeroJet Liquid Rocket Company,Sacramento, California, July 20, 1983

3-6 Combined Bimonthly SummaryNo. 40, Jet Propulsion Laboratory, California

Institute of Technology, Pasadena, California, p. 28, December20, 1953 to

February 20, 1954

3-7 Combined Bimonthly SummaryNo. 59, Jet Propulsion Laboratory, California

Institute of Technology, Pasadena, California, pp. 48-50, April l, 1957 to

June I, 1957,

4-I Dussourd, J., and A. Shapiro, "A Deceleration Probe for Measuring Stagnation

Pressure and Velocity of a Particle Laden Gas Stream", Jet Propulsion, January 1958

4-2 Liang, P., and M. Chang, "Development of a Three-Phase Combustion Code for

Modeling Liquid Rocket Engines," 2ist JANNAF Combustion Meeting, October,

1984

RI/RD85-312

VI-l
4-3 Jensen, R. and P. Llang, "Modeling of Dense Sprays from LOX/H 2 Coaxial

Injectors Under SupercrltIcal Conditions," 22nd JANNAF Combustion Meeting,

October 1985.

4-4 Cloutman, L. D., et al, CONCHA - SPRAY, A Computer Code for Reactive Flows

With Fuel Sprays, LA-9294-MS, May 19B2.

4-5 Becker, H. A., et al, Concentration Intermlttancy In Jets, lOth Symposium on

Combustion, pp 1253-1263, 1965.

4-6 Abramovlch, G.N., The Theory of Turbulent Jets, MIT Press, Cambridge, 1963.

5-I Ferrenberg, A., and M. Varma, "Atomization Data Requirements for Rocket Combustor Modeling," 21st 3ANNAF Combustion Meeting, Laurel, Maryland, October

1984.

5-2 Ferrenberg, A., and M. Varma, "Atomization Data for Spray Combustion Modeling," AIAA/SAE/ASME Joint Propulsion Conference, Monterey, California, July

1985

5-3 Ferrenberg, A., "Liquid Rocket Injector Atomization Research," Liquid Particle Slze Measurement Techniques, ASTM STP 84B, J. Tlshkoff, R. Ingebo, and

J. Kennedy, Editors, 1984.

5-4 Bachalo, W.D., "Method for Measuring the Size and Velocity of Spheres by

Dual-Beam Light-Scatter Interferometry," Applied Optics, Vol. Ig, February

1980.

5-5 Bachalo, W.D., C.F. Hess, and C.A. Hartwell, "An Instrument for Spray Droplet Size and Velocity Measurements," ASME ?g-WA/GT-13, 1979.

5-6 Hess, C.F., "A Technique Combining the Visibility of a Doppler Signal wlth

the Peak Intensity of the Pedestal to Measure the Slze and Velocity of Droplets in a Spray," AIAA-84-0203, January 1984.

5-7 Ingebo, R., and H. Foster, "Drop Size Distributions for Crosscurrent Breakup

of Liquid Jets In Airstreams," NACA TN 4087, October 1957

5-8 ZaJac, L., "Droplet Breakup In Accelerating Gas Flow, (Primary Atomlzatlon)," Part l, Rocketdyne Report R-9337-I, NASA-CR-134478, October lg?3

5-9 ZaJac, L., "Droplet Breakup in Accelerating Gas Flow, (Secondary Atomization)," Part 2, Rocketdyne Report R-9337-2, NASA-CR-lS447g, October IgTS

5-I0 Falk, A., "Coaxial Spray Atomization In Accelerating Gas Stream, Final

Report," Rocketdyne Report R-9753, NASA-CR-134B25, June 1975

RI/RDBS-312

VI-2
5-11 Dlckerson, R., "Like and Unlike Impinging Injection Element Droplet Sizes,"

3. Spacecraft, Vol. 6, No. 11, November 1969

5-12 Ingebo, R., "Drop Size Distributions for Impinging-Jet Breakup in Alrstreams

Simulating the Velocity Conditions in Rocket Combustors," NACA TN 4222, 1950

5-13 Lefebvre, A., "Alrblast Atomization," Prog. Energy Combust. Scl., Vol. 6,

1980

5-14 Lorenzetto, G., and A. Lefebvre, "Measurements of Droplet Size on a PlainJet Airblast Atomizer," AIAA Journal, Vol. 15, 1977

5-15 Kim, K., and W. Marshall, "Drop Size Oistrlbutlons From Pneumatic

Atomizers, '°AIChE Journal, Vol. 17, No. 3, May 1971.

5-16 Carroll, R., O. Connell, and C. Limerick, HPerformance Comparison of the

Current SSME and a New Main Injector - Main Combustion Chamber Using the

3ANNAF Methodology," 22nd 3ANNAF Combustion Meeting, Pasadena, California,

October 1985.

RI/RD85-312

VI-3
V
APPENDIX A

This appendix is a reprinting of pages 5 through 106 and 139 through 141 of the

"Atomization and Mixing Study - Interim Report" (Ref. l-l), published as NASA-CR170943, July 1983. The appendix is included in this report in order to provide

the reader with a complete and single report describing the state of the art of

atomization and mixing technology.

It should be noted that the work described in the main body of this report has

demonstrated that certain findings and statements of Appendix A are incorrect or

not generally applicable. In particular, the findings regarding gas/llquld mixing and llquld/llquld triplet mixing design criteria are superseded by the findings reported in Sections III and IV of this report.

RI/RD85-312

A-I
ATOMIZATION

Thls section describes the state of the art In the area of liquid rocket injector

atomization. The need for this information, the parameters of importance, the

manners in which the data are correlated and reported, the droplet size measurement techniques employed, and the specific correlations and data pertaining to

rocket engine injectors (triplets, pentads, coaxial, and some doublet data), are

described, discussed, and assessed. In the study of the atomization literature,

emphasis was placed on experimental programs and empirical results directly

related to liquid rocket injectors. The more theoretical or basic research

efforts are to be studied in a subsequent phase of this program. A bibliography

of the atomization reports reviewed to prepare this assessment is included In the

llst of references at the end of this section.

V

In liquid rocket engines, the combustion process generally is considered to be

evaporation limited, i.e., the evaporation of the propellants is the slowest step

in the combustion process and, therefore, very important to model correctly.

Droplet evaporation rate is a strong function of droplet size and velocity relative to the gas phase. Some computer codes calculate drop velocity and motion.

This is important in properly assessing evaporation, stability, spacecraft contamination by ejected propellant droplets, performance, and wall effects (e.g.,

wall film buildup, heat transfer). Droplet acceleration is due to an imbalance

between droplet inertial forces (a function of drop diameter cubed) and drag

forces (generally a function of drop diameter squared). Thus, the droplet size

is an important parameter in the assessment of droplet evaporation rate and

motion.

It Is possible to write the equations governing the motion and evaporation of a

droplet. The forms of these equations and most of the parameters are known

falriy well over many operating regimes of interest. The equations are ordinary

differential equations that can be solved readily. However, as with all differential equations, any such solution requires knowledge of the initial or boundary

conditions. And this is the problem--these initial conditions are not known well

enough. These conditions are the droplet slze and velocity distributions at the

RI/RD85-31 2

A-2
locations where the droplets are formed. Given these initial conditions, the

governing equations can be solved, and this is precisely what the combustor codes

do. However, errors in the initial conditions produce corresponding errors in

the predictions.

v

This problem has long been recognized and a number of experimental programs have

been performed to establish these initial conditions. Due to the complexity of

the physical processes occurring during atomization, these initial droplet conditions generally are characterized by empirical correlations. Some of these correlatlons and experiments are described later. Both mixing and atomization

experiments often are performed with propellant slmulants. This introduces a set

of corrections that must be employed to extrapolate to the actual propellants of

interest. Another set of corrections generally must be applied to extrapolate

the test data to the operating conditions (pressures, temperatures, etc.) of

interest. Thus, the establishment of these critical initial conditions depends

entirely upon a relatively small quantity of empirical data, relating the effects

of a few of the many parameters affecting these complex physical processes, and

several sets of corrections to this data.

The utilization of such atomization data and correlations in the combustor analysis codes is a major source of difficulty and error. This has been demonstrated

repeatedly in code development programs at Rocketdyne. The three major performance codes in use at Rocketdyne (TPP, CICM, and SDER) all attempt to use such

correlations. In all three cases, it has been found to be necessary to modify

the experimental correlations to force the codes into agreement with large-scale,

rocket engine performance tests. Such "calibration" of computer models with the

actual hardware they attempt to model is a standard procedure when dealing wlth

complex unknown phenomena, although it is obviously a technique of last resort.

Codes that are calibrated in such a manner can be relied upon to produce good

results as long as they are applied to designs and conditions not significantly

different from those that they were calibrated against. However, the accuracy of

the codes becomes increasingly questionable as they are applied to situations and

problems significantly different from the calibration points.

RI/RD85-312

A-3
The poor performance of these drop size correlations indicates that something is

wrong with the measurement techniques, the correlations developed from the measurements, and/or the manner in which they are applied. The assessment of the

state of the art, as described in the remainder of this section, provides reason

to suspect all of these.

V

PARAMETERS AFFECTING ATOMIZATION

The physical processes occurring during atomization cannot be reduced currently

to sets of equations derived from basic physical principles. The most common

case of the break up of a single Jet of liquid has been theoretically studied for

over lO0 years, and these theoretical studies have been unable to predict, to an

adequate degree, the characteristics of the droplets produced. Impinging streams

and other fan-formlng injectors also have been investigated theoretically. These

studies and experimental efforts, combined wlth the strictly emplrlcal investigations of others, provide an indication of the parameters of importance in the

atomization process. However, there is conslderable disagreement regarding the

relative importance of specific variables.

The properties of a liquid propellant that are considered of importance are the

surface tension, viscosity, and density. For a propellant injected as a gas, the

only thermodynamic property generally considered of importance Is the gas density. The geometric varlables of importance for Implnging-type injectors are the

orifice diameter, orifice length, orifice entrance conditions, number of orifices

(triplet or pentad), free Jet distance (i.e., distance from the orifice to the

impingement point), and impingement angle. Flow variables to be considered are

the velocities of the liquid streams and the existence of turbulence in these

streams.

For coaxial injectors, the geometric variables of potential importance are the

propellant flow areas, the inner tube (LOX post) wall thickness, and the recess

of the LOX post. The flow variables of greatest concern are the liquid velocity

and the relative gas to liquid velocity.

RI/RD85-31 2

A-4

V
Another parameter that has been shown to be of very great importance ls the

velocity of the combustion gases relative to the injected fluids. This parameter

affects the aerodynamic breakup of the large droplets and ligaments after they

have separated from the spray fan. This Is often referred to as secondary atomlzatton and many basic research efforts have been performed to evaluate this.

Separation of atomization Into primary and' secondary atomization processes ls

certainly an oversimplification, but it has been employed. The importance of the

combustion gas veloclty ls unfortunate, since the actual veloclty field In the

combustor cannot be determined adequately. The combustion gas velocity field

depends entirely upon the droplet evaporation rate and distribution, which In

turn is highly dependent on lnltlal droplet size, which In lts turn ls greatly

affected by the combustion gas velocity field. Thus, all of these phenomena are

interrelated highly and must be considered together. Even In cold flow tests,

the local gas velocity fleld In unknown. There ls no such thing as "spraying

into still air," as the spray itself transfers momentum to the gas and sets it In

motion.

 !ilj

_7

Another important parameter that is difficult to quantify ls the liquid velocity. Generally, this Is assumed to be the average velocity at the orifice exit

assuming the orifice Is flowing full. But this ts not the velocity of the liquid

in the fan, or of the ligaments, which is probably the velocity of greater

concern.

The effects of combustion on atomization are unknown. Matching the density and

velocity field of the combustion gases in a cold flow atomization test may not be

sufficient. Burning droplets may break up differently (secondary atomization)

than nonburnlng droplets due to the effect of the burning gas envelope about them.

J

Our knolwedge of many of the parameters affecting rocket engine atomization comes

primarily from the study of doublets. Details of many of these studies are

Included later In thls report where the relative importance of the parameters

affecting atomization are discussed In greater detail.

RI/RD85-31 2

A-5
ATOMIZATIONCORRELATINGEQUATIONS

The objective of atomization studies is to develop quantitative relationships

defining the effects of the various governing parameters on the characteristics

of the spray. The characteristics generally most desired are a representative

droplet size (a mean or average value) characterizing the spray, and a droplet

distribution defining how the number of droplets in the spray varies with droplet

size. Other parameters of interest are the distribution of droplets in space,

and the velocities of the droplets.

The data generally most desired is the number of drops of each size in a given

spray. Often the data is obtained in the form of numbers of droplets counted, n,

in each of many uniform size ranges (e.g., 5 to lOg, 15 to 20_, etc.) as

shown in Fig. la. However, this discrete form of data presentation has the

undesirable characteristic that as the width, AD, of the size ranges is varied

the count will change. In order to quantify the data in the form of a continuous

mathematical expression, the data is often converted to the form of Fig. lb.

Here, the number of droplets per unit size range (e.g., per micron) is plotted.

This continuous function is called a distribution function f, and is determined

by evaluating n(D) as AD approaches zero. n(D) is the number of dropAD AO AD

lets having diameters between D - _-- and D ÷ _--. This distribution

function is the mathematical expression that best defines the size distribution

of the droplets. A11 the other forms and techniques for expressing droplet size

dlstrlbutlons can be derived mathematically from this distribution function, f,

where

n__

f = llm AD

AD_ o

Another useful function is the fraction of droplets in the spray at diameter D,

which is

llm n/nt

f/nt = AD_O AD where nt _ total droplet count

RI/RD85-31 2

A-6
D

f = iim

AD_O

f

b. Number of distribution function

I

D D

P

N (D) " "_ n = J fdD

0 O

D

d. Cumulative number distribution

/D

AD

a. Number of dro_z}_e_r size_qroQ.u_r_

S D + 0.5-_D_

n or n J}dD

D,- 0.5_D

D _ 0.5_D

= 1 "_D3/ fdD

v

D - O.5_D

dV- I dN

T _D3 _'_

v

ORIGINAL PAQE M

OF POOR_AUTV

v:l

_ D3n

b

C. Volume of drqpsj)er size rqsoup__

I o D

V (D) = _ v " 1 n I D3 fdD

0 _- o

J

D

e. Cumulative volume distribution

I

R = V/Vto t

1

1.0

D

f. Normalized cumulative volum{

-BIr_TTa-n st

Figure I. Droplet-Slze Distributions and Relationships

RI/RD85 312

A-7
In some appllcatlons,_ ._ ._ the volume of 11quid In the spray as a function of drop

dtameter ts pf _nterest. Multiplying the number of drops in each 'diameter range

by the volume of a drop of that diameter converts the number distribution to a

volume distribution (Fig. lc). It ls also possible to construct a volume dlstrlbutlon function (not shown In Fig. 1) analagous to the size distribution function. Another useful representation Is a cumulative distribution. The cumulative number, N, of droplets at any diameter D Is the sum of the number of all

droplets having diameters less than D (Flg. 3d).

V

D D

N(O) = o_n = 0f f dO or (d_-_) D='_ = f (D)

Similarly, the cumulative volume distribution, V, of the droplets at any diameter

Dls the sum of the volumes of all drops In the spray havlng diameters less than

O (Fig. le).

D D

(11m V

V(O) = _ v = ] &D_o A--D) dO

0 0

or

1

(_-_) O = _ _ D3f

The normallzed cumulative volume distribution, R, (Fig. lf) ls the volume dlstrlbutlon divided by the total volume of all the droplets measured, i.e.,

V

R -

Vtot

The cumulative volume dlstrlbutlon (often normallzed) Is the most commonly utilIzed manner for graphically presentlng the data.

The mathematical expressions defining the drop size distribution that are

encountered most commonly are

dV = AD5 exp (-BD n)

dD

dV = BnOn-1 exp (-BD n)

dO

-1/2 dV = _i _ exp (__2 y2)

dY

where Y = In (O/D)

RI/RO85-31 2

A-8

Nuklyama - Tanasawa

Rosin - Rammler

Log probab111ty

V
and A, B, , and are the adjustable constants. Many other distribution

functions have been utilized and a more complete llst and description of these is

contained in Ref. II. Since:

dV l D3 dN l D3

d--_= _( _-_ = _ f

these distributions actually define the desired distribution function, f. Given

the distribution function, f, the cumulatlve volume (or number) distribution can

be obtained by integrating the distribution function, f, over various size ranges.

More often, the data is plotted in terms of the cumulative volume, or normalized

cumulative volume, versus drop diameter. For some droplet measurement techniques, particularly the frozen wax technique, it is this cumulative volume (or

mass) distribution which is measured directly. Evaluation of the slope of this

distribution then can be performed to define the droplet number distribution or

distribution function. Cumulative distributions tend to "smooth" the data, mask

inaccuracies due to too few droplet measurements, and reduce the apparent differences between different distributions. Usually, only the cumulative volume distribution is reported, so this problem is overlooked often. One very comprehenslvely reported investigation (Ref. 66), which presented all of the data and

dV

plots of V and _-_ versus D, demonstrates this problem. Figure 2 is a cumulative volume plot of the data from one set of droplet size measurements as presented in that report. The data appear to be in good agreement with the integral

of the particular distribution function chosen. However, the plot of the actual

dV

volume distribution, d--D' and the data (Fig. 3) demonstrate that this apparent agreement between the distribution function and the data is misleading.

RI/RDBS-312

A-g
1.000 I0.800

z

Om

Fr,- 0.600

LL

LLI

,_1

O

>

ILl

>

0.400

_J

0.200

0.000

0

0

0

0

I I

200 400 600 800 1000

MAX DIAMETER (MICRONS)

V

Flgure 2. Oroplet Cumulative Volume Distribution and Equation (Ref. 66)

t

RI/RD85-31 2

A-lO

V
0.00500

0.00400

w

>

>

0.00300

w

Z

0

0.00200

O

>

0.00i00

o.00000

®

®

®

(_ ®

_ @

® ®

0 200 400 600 800 1000

MID-DIAMETER (MICRONS)

Figure 3. Derivatives of Droplet Cumulative Volume Distribution

Data and Equation (Ref. 66)

RI/RD85-312

A-ll
Often the drop dlstrlbutlon Is characterized by a single value--a representative

drop diameter. Some such common representative diameters are the mean or average

drop diameter, volume mean diameter, Sauter mean diameter, or mass medlan diameter. The mean or average diameters are defined according to the following

general relationship:

V

Opq =

7. n I DIP

7- n I D1q

l

p-q

where I denotes size range considered.

n I = number of droplets In size range 1

DI = middle dlameter of size range I

Thus, DIO Is the 11near average diameter of all the droplets measured, D30 Is

the dlameter of the droplet having the average volume of all the droplets measured, and 032 ls the diameter of the droplet whose volume to surface area ratlo

ts the average of all the droplets In the spray (referred to as the Sauter mean

diameter). The mass median droplet dlameter Is the droplet whose stze Is such

that one half of the mass (or volume) of the spray Is contained tn droplets

having a larger diameter. On a plot of R versus D (Flg. lf), the mass median

diameter ls the diameter occurring at a value of R = 0.5. All of these representative diameters can be calculated from the distribution function, f.

Most of the correlations developed to define the effect of various geometric and

operational variables on droplet slze deftne this effect tn terms of the Influence these parameters have on one of these representative diameters. These

usually take a form:

Representatlve diameter = A X1 m X2 p X3 q ......

RI/RD85-312

A-12
where the X terms are the variables, or collections of variables, of interest,

and A, m, p, q ..... are the adjustable constants by which the relationship is

made to fit the data. In some instances, sums of terms, each similar in form to

the rlght-hand side of the above equations, are employed. It must be recognized

that such relationships as this do not completely characterize the spray, and

that two sprays with the same mass medlan'or Sauter mean diameter are not the

same. It is often the smallest drop size and/or the largest drop size that are

of greatest importance (e.g., in assessing stability and performance respectively). The mean or median droplet size does not provide this information.

Thus, it is important to also characterize the droplet size dlstrlbutlon--i.e.,

to establish the correlating equation defining the distribution function, f.

DROPLET MEASUREMENT TECHNIQUES

A variety of techniques have been employed to measure droplet sizes. All of these

techniques are subject to inaccuracies and questions associated with the basic

assumptions employed, their manner of use, and/or the quantities of data usually

obtained. Details of these techniques can be found in the literature, including

some of the references contained in the bibliography. The discussion here is

limited to the three primary techniques previously employed to obtain atomization

data for rocket engine injectors. The findings obtained regarding rocket engine

injection atomization utilizing these techniques is discussed subsequently.

Imaging Techn%qyes

These include photography and holography and have been the most extensively

employed methods for droplet sizing, lhey generally require a fairly dilute spray

and offer the advantage of actually "seeing" the droplets as they exist at the

point and time where knowledge of their size is desired. Although multiple exposure techniques can be employed to obtain droplet velocity data, none of the

experimental programs discussed herein did so. As will be discussed shortly,

velocity information is essential to the determination of accurate droplet size

distributions when imaging techniques are employed. Imaging techniques have been

employed to measure droplet sizes in reacting flows. This is an important and

valuable feature that apparently has only been employed for the case of a rocket

engine combustor in the investigations reported in Ref. 26, 72, 73, and 76.

RI/RD85 312

A-13
A major problem with the use of lmaglng techniques has been the need for human

analysis of the images. Although computerized techniques have been developed

recently for analysls of photographic images, al1 of the rocket Injector atomization work described herein employed at least some degree of human involvement in

the analysis of the droplet images. It is necessary for someone to determine

which droplets are to be counted (i.e., which droplets are in focus), and in most

cases, to manualiy measure the droplet sizes. Thls causes errors in two ways--

human Judgment and insufficient droplet counts to define the spray.

Another problem associated with imaging techniques is the time (i.e., cost)

required to manually identify, count, and measure the droplets. This often prevents the counting of a sufficient number of droplets to assure an accurate distrtbutton. A large number of droplets must be counted. The number of small

drops may be over 1000 times as great as the number of large ones, and yet these

large drops are often the most Important to lnclude, in Ref. 5, It Is caiculated

that It ls necessary to count 5500 droplets to be 95% confident that the Sauter

mean diameter ls correct to within 5%. Rarely are so many droplets counted per

sample with imaging techniques.

Perhaps the most important problem associated wlth imaging techniques is that

these techniques only measure the concentration of droplets in a given volume of

space (i.e., spatial distribution) rather than the true droplet distribution, the

temporal distribution. This problem is recognized in the older literature

(Ref. 11), but appears to have been neglected by many others. The nature of thts

problem ls discussed in detail in the following.

In a steady-state flow of droplets, the number of droplets, and the number of

droplets of each slze entering a particular region In space per unit time must be

constant. It ls possible to write a droplet number conservation equation (analagous to a mass conservation equation) as follows:

_1 " Pl A Vl

where N1 = number of droplets of size group I entering a region or control

volume (drops/sec)

Pl = local concentration of droplets of size I (drops/cm 3)

RI/RD85-.312

A-J4

V
i- ....

A = cross sectional area of region perpendicular to the direction of

flow (cm 2)

V i = velocity of the droplets of size group i (assuming all drops in

size group i are travelling at the same velocity) (cm/sec)

Now, the temporal distribution of droplets produced by an injector can be

obtained by counting the droplets per second of each size group i crossing A,

that is, by measuring the values of the Ni terms. As long as the droplets are

moving in one direction (i.e., the spray is not spreading), measurements of the

N i values at any location in the spray will not change. However, the imaging

techniques measure the droplet concentrations, i.e., the Pl terms. As long

as the droplet velocities remain constant as the spray moves downstream, these

PI terms also will remain constant. However, if the droplet velocities

change in such a way that the smaller drops are no longer moving at the same

speed as the larger drops, then the Pi terms also change. The N i values

must remain constant for this is a steady flow situation. Thus, an imaglng technique measures the Pl terms, and the ratios of the Pi terms is not the

ratios of the actual number of droplets of each size group in the spray. The

only time that the imaging techniques produce true dropS>size distributions is

when all the droplets move at the same velocity. This condition rarely, if ever,

occurs in nature or in experiments.

One particularly noteworthy effort that appears to demonstrate this effect is the

work of George (Ref. 72. 73. 76). In these experiments, measurements were made

at several axial locations downstream of the injector utillzlng an imaging technique (holography). In a11 these tests, the gas velocity exceeded the liquid

injection velocity. In such a case, the small droplets would be accelerated more

rapidly than the larger droplets. This would cause the spatial concentrations of

smaller drops to decrease faster than for the larger drops is we move downstream

from the injector face. Thus, we should expect to see more larger drops than

smaller drops in the holograms as we move downstream. This effect was observed

(Ref. 73) and was quite significant. The value of D30 was found to increase by

50 microns or more over a 2-inch change in axial distance. Also, a simple computer simulation of droplet dynamics in a constant velocity gas flow (Ref. 84)

demonstrates significant differences (40% or more variation in representative

droplet sizes) between temporal and spatial measurements.

RI/RD85-312

A-15
Thls problem of spatial versus temporal distributions places some doubt on the

utility of the results obtained with imaging techniques. Not only are the dlstrlbutlons measured not the true distributions, bu_ the dlstrlbutlons will vary

at different locations due to differences in velocities of the various slze droplets. This may account for much of the disagreement between various investigators. The only situation in which the temporal and spatial distributions are the

same is when the velocities of the droplets are not a function of droplet size.

Spraying Into "still" air will never produce this condition; and spraying Into

flowing air only will approximate this condition beyond some unknown distance

downstream from the injector (where the droplets and gas velocities are equal).It

is the temporal distribution that is needed for the combustor models.

Liquid Droplet Capture Technique

r

This technique Involves the capture of a sample of the spray on a solid surface

(e.g., a glass slide) or in another liquid. The droplets captured are measured

under a microscope or photographed for later analysis. Most of the work utllizing this type of technique was performed before 1960 and the technique seems to

have been supplanted, to a large extent, by photographic and other methods. In

many cases the captured droplets are no longer spherical (e.g., droplets captured

on a surface) and corrections to account for this must be applied.

This type of measuring technique requires the use of highly nonvolatlle 11qulds

an_, when the droplets are captured in another liquid (e.g., a heavy oli or glycerine), further requires that the droplet liquid be immiscible in the capturing

liquid. This llmlts the choice of llqulds that can be utilized. Also, the droplets must be captured gently so as to prevent droplet shattering.

In many applications of this technique, it is obviously the temporal distribution

of droplets that Is obtained. For some sampling methods, however, there is some

question as to whether It is the spatial or temporal distributions that are

measured. Such questionable methods include the slide "waving" technique, where

a glass slide is passed rapidly through a spray.

RI/RD85-312

A-16
_j

Like the imaging technique, the liquid droplet capture technique requires conslderable manpower to count and size the droplets. Thus, the slze of the sample

counted may be a serious source of error. Also, this technique requires the

spray to be diluted sufflclently to prevent a significant amount of coalescense

of the droplets in the sample. In order to accommodate this requirement, one

technique often employed is a spray spl_tter: The spray impinges on a plate contalnlng a hole or slit through which only a portion of the spray may pass. Only

this small portion is allowed to fall on the collection surface. This same procedure also is used occasslonally with imaging techniques to dilute the spray.

One aspect of this spray splitting procedure that occasslonally is overlooked is

the effect and probability of droplets colliding with the edge of the splitter

plate. Such collisions can shatter droplets thereby causing the sampled spray to

have a droplet distribution different from that of the main spray. This problem

is analyzed in some detail by Dickerson (Ref. 47).

Droplet Freezing Technique

This technique has been applied extensively in the study of rocket engine injectors. Much of this work that is related directly to rocket engine injectors was

performed at Rocketdyne during the period 1967 through 1975, and utilized wax as

the injected liquid. Fluids other than wax have been used and droplet capture

and freezing in liquid nitrogen also has been performed. All of the work

described herein utilizing this technique was done with wax.

The frozen wax technique offers several advantages over other methods. The

liquid wax is injected into the atmosphere or a large pressure vessel where the

droplets rapidly cool and solidify, and then are collected and sampled. The sam

ple then is subjected to a sieving operation where the wax droplets are separated

into size groups. Each size group then is weighed and a plot of droplet mass

(volume) versus size is constucted. Thus, the cumulative volume, volume distribution, and mass median diameter are measured directly, without the great time

and manpower associated with the sizing and counting of individual droplets.

Also, the true or temporal distribution of droplets is measured, since all the

droplets produced by the spray over a long period of time (several seconds) are

collected. And finally, the number of frozen wax droplets included in the sample

is on the order of millions. This technique does not suffer from a lack of a

sufficient sample size to be statistically accurate.

RI/RD85 312

A-17
One serlous disadvantageous feature of the 'hot-wax technique Is the 11mlted

cholce of materials that can be applled conveniently. Since the properties of

the actual propellants are different from the stmulants, It ls necessary to

establish the effects of these properties (surface tension, denslty, and viscosIty) on the atomization process. Thus, the capability tO perform tests w_th d_fferent flulds having widely varylng properties ls Important. Another feature of

the wax technique that merits consideration Is the density Increase upon freezIng. Because of this, some earller investigators have corrected the wax droplet

sizes by multiplying the measured droplets' volumes by the ratio of the solid to

11quid densities. However, the physlcs of the freezing phenomena Indicates that

the droplets should freeze on the outside first. If thts ts correct, then the

frozen drops should be hollow and no density change drop size correction Is

required. Dlckerson (Ref. 47) has discovered that the droplets indeed are

hollow, and that the volume of the central void ls approxlmately equal to the

size change due to freezing--at least for the larger drops.

One of the most serlous criticisms of most of the hot-wax lnvestlgatlonslnvolves

the problem of deftn_ng the temperature (and hence the properties) of the liquid

wax during atomization. In most Investigations, the hot 11quid wax Is Injected

Into a relatively cold gas (e.g., the atmosphere). For these cases, It Is necessary to question whether the wax has cooled significantly prlor to the completion

of atomization. ZaJac (Ref. 58) presents data showlng that the surface tension

and viscosity of the particular wax utilized (shell 270) Increase by 1_2%and 83%,

respectively, between 93 C (the nominal Injection temperature) and 66 C (slightly

above the wax fusion temperature). Certainly, the surfaces of the wax ligaments

and droplets must be cooler than the bulk wax Injection temperature. Thus, the

wax properties at the injection temperature may not be the same as those existing

during atomization. Longwell (Ref. 1) presents results suggesting the wax technique erroneously may glve large droplets due to viscosity Increases as the

liquid cools during atomization. However, Hasson and Mlzrahi (Ref. 23) present

extenslve data demonstrating that the wax technique produces significantly and

erroneously small droplets (they corrected for an assumed shrinkage of the droplets upon freezing, but this correction ls not great enough to account for the

observed difference). Several investigators (Ref. 29, 70, 71, 78) performed hotwax experiments In which the wax was lnJected Into hot gas and subsequently

RI/RDS5-312

A-18

V
cooled after atomization was complete. In these investigations the potential

problem of wax cooling during atomization should have been eliminated or

minimized.

DROPLET SIZE MEASUREMENT RESULTS AND CORRELATIONS

FOR ROCKET ENGINE INJECTORS

This section presents all the pertinent atomization results that were found

relating to triplet, pentad, and coaxial injectors. Since very little data pertaining to these injector types has been found, and since most of our knowledge

of atomization comes From the study of llke doublets, a discussion of llke doublets also is included.

The expressions relat)ng representative droplet diameter to injector geometry,

operating conditions, and environment vary with each investigator. The most

common representative diameters utilized are the Sauter mean (D32), volume or

mass mean (Djo), and mass median (D). Conversion between these diameters

requires that the droplet size distribution be known, and a generalized converslon requlres that the distribution function be known and Integrable. Generally,

such information is not available, so a direct comparison between these representative diameter equations cannot be accomplished (one exception to this is

described later). However, inspection of the exponents of some of the more

important variables (e.g., liquid velocity, VL, and orifice diameter, dj),

indicates considerable disagreement between these equations. This may be due to

the previously discussed questions and problems regarding the measurement technlques, testing over different ranges and conditions, the use of d_fferent

fluids, unmeasured and/or uncontrolled variables (the most important being the

local gas velocity, Vg), and/or other unknown causes. In a few cases, these

drop size equations contain variables that are not varied significantly during

the testing. In many cases (all cases for the triplet, coaxial, and pentad

injectors), all of the potentially significant variables have not been tested.

The equations developed from such data are incomplete. All of the droplet size

equations described herein are strictly empirical or are based only in part on

very limited theoretical considerations.

RI/RD85 312

A-19
Like Doublet Correlations

The most extensively studled rocket englne type Injector Is the like doublet.

Much can be learned through a revlew of these doublet studles about atomization

processes and measurements In general that Is relevant and important to the study

of triplets, pentads, and even coaxlal Injectors. Table 1 Is a list of selected

impinging doublet Injector atomization studies'showing the droplet size correlations derived and the conditions of the experiments.

One of the earllest studies of impinging element injectors was the work of

Tanasawa, et. al. (Ref. 9). This study employed totally opposed impinging

streams to develop the equation presented In Table 1, and also Indicated that

llquld viscosity ls of llttle importance. Although It Is not clear, It appears

that other properties were varied with viscosity. Subsequent studies indicated

that viscosity effects cannot be neglected.

Prof. Oombrowski of Imperlal College performed a number of investigations (Ref.

32, 33, 34, 35, 40, 45) of atomization of impinging streams utilizing photography. Many of these experiments were performed In a pressurizable vessel with a

weak gas flow to prevent droplet reclrculatlon. Some of these results (Ref. 33)

show the surprising result that increasing chamber gas density first causes a

decrease In droplet size, but at higher chamber pressures (>1 mPa) the effect

of gas density ls to increase droplet size. In order to cover both these

regions, two droplet size equations were fit to the data as shown In Table ].

Reference 33 also presents data showing the effect of chamber gas density on ligament and droplet velocity. Dombrowskl (Ref. 40) also demonstrates the differences in the spray fans and the drop]ets formed by laminar and turbulent

streams. Turbulent stream droplet size correlations did not fit laminar stream

data. Measurements of the liquid velocities in fans were performed and show that

these liquid veloclties can exceed the average stream veioclty if the streams are

flowing laminar. An explanation of thls phenomena ls presented. This difference

between ]amlnar and turbulent streams' drop sizes is an important finding that

was not considered by some later investigators. Oombrowskl's work utilized

photography that measures the spatial distribution or concentration of droplets.

RI/RDB5-312

A-20

V
i

TABLE I. SELECTED L!_

RE FERENCE MEASUREMENT TECHNIQUE CORRELAT ION _-

TANASAWA, ET. AL. DROPLET CAPTURE

(1957) (VARIOUS FLUIDS)

"32:1.73VLOS\°L/

DOMBROWSKI AND PHOTOGRAPHY (WATER, /_\

HOOPER (1962) ETHANOL, AND WATER/ = .0077/o\.16 PL .1

GLYCERINE SOLUTION D32

D32

.0012 < pg < .009 g/cm 3

= . 0468

AP2 kOL/

.oiooo .o25gjcm3 ii: 

z_P IN psi

DROMBROWSKI AND

HOOPER (1964)

PHOTOGRAPHY (WATER) 4

D32 = VL.79 (sin_) TM

INGEBO (1958) PHOTOGRAPHY (n-HEPTANE)

8 INCH FROM INJECTOR D30 =

.3[VLIdj)I/2 + .0125[VL-Vg

D ICKERSON ;

ET. AL. (1968)

HOT WAX SPRAYED INTO AIR d..57

= 8.41 x 105

v_

ZAJAC (1971) HOT WAX SPRAYED INTO AIR

dj'57 Ip_.I- 52 JETS _=_

• _=__

= 3.7 x 105 LAMIN_

c, = 60_

TURBU_:_

b = 2.85 x 106 _kpj/ JETS

: (I.44-.0073e)

t FaUx_ m
:iil _ii[i

ill

i ̧ I,
DOUBLET REPRESENTATIVE DROPLET SIZE CORRELATIONS

ORIGINAL PAGE IS

oF POOR J/UJTY

CONDITIONS

= 180 °

Vg = O

d. = .04 - .I cm

J

VL = 1,0 - 5.0 m/s

o = .029 - .075 Nt/m

PL = .005 - 3.5 POISE

RESULTS INDEPENDENT OF VISCOSITY

= II0 o

d. = .053 cm

J

AP = 25 - 100 psi

PL = 800-1000 kg/cm 3

o = .024 - .073 Nt/m

PL : i - 5 cP

vg--o

NOTED TFLAT RAPID DROP DECELERATION CAUSED DROP CONCENTRATION

TO INCREASE

VL = 7.3 - 19.5 m/s

= 50 - 140 "_

TURBULENT FLOW 9NLY LAMINAR

DATA DID NOT FIT

d. = .074 - .226 cm

J

VL = 9.]5 - 30.5 m/s

Vg = 19.8 - 91.4 m/s

= 90°

_'L: "366 cP

= 680 kg/m 3 PL

o = .020 Nt/m

L_V t VL

D30 = \p-_--/

(INGEBO 1957)

d. = .066 - .206 cm

J

VL = 15.5 - 45.3 m/s

a = 60 ° V = 0

g

WAX PROPERTIES (?)

PL = 3.0 cP

PL = 764 kg/m3

o = .017 Nt/m

ADJACENT FANS (WATER) HAD NO

EFFECT ON _, BUT DID AFFECT

DISTRIBUTION

d. = .152 - .206 cm

J

VL = 9.14 - 67.0 m/s

= 45-90 ° Vg = 0

P RATIO OF CENTERLINE TO

c = AVERAGE DYNAMIC PRESSURE

Pj (A MEASURE OF VELOCITY

PROFILE)

WAX PROPERTIES ORIFICE ENTRANCE CONDITIONS,

MISIMPINGEMENT, JET DISINTEGRATION, TURBULENT VS LAMINAR,

VELOCITY PROFILE, L/d (ORIFICE),

L/D (FREE STREAM), ET. AL.,

INVESTIGATED.

RI/RDB5-31 2

A-21

TABLE 1. (Conclt

REFERENCE MEASUREMENT TECHNIQUE CORRELAT ION

GEORGE (1973) HOLOGRAPHY

N2H 4 DROPS BURNING IN

N204

H20 DROPS IN N2

i" TO 4" FROM INJECTOR

D30 = Ao+AIVL+A21nPL*A3 12o

COLD FLOW

d .

J HOT FLOW

D30 BI + B2dj + B3W L

dj - .15 - .244

VL = 18.6 - 38.7_

N_

= 60 ° _

L_

Vg = 75.6 - 114

7_AJAC (1973) HOT WAX INJECTED INTO

WARM, FLOWING N2

IN A VARIABLE AREA DUCT

FOR-i< -_gm-VL < 1.25

- VL

: T IDc (1-1.52x10-3Dc)

FOR _ > 1.25

VL

WHERE bc = 2.01X105dj'375/VL "75

"r = 1 + 5.8 x 10-6 t

---10.9 DO'66/VL "09

41

L>5cm

ALL D VALUES IN MICRONS, ALL OTHER VALUES IN THE EQUATION SHOULD BE INPUT IN CGS UNITS

rau 

_ded)

CONDITIONS

cm

m/s

_Is

pg = 24, - 36. kg/m3 (HOT)

pg = 1.1 kg/m 3 (COLD)

PL = 1001 - 1017 kg/m3 (HOT)

PL = 103 kg/m3 (COLD)

HOT AND COLD FLOW DID NOT

CORRELATE

DROP SIZE DISTRIBUTION WAS

SPATIALLY UNIFORM

D30 INCREASES WITH AXIAL

DISTANCE

VL = 23.2 - 76.2 m/s

= 60 °

WAX PROPERTIES

I kg/m 3

Pg

Dc = DROP SIZE WHEN Vgm = VL (140 < Dc < 360)

= DROP SIZE WHEN V = 0

o gm

Vgm= MAX GAS VELOCITY (Vg m < 305 m/s)

L = LENGTH OVER WHICH GAS IS ACCELERATED

i

i

z

! RI/RD85-31 2

A-22

FQU 

As previously discussed, this is not the true or temporal distribution, and the

spatial concentration can vary with location and with changes in the droplet

decelerating forces (in this case, changes in gas density). In Ref. 33, it is

noted that rapid drop decelerations at high gas densities caused droplet concentrations to increase. Apparently, the possibility that the amount of increase in

droplet concentration might vary with the size of the droplet was never

considered.

During the late 1950s and early 1960s, a number of investigations of atomization

were performed at NASA-Lewis (Ref. 12, 13, 15, 21, 25, 26, 39, 49). lhese

include studies of the effects of various parameters on_Implnglng stream fans,

single stream crosscurrent injection, and photographic measurements of droplet

sizes produced by impinging streams. One of the most often quoted references of

these is the work of Ingebo (Ref. 15). He performed experiments utilizing llke

doublets injecting heptane into a Iow-veloclty gas stream. The heptane was

injected in the direction of the gas flow (concurrently) and the droplets were

photographed and counted throughout the spray at a distance of 8 inches from the

injector. The heptane streams were flowing _n the turbulent regime. One of the

most unique and important features of this work was the use of the flowing gas

stream to simulate the combustion gas motion in a rocket combustor. Unfortunately, one of the greatest problems then (and now) is that we are unable to adequately define the actual combustion gas velocity field _n a rocket combustor.

And Ingebo's work demonstrates the great importance of this effect. The droplet

size correlating equation developed from this data (Table l) contains a term

(VL - Vg), which accounts for this gas ve]oclty effect on droplet s_ze.

Ingebo also utilized the data to evaluate the constants in a Nuklyama Tanasawa

general droplet size distribution equation to obtain:

dR _3.915_ 6 D 5

exp I-3"915 D/D30)

RI/RD85-312

A-23
Furthermore, In an earlier Investigation (Ref. 13), Ingebo established a relationship for the effect of Injected liquid properties on droplet size as:

.25

D30 (_LLa)

However, this relationship was determined in experiments Involving crosscurrent

Injection of slngle streams lnto flowing gases. Its applicability to cocurrent

Injection may be questionable, yet, it ls often utilized.

In a subsequent study (Ref. 26), Ingebo utilized a moving camera to photograph

burning ethanol droplets and measured their velocities. At a distance of 0.1 m

from the Injector, the drops were observed to be traveling at a higher velocity

than their injection velocity. Most Importantly, the small droplets were

observed to have been accelerated much more than the larger droplets. 35 mlcron

droplets had undergone a velocity increase g times as great as 344 micron droplets. Again, this would indicate that the spatial concentration of each size of

droplets would be rapidly and differently varying with distance from the injector. The effect this would have on the measured spatial droplet slze distribution apparently was not considered.

==- H-=

V

In 1964, several studies (Ref. 36 through 38) were reported by investigators at

AeroJet General. Brown (Ref. 38) captured droplets on glass slides that were

produced by the injection of a stream of liquid into flowlng cold and hot (up to

nearly 1300 K) gas. One of the important features of thls work was the recognition that the spray affects the gas veloclty. An attempt was made to quantify

this effect In a droplet size relationship wlth a term containing the mass flowrate ratio of liquid to gas. In another of these investigations, Wolfe and

Anderson (Ref. 37) performed 'experiments and developed a relationship for the

breakup of large droplets (i.e., secondary atomization) based upon the earller

work of Weiss and Worsham (Ref. 29). This relationship Includes a liquid properties effect of the form,

112 I13 -I16

D30= a gL PL

RI/RD85-312

A-24
In the early 1970s, photography and the new technique of holography were utilized

at AFRPL to measure droplet sizes (Ref. 54, 66, 72, 73, 76). Kuykendal (Ref. 54)

investigated the effects of liquid velocity, orlfice diameter, impingement angle,

stream alignment, orifice length, and surface finish for like doublets flowlng

water. Droplet size equations were developed to define these effects, but the

average drop sizes were based upon a relatively small drop count (occasionally

less than 100), and these equations appear to disagree greatly with most other

similar studies. George (Ref. ?2, 73, 76) utilized holography to measure droplet

slzes In both hot flow (hydrazine drops burning in nitrogen tetroxlde) and cold

flow (water In N2). The form of the droplet size correlations developed in

that effort are presented in Table 1.

During the late 1960s and early 1970s, a very elaborate hot-wax capability was

developed at Rocketdyne and many atomization investigations were performed (Ref.

47, 48, 50, 51, 52, 55, 56, 58, 60, 61, 63, 64, 65, 67-71, 74, 78). Two of the

most comprehensive of these investigations are those of Dickerson, et. al. (Ref.

47) and ZaJac (Ref. 58). Their correlations for llke doublets are presented in

Table I. Both of these efforts were performed by spraying wax into "still" air.

Dickerson's droplet size correlatlons, as reported In Ref. 47, are not in

agreement wlth his subsequent paper (Ref. 52). Discussions with Dickerson

revealed that the earlier liquid velocity data was Incorrect, and the correlations of his latter paper include the correction. Dickerson evaluated the atomization characteristics of, a variety of impinging injectors, with great emphasis

on doublets. Experiments were performed wlth impinging fans from unlike pairs of

llke doublets utilizing water as the other fluid. These tests indicate that

impinging fans tend to broaden the droplet size distribution but have little

effect on D. Droplet slze distributions for several of the injectors tested are

presented In Fig. 4. Note that both axes have been normalized in such a way that

all distributions must pass through the point (I.0, 0.5). Also, as previously

discussed, it is the slope of these cumulative volume distributions that truly

defines the spray. Thus, the apparently small differences in the plots of

Fig. 4.are, in fact, large differences of great importance.

RI/RD85- 312

A-25
f

...i

/

/

/

PENTAD, FUEL SIDE

LIKE DOUBLET, IMPINGING FANS

PENTAD, OXIDIZER SIDE

LIKE DOUBLET, NON-IMPINGING

0 .4 .8 1.2 1.6 2.0 2.4 2.8 3.2

D/D, NORMALIZED DIAMETER

Figure 4. Normalized DropleL Size Distributions

for Selected Injectors (Ref. 52)
V

V

The single most comprehenslve study of atomization of rocket engine-type injectors ls the work of ZaJac (Ref. 58). Zajac examined the effects of ltqutd velocIty, orifice diameter, velocity and diameter ratios, orlflce length, free Jet

length (distance from orlflce to impingement point), angle of impingement, orifice entrance conditions (geometric and flow conditions), mlslmpingement, and

propellant miscibility for like and unllke doublets are well as triplets and pentads. In addition, he measured_transient pressure distributions within the free

streams (a measure of velocity profile and turbulence). ZaJac found that streams

flowing turbulent acted considerably different than laminar streams wlth regard

to atomization (Dlckerson had neglected this, but earller investigators, e.g.,

Dombrowskt (Ref. 40), already had indicated this). Thus, It was necessary to

establish two droplet representative diameter equations, one for turbulent and

one for laminar. Velocity profile also was found to be important, but only In

laminar flow. Free stream breakup prtor to impingement was shown to be important

and can occur at a free stream length of from 5 to 10 orifice diameters In turbulent streams. The much higher gas densities in a real combustor could cause

breakup in shorter lengths.

The state of the art circa 1971 was that the wax technique yielded sufficient

quantity and apparent quality of data to define droplet sizes and slze dlstrlbutlons of hot wax droplets sprayed from like doubltes into still air. Several

problems remained, as follows:

°

.

3.

How valid is the hot-wax technique? Does wax significantly change properties before atomization is complete?

How can hot-wax results be correlated to that of real propellant?

What is the effect of the actual rocket combustor environment (hot,

hlgh. denslty combustion gases moving at high velocity) on the atomization process?

In an attempt to solve some of these problems, tests were,performed wlth comblna

tlons of waxes to examine viscosity effects (Ref. 69) and a large pressure tank

was utilized to simulate hlgh-denslty gases (Ref. 64, 6g, 74, et. al.). In addition, several attempts (Ref. 65, 68, 74, et. al.) were made to validate these

RI/RD85- 312

A-27
droplet size correlations by utilizing them In computer models of rocket engines

and comparing the results of these models wtth the actual hot-ftre tests the

models were attempting to simulate. In one program (Ref. 65), a test engine

operating on wax and liquid oxygen was utilized. Although all of these efforts

reported some degree of success, these three baslc questions still remain essentially unanswered.

One of the most unknown aspects of this problem was (and Is) the effect of the

combustion gas ve'locity on droplet size. The actual velocity field existing In

and around the spray In cold-flow experiments Is never measured. The actual velocity field exlstlng in and around the spray In an operating engine also is

unknown. And finally, the effect of a known flowfleld on the formation and

breakup of a spray fan or stream (primary atomization) Is essentially unknown.

There is, however, a considerable body of work performed to evaluate the effects

of gas flowflelds on the deformation and breakup of Individual droplets (secondary atomization). Such efforts demonstrate the great complexity of this latter

process.

In an effort to establish the effect of gas velocity on the size of droplets produced by Impinging ltquld streams, experiments have been performed in lowpressure wind tunnels. In such experiments, the gas velocity ls defined as the

velocity that existed prior to the introduction of the spray. The effect of the

spray on the gas velocity, although often recognized, ls not taken into account--

very crudely Included by Brown (Ref. 38), and ls not measured. Similarly, the

liquid veloclty in the gas Is assumed to be the average liquid velocity at the

injector orlflce exit, and not the actual llquld velocity in the spray fan.

Thus, In attempting to correlate this very important effect of relative gas velocity (gas velocity relative to liquid velocity), the velocities used are incorrect and are, at bes t , only representative of the true velocities. Despite this,

these experiments do provide an indication of the importance of the relative gas

velocity. Probably the most extensive of these efforts for impinging like doublets are the work of Ingebo (Ref. 15) and George (Ref. 76), as previously discussed, and the latter investigations of ZaJac (Ref. 70 and 71).

_J

RI/RD85-312

A-28
z_

ZaJac utilized a doublet injecting hot wax cocurrently into a ducted hot

( 60 C) nitrogen stream. He separated the atomization process into two parts

(i.e., primary and secondary) and studied these separately. In the primary atomization study, the effect of constant velocity and accelerating gas streams on

the sizes-of droplets initially formed was investigated. In the secondary atomization study, known droplet size distribution sprays were subjected to accelerations to observe droplet breaking. The rate and degree of acceleration was controlled by varying the length and area of the duct downstream of the injection

1ocatlon.

ZaJac found that many of the parameters investigated in his previous work were of

little importance compared to the effect of relative gas veloclty. Much of his

data was plotted in the form of Fig. 5, showing droplet size versus a nondlmenslonallzed relative gas velocity. Note that all of the investigations in which

the liquid was injected into "still" air would be plotted at the -l value of the

nondlmenslonallzed velocity. Shown in Fig. 5 are volume mean diameter data from

Ingebo showing the effect of gas velocity on droplets produced by two different

injectors, mass median diameter data from Zajac, and the calculated droplet size

based upon tests with V = O. The data from ZaJac presented here was obtained

g

with a constant gas velocity (i.e., duct area remained constant). Figure 5

demonstrates the great efect of gas velocity on droplet size.

Based upon his experiments with accelerated and constant velocity flows, ZaJac

constructed the droplet size correlation equations shown in Table I. These

equations compute the mass median droplet s_ze based upon the gas and liquid vel

ocities, the droplet size, D occurring when the maximum gas velocity equals

c

the injected liquid velocity (i.e., Vg-VL/V L = 0), and a parameter , which

includes the distance over which the gas is accelerated (at V = constant, L =

_ c g

infinity). The parameter Dc is computed from the liquid velocity and orifice

diameter. A study of the de_Ivatlon of these equations indicates they are

applicable to turbulent flow only.

The correlation of ZaJac is, to some extent, supported by the earlier work of

Ingebo, and the very important effect of relative gas velocity is demonstrated.

Unfortunately, the application of such results to real combustors is difficult

since the combustion gas velocity cannot be defined adequately.

RI/RD85-312

A-29
i

.,=.=

¢3

30O

200

lOO

[]

I

-1

®

A

®

®

®

Z_ D30 , DJ = .226 CM, V L = 19.8 M/S (INGEBO)15C7H14

D30 , DJ = .074 CM, V L = 19.8 M/S (INGEBO)15 C7H14

D, Dj = .239 CM, V L = 43.3 M/S (ZAJAC) 70 WAX

[] SAME AS ®, COMPUTED FROM ZAJAC 58, V G = 0, WAX

0

/k

<>

0

I I I I I

0 1 2 3 4

Vg - V L

VL

Flgure 5. Effect of Gas Veloctty on Droplet Size
Since about 1975, there has been very little atomization work directly relateable

to rocket engine llke doublet injectors. Th_s is certainly not because the problem has been considered solved. Despite all the earlier efforts to define the

initial droplet sizes produced by llke doublets in combustors, our knowledge in

this area is very crude and/or qualitative. All of the droplet size data to date

is of questionable accuracy and/or validity due to real or possible droplet size

measurement technique problems as previously discussed. The droplet size correlations and distributions developed from this data are generally, strictly empirical. They are mere curve fits of the test data and, as such, may be neglecting

important untested variables and are certainly not of the proper form. These

correlations are based upon data that demonstrated poor or usually unknown

repeatability, considerable spread, and often a relatively low quantity of droplet counts. To some extent, these features of the data aremasked by the extensive us_ of semilogarlthmlc plots of the data and cumulative droplet size distribution plots.

!iii, J

©

Perhaps the greatest problems involve the application or utility of the atomization data. Extrapolation of the cold-flow data using wax or other liquids to the

actual propellants and to the conditions existing in a rocket combustor requires

many questionable assumptions and estimates. One of the most important and,

unfortunately, most questionable of these extrapolations involves the combustion

gas velocity, as previously discussed. Also, since the correlations developed

are empirical, extrapolation to any conditions outside the ranges tested is dangerous. And finally, the attempts to utilize the correlations In rocket combustor codes have not been successful. All of the major rocket combustor codes in

use at Rocketdyne (i.e., 1PP, SDER, CICM) have arbitrary multlpllers of the

initial droplet sizes, either as a part of the code or as an input, in order to

force agreement between the codes and hot-fire engine test data.

Pr___ertles Correlations for Like Doublets

In addition to all the llke doublet "lessons learned" discussed above that are

applicable in general to rocket englne-type injectors, these studies provide the

only known corrections or correlations by which we may relate real propellant

atomlzatlon to that of the slmulants used in atomization experiments. Although

RI/RD85- 312

A-31
many of the droplet size correlating equations contaln liquid properties effects,

probably the two properties correlations quoted most generally are

V

and

D30 (p_)

.25

Ingebo (Ref. 13)

.333 .5 -.167

D30 PL _ PL Wolfe and Anderson (Ref. 37)

Ingebo's correlation comes from a droplet size correlation equation defining the

droplet sizes produced by the breakup of a single stream injected transversely

into an airstream. Wolfe's and Anderson's correlation is based on the breakup of

already formed droplets In gas streams (i.e., secondary atomization). The

applicability of either of these relationships to like doublets can be questioned. In addition, no attempts to establish the effect of liquid properties on

droplet slze distributions were found In the literature. Also, properties correlations for unlike doublets, triplets, pentads, or coaxial injectors, or any

gas/llquld injector apparently do not exist.

Another aspect of the liquid properties correlations problem that often Is overlooked is the actual values of the properties of the real propellants and the

slmulants at their injection conditions. Since liquids are generally, relatively

incompressible, since viscosity usually is not considered to be a function of

pressure, and slnce density, viscosity and surface tenslon data for many propellants and test fluids Is readily available only at room temperature and one

atmosphere or at the liquid's normal boiling point (for cryogenics), these room

temperature and one atmosphere or NBP properties data often are utilized. Thls

can cause considerable error. Liquid oxygen is a propellant of considerable

interest which serves as a good example. For LOX at 134 K (the SSME preburner

LOX injection temperature) the density increases by ll% and the viscosity

increases by 52% between 17 and 340 arm (data from NBS Table TN 384). LDX

properties are, of course, a fairly strong function of temperature, and choosing

the wrong temperature (e.g., using NBP data) also can cause great errors.

RI/RD85-312

A-32
Surface tension Is a particularly difficult property for which to find nonroom

temperature and one atmosphere or non-NBP data. Surface tension is a strong

function of temperature and techniques are available to compute ±he effect of

temperature. For LOX, the surface tension changes from 13.2 dynes/cm at Its NBP

of 90 K to 6.4 dynes/cm at its SSME injection temperature of 134 K. As part of

an attempt to determine the effect of pressure on surface tension, papers were

found that indicated a very strong effect (e.g., O.K. Rice, "The Effect of Pressure on Surface Tension,: 3ournal of Chem. Physics, Volume 15, #5, May 1947).

However, based upon discussions wlth Prof. A. Adamson and Dr. R. Massoudl of the

University of Southern California's Chemistry Department, this effect apparently

Is not due to pressure, but rather to the absorption of gases Into the liquid.

The effect of pressure alone on surface tension should be on the order of a I%

increase per lO0 atm pressure. This absorption of gases also probably would have

a great effect on other properties. Since the time available for absorption,

i.e., the tlme between injection and atomization Is so short, very little absorption would be expected. If thls is the case, the effect of pressure on surface

tension should be of little concern.

Triplet Correlations

Very little data was found regarding the atomization characteristics for triplets. This data Is synopsized in Table 2. All these investigations were performed at Rocketdyne utilizing the hot-wax technique. In all these tests, the

wax was injected into "still" air at ambient pressure.

As a small part of ZaJac's earlier investigation (Ref. 58), a particular liquid/

liquid triplet having all three holes the same size was subjected to atomization

testing. In order to separately evaluate the droplet size produced by the inner

and outer streams, wax and hot water were employed. The wax was injected through

the inner orifice and the water through the outer orifice, and the liquids then

were reversed on a subsequent test. The only variables investigated were the

liquids' velocities, and these were varied in such a way as to maintain a constant mixture ratio. Most of these tests were performed under laminar flow conditions. At hlgh velocity (turbulent flow), the data begins to markedly deviate

from the correlating equation presented In Table 2.

RI/RD85-312

A-33

REFERENCE MEASUREMENT TECHNIQUE

ZAJAC (1971) WAX AND WATER

McHALE AND

NURICK (1974)

WAX AND WATER

CORRELATION

Dc & b° = 3.03 x 104 VL

TABLE 2.

MEHEGAN, ET. AL.

(1970)

WAX AND HOT GAS (N2)

CENTER ORIFICE GAS

-.575

DROPLET SIZE CORRI ..

L

VL (cm/s) IS VELOCITY OF LIQUID IN

ORIFICE OF INTEREST

(CENTER OR OUTER)

PLOTS NEW DATA AND ZAJAC's VS We (L/d)

PLVL2d

W -

e a

L = ORIFICE LENGTH

2 DATA POINTS

dc = do

dc OR do

d

__0_o: .52

dc

L =ORI

pg 1.1 ....

d =1.4

C

d = .52

0 _.

pg: i.I

M

SUBSCRIPT:

C - CENTER

0 - OUTER

_LATIONS FOR TRIPLETS

1

ORIGINAL PAGE IS

of Poor 0uALtry

CONDITIONS!

f

WAX AND WATER PROPERTIES

- 45.7 m/s MR _ CONSTANT

.17 cm

kg/m 3

(pV2d) c

(pV2d)o

.17 cm

].0

PcVc2dc

PoVo22do

= .34 - .38

FICE LENGTH = 7-50d

kg/m3 (Pc = ATMOSPHERIC)

POSSIBLE DEVIATION AT HIGH

VELOCITY

WATER TURBULENT

WAX TURBULENT ABOVE _23 m/s

LAMINAR ONLY

FOR TRIPLET GREATER THAN

FOR UNLIKE DOUBLET AT SAME

VL

cm

cm

kg/m 3

PcVc 2 = 7.58 x 104 Pa (11 psi)

X

-P = 5.39 AND 9.74 (DEFINED SAME AS FOR

dc PENTADS

RI/RD85-312

A-34

l

I

McHale and Nurlck (Ref. 74) performed the most extensive study of triplets to

date, utilizing the wax and hot-water technique previously used by ZaJac.

Although they did not fit the data wlth an equation, they did plot the data versus Weber number times the orifice length to diameter ratio (Fig. 6). All of

their work was done with laminar flow and the Iow-veloclty triplet data of

ZaJac's is included in Fig. 6. The fact that the droplet size was a strong

function of the orifice length to diameter ratio is in agreement with the work of

ZaJac on doublets. Since the flow is laminar, the velocity profile significantly

can affect atomization, and the length to diameter ratio determines how close to

fully developed the velocity profile has become. This effect is included in

ZaJac's correlation for doublets in the Pc/P_J term (Table l).

The data of Fig. 6 confirms ZaJac's earlier work regarding the lack of difference in drop size when by the central and outer streams are of the same size.

Note that the ordinate is normalized by the orifice diameter. Also, Fig. 6 is a

logarithmic plot. As previously mentioned, such plots tend to minimize slgn_flcant differences and scatter in the d_ata. At low Weber numbers, which ZaJac did

not investigate, the normalized mass median drop size for the center and outer

streams was no longer the same.

_z_  ̧

Figure 6 contains all the known droplet size data for llquld/llquld triplets

(except for a few_h%gher velocity tests of Zajac). All of the data in Fig. 6 is

from laminar flowing orifices, while most future triplet designs would be flowing

turbulent. All of the data was obtained with wax and hot water, so there is no

way to evaluate liquid properties effects and, hence, to extrapolate this data to

actual propellants. And finally, none of this data includes the very important

effects of combustor gas velocity.

The only data available regarding gas/llquld impinging elements was obtained by

Mehegan, et. al. (Ref. 55). Two tests were performed with triplets utilizing two

wax streams impinging on a central hot, gaseous nitrogen stream. These test conditions also are presented in Table 2.

RI/RD85-31 2

A-35
_O

¢._ O0

I

..J

t_

5.0

1.0

E3

0.5

0.1

FROM_

ZAJAC

1971

SYM BOL

o

[]

,fzX

©

OUTER

dl

iCM)

.051

.076

.17

.076

.076

INNER

d2

(CM)

.051

.076

.17

.109

.147

I

dl/d2

1.00

1.00

1.00

0.689

0.522

(L/d)1

13,5

13.5

50.0

13.5

13.5

(L/d)2

15.0

15.0

50.0

10.5

7.8

Rewax< 10 4

pVl 2 2d I =,0 2 V2 2 d 2

P • I l I Ii

• _ OUTER ORIFICES (DASHED)

(SOLID)

I

5X10 4 105 5X10 5 106 5XLO6

(WEBER NUMBER)0/D )

Figure 6. Correlation of Relative Droplet Diameter wlth Weber

Number and Orifice L/d for Triplets (Ref. 74)

( ( (
Pentad Correlations

The state of the art for pentad atomization knowledge is essentially the same as

for triplets. What little data is available was obtained from hocketdyne wax

tests. All of these tests were performed by injection of the propellant slmulants Into "still" alr. The data Is synopsized in Table 3.

As a part of Dickerson's (Ref. 47) investigation of injector atomization characteristics, a number of tests were performed on a set of pentad injectors. Droplet size correlating equations were developed relating the mass median drop size

to the orifice diameters and injection velocities. Separate equations were

obtained for the inner and outer orifices. Wax and hot water were used as the

test liquids in a manner similar to the previously discussed triplet tests. In

addition, droplet slze distribution data were obtained. Normalized volume distribution plots from this work were presented earlier (Fig. 4), and show the

different distributions obtained for the center and outer orifices. In addition,

the droplet size distribution equations for this data are presented in Ref. 52.

As previously discussed, the droplet size correlating and distribution equations

presented in Ref. 47 are incorrect, and the equations in this latter paper (Ref.

52) are correct. Dickerson also notes that the quality of the wax spheres was

poorer than usual for these pentad tests.

ZaJac (Ref. 58) performed a few similar tests and found that the very few higher

velocity tests were In crude agreement with the correlations of Dlckerson. Most

of Za_ac's tests were at lower velocities and were in great disagreement wlth

Dickerson's correlating equation (Dlckerson dld not perform tests at these lower

velocities). The deviation at the low velocities Is speculated to be due to velocity profile and/or laminar flow effects. ZaJac speculates that the flow regime

of the outer streams Is more important than that of the inner stream.

As a part of the investigation of Mehegan, et. al. (Ref. 55) of gas/llquld injectors, atomization characteristics were determined for a set of pentads. These

experiments employed wax and hot gas, with the central orifice always flowing the

gas. These tests were performed at atmospheric pressure with variations In gas

and liquid velocity and orifice sizes. No correlating equations were developed.

RI/RD85-312

A-37

ORIGIN ,AL PAGE

TABLE 3. DROPLET SIZE CORREI.

REFERENCE

I

MEASUREMENT TECHNIQUE i

DICKERSON (1969) WAX AND WATER

ZAJAC (1971) WAX AND WATER

CORRELATION

= 8.26 x 105 dc'ld°'12

c V "086V .89

c o

d .68

= 5.66 x 106 o

o Vo-56 Vc .57 dc "35

= 30

pg = I.

do AND dc = .01

WAX AND H20 PR_I

AGREES WITH DICKERSON EXCEPT AT

LOW VL

(SEPARATE EQUATIONS SHOULD BE DERIVED

FOR LAMINAR)

= i.

Pg

d AND d =

o c

VL = 9

MEHEGAN, ET. AL.

(1970)

WAX AND N2 (OR He )

CENTRAL ORIFICE GAS

GAS HEATED >140 F

D_d

°/d c

-D _ FOR > .8, OTHERWISE

c INDEPENDENT

MR NOT IMPORTANT

LARGER DROPLETS FOR PENTADS THAN

TRIPLETS

= 45!.

dc = .6r

d = .I

0

Vc (GAS) = 18

WAX PROPERTIES

SUBSCRIPT:

c = CENTER STREAM

o = OUTER STREAM

ALL d VALUES IN cm

ALL v VALUES IN cm/s

ATIONS FOR PENTADS

CONDITIONS

(ANGLE BETWEEN CENTER & OUTER STRE#MS)

kg/m 3

35 - .218 cm

_PERTIES

VL = 16.8 - 77.7 m/s _

DIFFERENT DISTRIBUTIONS FOR

INNER AND OUTER STREAMS

POORER QUALITY WAX SPHERES

kg/m 3 _ = 30 °

- .218 cm

4 - 18.3 m/s

VELOCITY PROFILE EFFECTS MAY

BE IMPORTANT

RELATIVELY FEW TESTS

2.1 cm

.52 cm

-914 m/s

pg = 1.1 kg/m 3

d

o .161 - .317

d

c

X FmLVL

d_cc: .293-]. 14 : 2.5 L gVc

(PENETRATION PARAMETER)

cqs 2 (90-e)l 1⁄2

RI/RD85-31 2

A-3B

However, the proporttonallty relationships shown In Table 3 were observed.

Higher gas dynamic forces (pg V_) tended to reduce drop size, and the

degree of penetration of the outer liquid streams into the central gas stream

also affected atomization. Thts degree of penetration Is quantified In a mixing

assessment parameter, the pentratlon parameter, which is defined as shown in

Table 3. The droplet size was observed to increase with the penetration factor

at higher values.

The capablllty to assess the atomization characteristics of pentads is only

slightly better than that of triplets. At present, there is no way to extrapolate to the actual propellants. Also, there Is no data by which to assess the

effects of combustion gas motion.

Coaxial Correlations

The standard coaxial injector (gas flowing through the annulus about a central

liquid stream) has been studied more extensively than gas/liquid triplets and

pentads. Again, all of this work was performed at Rocketdyne utilizing the hotwax technique. These efforts are synopsized In Table 4.

As a part of the investigation of gas/liquid injectors by Nehegan, et. al. (Ref.

55), a number of tests of the effects of inner tube recess, gas and liquid velocities, and mixture ratio on mass median droplet size were performed. These tests

were performed with a large coaxial element. The results of these tests are presented graphically In Flg. 7. Since it is generally believed that the liquid

stream breakup Is dependent strongly on aerodynamic forces, the velocity difference between the gas and liquid Is often of great concern. The effect of inner

tube recess also Is shown to be of great importance In Flg. 7. A smaller element also was tested, which indicated little effect of recess, but this data was

considered questionable. Also, the sizes of the droplets produced by this large

injector were quite large.

®

RI/RDS5-312

A-39

REFERENCE MEASUREMENT TECHNIQUE

I

!

TABLE 4. DROPLET SIZE CORRELATI[ ..

CORRELATION

MEHEGAN, ET. AL.

(1970)

BURICK (1972)

McHALE AND

NURICK (1974)

WAX AND HOT N2

SPRAYED INTO ATMOSPHERE

WAX AND HOT N2

SPRAYED INTO PRESSURIZED

TANK

WAX AND HOT N2

SPRAYED :NTO PRESSURIZED

TANK

WL

g

R = RECESS iN UNITS OF dL

INCREASE Va (BY DECREASING Y AT

CONSTANT WG) DECREASES D GREATLY

FOR SMALLER Y

INCREASE VL (BY DECREASING dL AT

CONSTANT WL) DECREASES

i

D/Y DECREASES AS pqVg 2 INCREASES

BUT THE EFFECT IS EESS AT

HIGHER VL

dL = CENTER TUi

Y = ANNULUS G_

WL

-- : 5.7 - 13.[ ....

= 12.'_

VL 1.6 -

dL .177, .27

Y .013, .04

LVL- 3 - 7.5 _i_

w9

VL = 12.2 - 42

!

pq = .9 - 4.3 1 

VL = 1.2- 16.[_;

V = 91.4 - 301

g

WL _ 6

g

FALK (1975) WAX AND HOT N2

SPRAYED INTO FLOWING GAS

(f VL (CONSTANT d L) SMALL EFFECT

: 1/Vg (ESPECIALLY WHEN VCG = O)

_ WL/_g (ESPECIALLY AHEN VCG = D)

_ I/L (NEGLIGIBLE EFFECT EXCEPT AT

LARGE Do

Do = DROP SIZE WHEN VCG = 0

d L = .14 - .41

V = .25- 1.4i

V9 = ANNULUS Gi

= 61-305 m/ii

VL = 23 - 76 m;

Vcg = SIMULATEDi"

VELOCITY !.i

L : LENGTH OV[

ACCELERAT!

i : .7

i _----L

NS FOR COAXIAL IN3ECTORS

_E ID = .704 cm

_P = .259 cm

.345 cm

.104 cm

m/s

7 m/s

g/m 3

m/s

m/s

VELOCITY

COMBUSTION GAS MAX

61-244 m/s

R WHICH GAS WAS

ID = 5-20 cm

V = 112 - 339 m/s

g

b

CONDITIONS I

R = 0 - 2dL

GENERALLY THESE LARGE

COAXIAL ELEMENTS PRODUCED

LARGE DROPS

A SMALLER ELEMENT WAS

TESTED AND SHOWEDDIFFERENT

EFFECT OF R. THIS DATA WAS

CONSIDERED QUESTIONABLE

R : 0 - 4 dL

pg = 3.1 - 22.7 kg/m 3

POST REcEss HAD NO EFFECT

DIFFERENT RESULTS FOR

DIFFERENT SIZE ELEMENTS

PLOTS OF DISTRIBUTION DATA

R = 0 - Id L

dL = .41 cm

Y = .13 cm

SLIGHT D DECREASE AS R

INCREASES

_VL

-- = .1, .5, 2.0

g

SAME DROP SIZE DISTRIBUTION

AS ZAJAC LIKE DOUBLETS

|

!

i

RI/RD85-312

A-40

I,.-..l

Go

t

1400

1200

_D

1000

X

CO

cc

L,U

I--

,,, 800

:¢

frO

Z

o 6oo

er400

2OO

0

I

6

d L = .704 CM

Yg = .259 CM

0Dg = 1.37 CM

I 1 I I l I

7 8 9 10 11 12

MR

Vg -- V L = 360 ft/sec(110 m/sec)

RECESS = 0

RECESS= 1D L

RECESS = 2D L

Vg- V L = 1080 ft/sec (329 m/sec)

I I

13 14

Figure 7. Mixture Ratio Influence on Dropslze

for Coaxial Injectors (Ref. 55)
Investigations by Burlck (Ref. _64 and 67) and McHale and Nurlck (Ref. 74) utilIzed hot wax and nitrogen injected into a pressurized chamber to examine coaxial

element atomization. The combined work of these two studies, along with the previously discussed work of Mehegan, indicate that recess only reduces D at low

pressures and/or for large elements. Burlck correlated hls data as shown in

Fig. 8. Again, the normalization and logarlthmlc plotting of the data masks the

"spread" of the data. Although McHale and Nurlck were investigating primarily

the atomization characteristics of nonclrcular orifices, they did perform limited

tests on circular orifices. Their data indicates that increased annulus gas

dynamic pressure (pg V_)_ reduces droplet size, especially at low liquid

velocity. References 67 and 74 present droplet size distribution plots.

Although McHale and Nurlck state that recess Is a major factor influencing droplet size, this conclusion is based upon tests of all of their injectors, which

are prlmartly noncircular. The llmtted testing performed with circular coaxlal

elements indicates a 10 to 20% reduction in drop size as recess ls increased to

R = d L. Even the noncircular elements do not show an effect of recess anywhere

near as significant as that found by Mehegan, et. al.

Falk (Ref. 78) investigated the atomization characteristics of coaxial elements

injecting wax and hot nitrogen cocurrently (i.e., axially) into a duct flowlng

hot nitrogen. This work utilized the same test apparatus and techniques as the

analagous work of Zajac (Ref. 70 and 71) on llke doublets. One potentially very

important finding of this work was that the droplet size distribution of these

coaxial injectors could be described by the distribution function defined by

Zajac for llke doublets. Also, the mass median droplet sizes observed In both of

these investigations were essentially the same at high relative gas velocity.

This would seem to indicate that the manner in which the liquid Is broken up

(i.e., the type of injector) has no effect on the ultimate droplet size in the

presence of a sufflc_ently accelerating combustion gas. If this is truly the

case, it is a most important discovery that wlll direct the course of future

studies of rocket engine injectors.

RI/RD85-312

A-42
..J

"0

IC3

1.0

0.6

0.4

0.2

0.1 -

0.6

0.1

/k

I /k• RECESSED FLUSH POSTPOST i Z_

1 i I i I i , I I ] , I J I i i _ l

o.2 o.4 o.6 1.0 2.o 4.o 6.o 10.

(Vg -- VL)/IVLVV I/_Vi)

20

_L

Figure 8. Correlation of Cold-Flow Atomization for Coaxial Injectors (Ref. 67)

RIIRD85-312

A-43
Some of Falk's results, showing the effect of simulated combustion velocity on

droplet size, are presented In Fig. 9. This data indicates that injectors, which

form larger droplets when no combustion gas motion Is simulated (i.e., when

V = 0), show more effect of this gas motion than injectors producing smaller

cg

droplets. Recognizing this important influence of the relative combustion gas

velocity on the droplet size, Falk correlated 'the data in a manner shown In

Fig. 10. This correlation is based only upon the relative, simulated, combustlon

gas velocity and Do , the mass median droplet size produced by an injector In

the absence of this gas flow.

r _

ATOMIZATION SURVEY - FINDINGS, CONCLUSIONS, AND RECOMMENDATIONS

The state of the art regarding our knowledge of atomization processes Is generally quite poor. The physics Is poorly and, at best, only qualitatively understood. Only very rudimentary quantitative theories exist. The available data

and correlations are generally of questionable validity and/or utility. Many of

the most critical parameters are unknown (e.g., combustion gas velocity field,

multiple element effects) and/or are not simulated In tests (e.g., gas densities,

real propellant fluid properties, combustion gas motion). Thls sad state of

affairs appears to be attributable to two primary causes: the great complexity

of atomization processes, and the inaccuracies, errors, and limitations associated wlth droplet size measurement techniques. Nevertheless, the available data

does provide information regarding the importance and relative effects of a number of variables on droplet size.

V

Probably the most critical of these parameters affecting droplet size Is the combustion gas velocity field. Thls is unfortunate since the actual velocity field

In a rocket combustor, and In atomization experiments, Is unknown. Combustion

gas velocity also Is the one parameter that greatly increases the complexity of

the atomization assessment problem. Thls Is due to the fact that atomization Is

highly dependent on the combustion gas velocity field, and In turn, the combustion gas velocity field Is established by the rate of combustion, which is determined by the rate of propellant evaporation, which Is highly dependent on how

well the propellants are atomized (i.e., initial droplet sizes) and mixed. Thus,

all of these problems are coupled and the solution of any one requires at least

an approximate solution of each of them.

RI/RD85-312

A-44
I

4:,.

_o

CO

200

180 ̧

160

U

"- 140

120

100

I / Vgac

I

I

I

-0

I

Gas/Llautd Coaxtal Element

Vg - 30S m/see (1000 ft/sec)

VL - 22.9 m/see (75 ft/sec)

_/L _gl0." 2.0 L 16 cm (4.0 Inch]

A dL• .411 c.m(0.162 ]nch)

dL• .239 an (0.094 Inch]

• d L = .140 cm (0.055 Inch)

I

I

I

I

i

4k_

A

I I I I I I I I

0 1 2 3 4 5 6 7

(v - v,)/vL cq I.

! I

8 9 10

Figure 9. Influence of Combustion Gas Simulant Velocity on

Mass Median Dropslze - Coaxial Injectors (Ref. 78)
.¢. ¢0

Cn _n

I

r_3

1.0 _,

0.4

o __

nO

COAXIALELEHENT

L - 10.16 cm (4-tnch)

110_ _ 1_o _ 310.

AV - (Vcg VL) - -15 to 200 m/sec

(-50 to 650 ft/ser)

0 0

0

In

I ! I I I

6 7 8 9 10

,o,v,l

Figure 10. Correlation of Droplet Breakup (Atomization) Data

- Coaxial Injectors (Ref. 78)

( (
F ....

All of the droplet size measurement techniques applied to atomization studies

have serious limitations and potential and/or known sources of error. Imaging

techniques measure the spatial concentrations of the various size droplets. Such

spatial concentrations can be utilized only rarely to define the actual droplet

slze distribution or representative droplet size characterizing all of the droplets produced by a given spray (temporal distribution). Spatial and temporal

distributions are often qu_te different. Thus, the photographic techniques and

the droplet freezing (1.e., hot wax) technique do not measure the same thing.

In order to utilize cold-flow atomization data, it is necessary to be able to

account for the effects of the different liquids' properties on the droplet

sizes. The only data available for this purpose applies to llke doublets, Is of

questionable validity and applicability, and differs from one investigation to

another. No methods have been proposed to accomplish this properties effects

correlation for any gas/llquld injector or for any llquld/llquld injector except

llke doublets. No attempts have been amde to assess injected fluids properties

effects on droplet size distributions.

Very little information could be found regarding the atomization characteristics

of triplet, pentad, and coaxial injectors. Such data, as Is available, is presented along wlth a representative sampling of the data for llke doublets.

The following actions are recommended for the purpose of (1) improving our knowledge of atomization processes, (2) developing the droplet size data required by

the combustor analysis codes, and (3) utilizing the data in such codes. These

actions are divided into near and long-term approaches.

Near term: For the most immediate future, it Is recommended that droplet size

data for combustor analysis be determined in the following manner. First, the

existing data can be utilized (it should be verified first, however) and/or tests

can be performed to better define DO , the droplet size produced in the absence

of any simulated combustion gas motion. Thls can be considered primary atomlzatlon. Then the data and correlations of Falk (Ref. 78) and ZaJac (Ref. 70 and

71) can be employed to estimate the effect of gas velocity on droplet size. In

order to do this It Is, of course, necessary to estimate the combustion gas

RI/RD85- 312

A-47
velocity where the injector is to be employed. Thls can be accompllshed through

the use of combustor performance computer codes that generally compute the axtal

velocity of the gas. Thus, at least the major (hopefully) gas phase velocity

component will be estimated. Since the computed axial gas phase velocity will

depend on Initial droplet sizes, a few iterations of this process may be necessary. That ts, the codes can be used to predict Va, which then can be used to

estimate D, which will be Input to the codes to predict a new Vg, etc.

Another problem In the use of cold flow droplet size data ls that It Is necessary

to convert from the test flulds to the real propellants. With great reservation,

and only because no better information Is avallable, the properties effects correlations of Ingebo (Ref. 13) or Wolfe and Anderson (Ref. 37) are recommended for

this purpose, when llquid, like impinging elements, are being considered.

The method described above provides a rudimentary technique for estimating a

representative droplet size. Drop size distributions In general, and representative droplet slze information for gas/liquid injectors, cannot be estimated via

this technique due to the lack of data regarding combustion gas velocity effects

and fluid properties' effects on atomization. Even when applied to the case of

like doublets, which have been most extensively studied, this technique may be

little better than a consistent guessing method.

In order to better utllize this technique and improve Its accuracy the following

are recommended:

1 •

2.

•

Experiments to investigate gas veloclty effects on droplet sizes

Additional tests to better define Do for the injectors of greatest

Interest, especially gas/liquid injectors. Most of the geometric and

operational variables have not been tested

Experiments to establish fluid properties effects for all types of

injectors, like and unlike liquid and gas/liquid Injectors, and separate

effects for primary and secondary atomization

RI/RD85-312

A-48

V
-v

Such studies and experiments will provide the basis for improvements to atomization assessment methods and will establish the nature, feasibility, and desireability of pursuing the long-term approach.

Long term: As previously discussed, due to the importance of the combustion gas

motion on atomization, the problem becomes coupled with those of droplet evaporation, combustion, and three-dlmenslonal fluid mechanics with momentum and mass

sources and sinks. Unless some simplifying assumptions are identified earlier,

the only available solution would consist of a coupling and solution of all the

equations governing these processes. This would probably involve a long-term

effort consisting of a number of programs to model (probably with a computer

code) various parts of the problem, experimentally verify these models, and combine them in one comprehensive model. Such an approach offers the greatest

potential for a comprehensive, accurate, proven solution to the problem of spray

definition for rocket engine injectors. If a satisfactory measurement technique

exists, experiments with operating, small-scale rocket combustors should be performed to validate the atomization model. In its ultimate form, such a model

would include multiple element effects and would predict mixing efflclencles.

RI/RD85-31 2

A-49
NOMENCLATURE

Someatomization nomenclature is defined in text.

A

d

D

D

D0

D

C

f

nt

N

area (cm 2)

injector orifice diameter (cm)

droplet diameter (microns)

mass median diameter

droplet mass median diameter observed when Vg = 0

droplet mass median diameter observed when Vg = VL

droplet distribution function, (drops/mlcron)

f = llm 9_

AD_o AD

length over which gas is accelerated in

accelerating gas flows (cm)

number of droplets counted in a given slze range

total number of droplets counted

cumulative number distribution,

D

N(D) =_n

0

atomization studies in

v

V

AP

R

V

V

gm

V

V

Vtot

VI

flowrate of droplets of size group I

injector orifice pressure drop (Pascals)

normalized cumulative volume distribution, R = V/Vto t

mass rate of flow (kg/s)

velocity (m/s)

maxlmum gas velocity

cumulative volume distribution,

D

v (o) = X v

0

volume of all drops In a given size range (cm3)

total volume of all droplets counted

velocity of drops of size group I

RI/RDBS-31 2

A-50
Vcg

Y

Z

P

oi

simulated combustion gas maximum velocity (gas/liquid injectors only)

annulus gap for coaxial injectors (cm)

axial spatlal coordinate

Impingement angle

viscosity (cP)

density

concentration of drops of slze group I (drops/cm 3)

surface tension (dynes/cm)

Subscripts

g

I

J

L

L

S

gas (either local chamber gas or injected gas)

slze group of droplets

Jet or orifice

large drops

liquid

small drops

©

RI/RD85-312

A-5I
REFERENCES

1. Longwe11, 3.P.: "Fuel 011 Atomization," D. Sc_ Thesis, HIT, 1943.

2. Hlnze, J.: "Critical Speeds and Sizes of Liquid Globules," ADpl. Scl. Res.,

1948, Vol A1.

3. Lane, W.: "Shattering of Drops In Streams of Air," Industrial & Englneerlnq

Chemistry, Chemical Defense Experimental Establishment, England, 3une 1951,

Vol. 43, No. 6.

4. Mugele, R., and H. Evans: "Droplet Slze Distributions In Sprays," Industrlal & Englneerlng Chemistry, Chemical Defense Experimental Establishment,

England, 3une 1951, Vol. 43, No. 6.

5. Bowen, I., and G. Davies: "Particle Size Distribution and Estimation of

Sauter Mean Diameter," Shell Petroleum Co. (London), October 1951, Report

No. I.C.T. 28.

6. Heldmann, M., and 3. Humphrey: "Fluctuations In a Spray Formed by Two

Impinging 3ets," ARS Journal, May-June 1952, Vol. 22, No. 3.

7. Moore, B., and A. Hussman: "Fuel Injection Studies with Impinging Jet Nozzles", Penn State University, October 1955, ONR Contract 656(07), AD 119844.

8. Fraser, R., and P. Elsenklam: "Liquid Atomizatlon and the Drop Size of

Sprays," Trans. Instn. Chem. Engrs., 1956, Vol. 34.

9. Tanasawa, Y., S. Sasakl, and N. Nagai: "The Atomization of Liquids by Means

of Flat Impingement," Technoloqy Reports of the Tohoku Unlverslty, 1957,

Vol. 22, No. 73.

lO. Graves, C., and D. Bahr: "Atomization and Evaporation of Liquid Fuels,"

Basic Considerations in the Combustion of Hydrocarbon Fuels wlth Air, NACA

Report 1300, Chapter l, 1957.

II. Putnam, A., et. al.: Injection and Combustion of Liquid Fuels, Battelle

Memorial Institute, March 1957, WADC-TR-56-344, AD I18142.

12. Heldmann, M.F., R.J. Priem, and J.C. Humphrey: "A Study of Sprays Formed by

Two Impinging Jets," March 1957, NACA TN 3B35.

RI/RD85-312

A-52
_J

13. lhgebo, R., and H. Foster: ."Drop Size Distributions for Crosscurrent Breakup of Liquid Jets in Airstreams," October 1957, NACA TN 4087.

14. Fuhs, A.: "Spray Formation and Breakup and Spray Combustion," February

1958, AFOSR-TN-58-414, AD 158217.

15. Ingebo, R.: "Drop Size Distributions. for Impinging Jet Breakup in Airstreams Simulating the Velocity Conditions in Rocket Combustors," March

1958, NACA TN 4222.

16. Weiss, M.A., and C.H. Worsham: "Atomization in High-Veloclty Airstream,"

ARS Journal, April 1959.

17. Rossman, T.: "A High-Speed and High-Resolutlon Photographic Technique for

the Observation of Propellants Injected into a Firing Chamber," May 1959,

AFOSR-TN-59-8 (2 parts).

18. Benson, G., M.E. Walil, P. Myers, and O. Uyehara: "Fluorescent Technique

for Determining the Cross Sectional Drop Size Distributions of Liquid

Sprays," ARS Journal, May 1960, Vol. 30, No. 5.

19. Rossman, T.: "Observation of Propellants Injected into a Firing Rocket

Chamber," July 1960, AFOSR-TR-60-98.

20. Dykema, 0.: "A Study of Injector Geometry by Spray Analysis," Rocketdyne

Report 60-19, August 1960.

21. Heidmann, M.: "Photography & Analysis of Time Variation in Drop Size Distribution of a Liquid Spray," Fifth International Congress on High-Speed

Photography, Paper N-7, Washington, D.C., October 1960.

22. laylor, G.: "The Dynamics of Thin Sheets of Fluid," Proceedings of the

Royal Society of London, November 1970, Vol. 259 A.

23. Hasson, D, and J. Mizrahi: "The Drop Size of Fan Spray Nozzles: Measurements by the Solidifying Wax Method Compared With Those Obtained by Other

Sizing Techniques," Trans. Instn. Chem. EnSOl_E=., Chemical Engineering Department, Technion, Halfa, Israel, Ig61, Vol. 39.

24. Ungureanu, C.: "Some Results Concerning the Distribution of Drops of Fuel

Atomized Through Low-Pressure Injectors," Rumanlan periodical _En_eLq_etIca,

1961, Vol. 4, No. 3, (FID-TT-63-29/I+2+3+4).

RI/RD85-312

A-53
25.

26.

Heldmann, M.F., and H. Foster: "Effect of Impingement Angle on Drop Size

Distributions and the Spray Pattern of Two Impinging Water Jets," July 1961,

NASA TN 0-872.

Ingebo, R.: "Size Distribution and Veloclty of Ethanol Drops in a Rocket

Combustor Burning Ethanol and L1quld Oxygen," ARS Journal, Aprll 1961 (also

see NASA-TN-D-290, June 1960).

27. Popov, M.:

July 1961.

28. Fraser, R.:

"Model Experiments on the Atomization of Liquids," NASA-TT-F-65,

"Liquid Atomization," Journal of Royal Aeronautical Soclety,

Imperial College, London, November 1961, Vol. 65.

29. Neiss, M., and C. Worsham: "Nozzle Sprays in Air Streams," Chemical Engineering Science, Esso Research & Engineering Co., New Jersey, December 1961,

Vol. 16, No. 1 and 2.

30. Lewis, O.: "Studies of Atomization and Injection Processes in the Liquid

Propellant Rocket Engine," Rocket Propulslon Establishment TM No. 241, Westcott, Great Britain, December 1961.

31. Lambris, S., L. Combs, and R. Levlne: "Stable Combustion Processes in

Liquid Propellant Rocket Engines," 5th Colloquium of the Combustion and Propulsion Panel, 1962, NATO, AD 283525.

32. Dombrowskl, N., and P. Hooper: "The Effect of Ambient Density on Drop Formation in Sprays," Chemical Engineerlnq Science, Imperial College, London,

1962, Vol. 17.

33. Dombrowski, N., and P. Hooper: "The Performance of Characteristics of an

Impinging Jet Atomizer in Atmospheres of High-Ambient Density," Fue_____!l, 1962,

Vol. 41.

34. Fraser, R., P. Eisenklam, N. Dombrowskl, and D. Hasson: "Drop Formation

from Rapidly Moving Liquid Sheets," A. I. Ch.E. Journal, Imperial College,

London, November 1962, Vol. 8, No. 5.

35. Oombrowskl, N., and P. Johns: "The Aerodynamic Instability and Disintegration of Viscous Liquid Sheets," Chemical Engineering Science, Imperial College, London, 1963, Vol. IB.

RI/RD85-312

A-54
36. Brown, R., and K. Leonard: "Methods of Describing Droplet Slze Distrlbutions from Atomized Solutions," keroJet Report 0395-04(15) SP, kO 434106,

March 1964.

37. Wolfe, H., and W. Anderson: "Kinetics, Mechanlsm, and Resultant Droplet

Sizes of the Aerodynamic Breakup of Liquid Drops," AeroJet General Report

0394-09(18)SP, kO 437340, April 1964.

38. Brown, R.E.: "The Atomization of a Solution of 2, 4 - Dlhydrobenzophenone

in Bls (2 Ethylhexyl) Hydrogen Phosphate," AeroJet General Report 0395-

04(20)SP, A0601462, May 1964.

39. Clark, R.: "Breakup of a Liquid Jet %n a Transverse Flow of a Gas," August

1964, NASA TN 0-2424.

40. Oombrowskl, N., and P. Hooper: "A Study of the Sprays Formed by Implnglng

Jets In Laminar and Turbulent Flow," Journal of Fluid Mechanics, 1964, Vol.

18, Part 3.

41. Lewis, J.: "Some Basic Studies of Liquid Propellant Injection Processes,"

Journal of the Royal Aeronautical Society, Rocket Propulsion Establishment,

England, November 1964, Vol. 68.

42. Lapple, C., J. Henry, and O. Blake: "Atomizat%on - A Survey and Critique of

the Literature," Stanford Research Inst., Report No. 6, April 1967, kO

821314.

43. Luna, R., and W. Kllkoff: "On Aerodynamlc Breakup of Liquid Drops," Sandla

Laboratories Research Report SC-RR-66-2716, June 1967.

44. Hiroyasu, H.: "Mathematical Expressions for Drop Size Distribution in

Sprays," NASA-CR-72272, November 1967.

45. Dombrowskl, N., and G. Munday: "Spray Drying," Biochemical and Blolo_Jlcal

Engineering Science, Academic Press, 1968, Chapter 16.

46. Adelberg, M.: "Mean Drop Size Resulting From the Injection of a Liquid Jet

into a High-Speed Gas," AIAA Journal, June 1968, Vol. 6, No. 6.

47. Dickerson, R., K. Tate, and N. Barslc: "Correlation of Spray Injector Parameters wlth Rocket Engine Performance," Rocketdyne Report R-7499, AFRPL-TR68-147, June 1968.

RI/RD85- 312

A-55
48. Sloat, T.: "Dropslze Measurements with Large-Thrust Coaxlal and Trtplet

Elements," Rocketdyne Internal Letter [C69-344'24, June 1969.

49. Ingebo, R.: "Maximum Drop Diameters for the Atomization of Liquid Jets

Injected Concurrently Into Acceleratlng or Decelerating Gas Streams," NASATN-D-4640, July 1968.

50. Knight and Nur%ck: "Interlm Report, Correlation of Spray Oropslze Distribution and Injector Variables," Rocketdyne Report R-7995, September 1969.

51. Sloat, T., and D. Campbell: "Atomization Characteristics of Gas/Liquid

Injectors," 6th Liquid Propellant Combustion Instability Conference,

Chicago, September 1969.

52. Dickerson, R.: "Like and Unlike Implnglng Injection Element Droplet Sizes,"

3. Spacecraft, November 1969, Vol. 6, No. ll.

53. Walkden, A., and R. Kell: "Characteristics of High-Flux Sprays From CollidIng Water Jets," Trans. Instn. Chem. Enqrs., General Electric Co., Wlmbly,

Middlesex, England, 1969, Vol. 49.

54. Kuykendal, W.: "The Effect of Injector Design Variables on Average Drop

Slze for Implnglng Jets," AFRPL-TR-70-53, Hay 1970.

55. Mehegan, P., D. Campbell, and C. Scheuerman: "Investigation of Gas Augmented Injectors," Rocketdyne Report R-8361, NASA-CR'72703, September 1970.

56. Nurick, W., and R. McHale: "Nonctrcular Orifice Holes and Advanced Fabrication Techniques for Liquid Rocket Injectors," Rocketdyne Report R-8224,

NASA CR-108570, September 1970.

57. Huang, 3.: "The Breakup of Axisymmetrtc Liquid Sheets," J. Fluid Hech,

1970, Vol. 43, Part 2.

58. ZaJac, L.: "Correlation of Spray Dropslze Distribution and Injector Variables," Rocketdyne Report R-8455, Contract NAS7-726, February 1971.

59. Klm, K., and W. Marshall: "Drop Slze Distributions From Pneumatic Atomlzers," A. I. Ch.E. Journal, University of Wisconsin, May 1971, Vol. 17,

No. 3.

V

RI/RD85-31 2

A-56

V
v

60. Nurlck, W.: "Analysis of Sprays From Rocket Engine Injectors," Journal of

Spacecraft and Rockets, Rocketdyne, July 1971, Vol. 8, No. 7.

61. ZaJac, L., and W. Nurlck: "Correlation of Spray Dropsize and Injector Varlables," 8th JANNAF Instability Conference, October 1971.

62. Slmpkins, P., and E. Bales: "Water Drop Response to Sudden Accelerations,"

Journal of Fluid Mech., Bell Labs & Stevens Inst. Tech., 1972, Vol. 55,

Part 4.

63. Combs, L.: "Catalog of Injector Spray Correlatlons," Rocketdyne, NASA Contract NAS7-746, 1972.

64. Burlck, R.: "Atomization and Mixing Characteristics of Gas/Liquid Coaxlal

Injector Elements," Journal of Spacecraft, May 1972, Vol. 9, No. 5.

65. Nurlck, W.: "Study of Spray Disintegration in Accelerating Flow Fluids,"

Rocketdyne Report R-9017, NASA CR-114479, June 1972.

66. George, D., and F. Spaid: "Holography As Applied to Jet Breakup & An Analytical Method for Reducing Holographlc Droplet Data," AFRPL-TR-72-72, September 1972.

67. Burlck, R.: "Space Storable Propellant Performance Program Coaxial Injector

Characterizatlon," Rocketdyne Report R-8973-2, NASA-CR-120936, October 1972.

68. Falk, A.: "Space Storable Propellant Performance Gas/Liquid Like Doublet

Injector Characterization," Rocketdyne Report R-8973-1, NASA-CR-120935,

October 1972.

69. Nurick, W.: "Physical Property Effects on Spray Atomization," Rocketdyne

IR&D Report ITR-73-O14-C, September 1973.

70. ZaJac, L.: "Droplet Breakup in Accelerating Gas Flow, (Primary Atomtzatlon)," Part 1, Rocketdyne Report R-9337-1, NASA-CR-134478, October 1973.

71. ZaJac, L.: "Droplet Breakup in Accelerating Gas Flows (Secondary Atomization)," Part 2, Rocketdyne Report R-9337-2, NASA-CR-134479, October 1973.

72. George, D.: "Rocket Injector Hot-Firing and Cold-Flow Spray Fields," AIAA/

SAE 9th Propulsion Conference, Las Vegas, November 1973.

R1/RD85--312

A-57
73. George, O.: "Rocket Injector Hot-Firlng and Cold-Flow Spray Fields," lOth

3ANNAF Combustion Meet, December 1973, CPIA No. 243, Vol. 3.

74. McHale, R., and W. Nurlck: "Nonclrcular Orifice Holes and Advanced Fabrication Techniques for Liquid Rocket Injectors (Phases I, II, III,& IV),"

Rocketdyne NASA CR-R-9271, May 1974.

75. Buschulte, W.: "Liquid Propellant Atomization by Injector Elements and Its

Effects on Combustion Chamber Efficiency," Israel Journal of Technology,

1974, Vol. 12.

76. George, D.: "Droplet Size Distribution Functions For Rocket Combustor Spray

Fields," llth 3ANNAF Combustion Meet, December 1974, CPIA No. 261.

77. Anon: "Analysis of Rocket Engine Injectlon/Combustlon Processes," Rocketdyne Report R-9668P-l, March 1975.

78. Falk, A.: "Coaxial Spray Atomization in Accelerating Gas Stream, Final

Report," Rocketdyne Report R-g753, NASA-CR-134825, June 1975.

79. Rizkalla, A., and A. Lefebvre: "The Influence of Air and Liquid Properties

on Alrblast Atomization," Journal of Fluids Engineering, Cranfleld Inst.

Tech., Trans ASME, September 1975.

80. Rao, K., and A. Lefebvre: "Fuel Atomization in a Flowing Alrstream," AIAA

Journal, October 1975, Vol. 13, No. lO.

81. Farmer, W.: "BRL Particle Sizing Interferometer," University of Tennessee

Space Institute, Tullahoma, Tenn., circa 1980.

82. Lefebvre, A.: "Airblast Atomization," Prog. Enerqy Combustion Science, Purdue University, 1980.

83. Elkotb, M.: "Fuel Atomization for Spray Modeling," Progress in Energy &

Combustion Science, Cairo University, 1982, Vol. 8, No. l.

84. Ferrenberg, A.: "Liquid Rocket Injector Atomization Research," Proceedings

of ASTM Symposium on Liquid Particle Size Measurement Techniques, Kansas

City, June 1983.

RI/RD85-312

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V
_i MIXING

Cold flow mixing tests frequently have proven to be a significant aid in

predicting potential performance, or diagnosing problems with rocket engine

injector components. Cold flow tests are not sufficiently reliable so as to

serve as a replacement for hot-flre testing, but should be considered as complementary to hot-flre tests, aiding in minimizing the number of hot-flre tests

required to obtain an optimum configuration. In almost every case, an injector

or element that performs poorly in cold flow testing will not perform well in

hot-flre testing. However, the counter side of this statement cannot be applied

universally. An element can be excellent in cold flow mixing, but the combustion

reaction may override the hydromechnlcal mixing provided by the injection

streams. This effect is most notable with storable hypergollc propellants, where

a phenomena of reactive demlxlng "blowapart" is frequently a significant factor

in combustion performance. There have been other reports of combustion systems

suffering from reactive demlxlng, but none have been as well documented as the

hypergollc reaction systems.

Aerodynamic forces in the combustion zone also are factors that cannot be simulated in cold flow mixing tests. Gas forces in reclrculatlon can be strong factors influencing mixing and atomization. There are, however, useful correlations

between cold flow mixing and combustion results, and the relative cost factor

between cold flow and hot-flre tests generally is a rational reason for utilizing cold flow tests as an injector design and development tool.

The key objective, to establish correlations between cold flow mixing data and

hot-fire results, requires a large empirical data base as well as a consistent

assessment of the data and an applied scientific evaluation of the resultant correlating parameters. Therefore, an assessment criteria was established, which

allowed compilation of existing cold flow experimental data acquired within the

industry on element types suitable for LOX/hydrocarbon injector advancement.

The triplet, pentad, and coaxial element injection devices were selected for

study based on available hotflre and cold flow experience wlth LOX/hydrocarbon

RI/RD85-312

A-59
propellants. The cold flow correlating parameters used for these devices were

identified and an extensive literature survey conducted to obtain related cold

flow data. Data from the literature search was compiled Into a displayable format. The information then was plotted by the appropriate correlating parameter(s) against mixing efficiency, a standard measure of cold flow performance.

V

In addition to the literature survey, flve impinging triplet elements, one pentad

element, and three coaxial elements were fabricated for cold flow testing. The

sizing of these elements encompassed designs for both preburner (gas generator)

and maln injector mixture ratios at high chamber pressure. The propellant combinations were LOX/methane (gas/llquld), LOX/RP-1 (llquld/llquld), and LOX/propane (llquld/llquld and gas/llquld). The low-pressure cold flow mixing test program was conducted wlth these elements at several flow conditions. Measures of

mlxlng efficiency were established and plotted as a function of mixing parameters. Maps depicting mixture ratlo-normallzed mass distribution were constructed from the cold flow tests to provide a good visual indication of relative

mass and mixture ratio concentrations for the different element types.

IN3ECTOR MIXING CORRELATING PARAMETERS

Mixing correlation parameters are mathematical expressions based upon injector

element geometry and flow conditions. Their utility as injector design criteria

depends upon (1) their ability to be related to mixing efficiency and (2) the

existence of optimum values of these correlation parameters at which mixing will

be maximized.

Numerous correlating parameters have been proposed for different injector configurations, propellant conditions, and hot flre related operating conditions. The

scientific basis for the parameters generally has been derived from momentum and

stream diameter relationships of the injection element. A survey of available

literature showed that of these relationships, most correlating parameters were

derived for llquld/llquld Implnglng-type injectors. Many of the experimenters

have established formulas to plot data from numerous test conditions on a single

curve, or at least, within a family of curves.

RI/RD85-31 2

A-60

V
r= = The correlating parameters used in the literature survey data reduction and in

the subsequent low-pressure mixing tests are presented in Table 5. A description of these important parameters is discussed below. Illustrations of the

three element types studied under this program (coaxial concentric tube, triplet,

and pentad) are presented in Fig. II through 12, with the appropriate terminology

and physical parameters identified.

Rupe Factor/Rupe Number

The best example of an injector correlating parameter for mixing criteria is the

Rupe Factor, or Rupe Number, developed for use on unlike impinging doublets elements. This basic expression (Eq. 2) primarily was developed in the '50s by its

namesake, Jack Rupe of Jet Propulsion Lab (JPL). He ran a great number of cold

flow mixing tests and conducted related hot fire experiments. Applying stream

momentum and diameter ratios, he developed an expression, since referred to as

the "Rupe Factor," which indicated the best mixing when it equalled unity. This

parameter also can be expressed as the diameter ratio over the momentum ratio.

Since this expression is a ratio, the mathematical range of this factor from zero

to one is the same as from one to infinity, which is difficult to interpret. For

this reason, the expression has been revised to the "Rupe Number" (Eq. 3), which

has a total range from zero to one and an optimum value of 0.5.

This expression has been utilized widely for sizing of unlike doublets and has

demonstrated good correlation over a wide range of conditions. This does not

mean that a Rupe number of O.SB reflects a certain quantitative level of mixing

efficiency, but that in sizing an element for a given design, mixing, for most

cases, optimizes very near the 0.5 value.

Momentum Ratio

i

Other element types have been analyzed in a slmilar manner as the Rupe Number,

and modified momentum/diameter relationship expressions have been derived for

triplet and pentad impinging element patterns. These parameters are based on

more limited cold flow data and virtually no hot fire data, and should be used

more cautiously in universal application than the doublet expressions.

RI/RD85- 3i2

A-61
TABLE 5. INJECTOR ELEMENT CORRELATING PARAMETERS

I_o u_l

(,o

r_

PA_ETER COMMON APPLICATION

• MOMENTUM RATIO IMPINGING ELEMENTS

• RUPE FACTOR

• RUPE NUMBER

ELVERUM - MOREY

CRITERIA

• PENETRATION FACTOR

• VELOCITY HEAD

RATIO

m COAX PARAMETER

UNLIKE IMPINGING DOUBLET

ELEMENTS, LIQUID/LIQUID

IMPINGING TRIPLET ELEMENTS,

LIQUID/LIQUID

IMPINGING PENTAD ELEMENTS,

)LIQUID/LIQUID

IMPINGING ELEMENTS, GAS/

LIQUID, CENTRAL GAS

STREAM

GAS/LIQUID, LIQUID/LIQUID

IMPINGING AND COAXIAL

ELEMENTS

COAXIAL CONCENTRIC TUBE

ELEMENTS, GAS/LIQUID

• 2

PFU AFU (mox I

%x Aox _mfuI

_FU VFU Dox

mox Vox DFU

I + RUPE FACTOR

2 1.75

m •

(_in) /Pin _(Ain I

EXPRESSION OPTIMUM VALUE. LEGEND

(I)

(z)

(3)

(4)

2 1.25

(m:ou,l (°,n (5)

I i/o o1lo,o

_',n l_%ut / _out_ '

(7)

NONE

O.S

0.66

2.75

0.5

VARIABLES

= MASS FLOW RATE, KglSEC

V = VELOCITY, MISEC

A = ELEMENT DISCH FLOW

AREA, M2

O = ELEME_NT HOLE DIA, M

p = PROPELLANT, DENSITY,

Kg/M 3

MR = MASS MIXTURE RATIO, OXID/FUEL

O = VERTICAL ANGLE BEII_EEN INNER

AND OUTER STREAMS, DEGREES

SUBSCRIPTS

ox = OXIDIZER

FU = FUEL

g= GAS

L = LIQUID

in = INSIDE OR CENTRAL ORIFICE

)ut= OUTER ORIFICES

(e)

MR VL

1

( (::,, (
_i LIQUID

GAS

_s ©__ =s!.,_c,=ss o,_.T,, <cu. O,..T,,,

_ OUTER TUBE I.D. (FUEL SLEEVE)

Figure II. Coaxlal Concentrlc Element

RI/RD85-312

A-63

mm'

..... [ .... :
t c:_

i

f_D

LENGTH J /_,

_'_/_k_ ,,_/_ 2j0_ 0= IMPINGEMENT ANGLE

FUEL-OXID-FUEL

FUEL

FUEL-O)(ID-FUEL

FLIEL

Figure 12. Impinging Triplet and Pentad Elements

( ( (
As previously mentioned, the primary root for almost all impinging element mixing

parameters Is the momentum ratio. As a general rule, the momentum ratio always

Is expressed as the oxidizer total momentum over the fuel total momentum regardless of the number or placement of oxidizer streams relative to fuel streams

within the element. Relating this ratio to the values available to the designer,

we have the form of momentum ratio as shown In Table 5, (Eq. I). There is no

design optimum for this parameter and, again, thls Is a ratio with theoretical

values from zero to infinity, where values over one indicate that the oxidizer

has higher momentum than the fuel.

Elverum-Morey Factor

The equivalent of the Rupe Factor for triplet and pentad elements was developed

by Rupe's colleages, Elverum and Morey, and is based also on momentum/dlameter

(area) relationships as shown in Table 5, (Eq. 4). For the triplet element, with

two outer angled streams and a central axial stream, the relationships are set as

inner and outer streams rather than oxidizer and fuel streams, since both

fuel-oxldlzer-fuel and oxidlzer-fuel-oxldlzer triplets are in general use.

For llquld/llquld triplets, within the range of study by Elverum and Morey, the

optimum value for this expression was 0.66. Triplet injectors have been used

most commonly for hypergollc storable propellants, and use of the Elverum-Morey

Factor has been successful under these conditions. For the nominal mixture ratio

of ]lquld oxygen and liquid hydrocarbons*, very little data has been available.

A modified Elverum-Morey expression, Table 5 (Eq. 5), was designed for pentads

and has a purported optimum value of 2.75.

*The typical mixture ratio for storable propellant combinations, such as NTO/MMH

or UDMH/IRFNA, is between 1.5 and 2.5 ox/fu for main injector operation. The

mixture ratio for llquld/llquld LOX/hydrocarbon propellants, i.e., RP-I/LOX, is

optimum near 2.8 ox/fu for main injector maximum Isp, and near 0.4 ox/fu for

fuel-rlch preburner (turbine drive combustor) applications.

RI/RD85- 312

A-65
Penetration Factor

This parameter has been developed for gas/liquid triplet Injectors where two

liquld streams impinge on a central gas stream. It relates the predicted

penetration of the liquid streams to the central gas flow. Optimum mixing Is

predlcted tf the liquids barely penetrate to the Center, with liquid droplets

belng sheared off and entrained by the gas flow on the way. The penetration factor Is presented In Table 5, (Eq. 6).

A value of 0.5 Is the theoretlcal optimum. Lower numbers lnfer that the llquld

ls being deflected away by the gas or ls not fully penetrating the gas stream.

Over penetration, on the other hand, produces a liquid fan within the gas flow,

which also reduces the uniformity of gas/llquld mixing. This factor was created

from a comb_nation of analysis and cold flow experiments, and hot fire data

appears to support the basic premise. Pentads and other lmptnglng patterns wlth

liquid streams Impinging on a central gas core also would be expected to correlate with some form of the penetration parameter. However, data Is llmited for

these applications.

The use of this factor for the reverse case of gas streams Impinging on a central

liquld, or any other extremes in the density relationships, ls questionable.

Trlplets with the gaseous reactant In the outer streams have been used In numerous cases, but there Is little data on any correlatlng parameters. Some

llmlted lnformatlon suggests that hlgh levels of gas to 11quid momentum ratio are

beneficial to the m_xlng process In lmplng_ng element injectors.

Velocity Head Ratio

Another parameter that does not have a stated optimum value Is the velocity head

ratio shown In Table 5, (Eq. 7). This roughly relates to the very practical

RI/RD85-312

A-66
consideration of "Delta P" ratio, or pressure drop ratio. The usual starting

point in an injector design is based on the desired level of pressure drop at the

design flowrates. Isolation between feed system and chamber pressure disturbances generally dictates a desire for a high level of injection orifice pressure

drop, and system pressure limitations would like a low pressure drop. A compromise solution usually results in an injector delta P of about 15 to 20% of chamber pressure, and an initial starting point would be for both oxidizer and fuel

systems to be roughly the same value. Therefore, an injector design that has

velocity head ratios significantly distant from 1 would require some compensation in design approach (i.e., supplementary orifices, etc.).

r

As mentioned previously, there is no theoretical optimum for the velocity head

ratio, but the values close to 1 are desirable for system integration. Many

times, sizing the injection orifices to optimize one of the other parameters wtll

result in an unacceptable level of velocity head ratio. For this reason, the

velocity head ratio should be computed at the same time as the other parameters,

and evaluated and adjusted concurrently.

Coaxial Parameter

The gas/llquld coaxial concentric tube injector element has had wide, successful

usage for hydrogen/oxygen combustion. Cold flow and hot flre experience wlth

thls element still has not provided a good correlating parameter. In this element, typical design practice has been to provide a Iow-veloclty central liquid

stream (liquid oxygen) sheathed by a hlgh-veloclty gas flow (gaseous hydrogen or

fuel-rich preburner gases) as shown in Flg. II. Mixing and atomization are provided primarily by the shear forces between gas and liquid and by the momentum of

the expanding gases.

Recessing the liquid stream upstream of the exit plane of the outer (gas) stream

is popularly held to increase both atomization and mixing. Cold flow testing has

not establlshed a strong correlation wlth this practice, although hot fire

results generally reflect a performance increase that usually is accompanied by

an increase in face heating.

RI/RD85-312

A-67
Increasing the gas velocity (relative to the liquid velocity) generally improves

mixing. This design approach should not be employed blindly, since some references suggest that mixing can be impacted adversely by velocity ratios that are

too high. This would tend to suggest that some correlating parameter for optlmlzatlon may be possible. Very high gas velocity apparently can reverse the gas

liquid relationship, "blowing out" the center of the spray and dispersing excess

liquid to the outside of the spray cone.

A review of existing data, as a part of this effort, indicates trends that may be

useful for providing a general optimizing expression for the coax element. Falk

and Nurlck of Rocketdyne (NASA CR-72703 R-B361) have suggested the coaxial parameter presented in Table 5, (Eq. 8). However, no optimum value of this parameter

has been established. One of the objectives of the remainder of this program is

the establishment of a coaxial element mixing parameter.

MIXING TEST METHODS

Liquld/Liquld Mixing Test Methods

The liquld/llquld testing for mixing efficiency is relatively easy and low cost,

if facilities are available. The procedure for llquld/liquld mixing utilizes a

grld-llke sample device, which ducts the individual position captured liquid into

an appropriate sample container (Fig. 13). This technique utilizes two

immiscible liquids as propellant slmulants, typically water and a high-denslty,

low-vapor pressure solvent such as l,l,l-trlchloroethane. The fluids collected

in the sample tubes separate by the variation of density and their quantities in

each tube are measured (Fig. 14). Typically, the sample grid represents hundreds

of data points, and a computer data reduction process is required to provide

meaningful quantitative data.

Different fluid combinations have been employed for liquld/liquid mixing in an

effort to better match injected reactant conditions, while addressing concerns

for toxicity, flammability, and general questions of safety, convenience, and

cost. Other solvents used for these purposes have included many of the lower

RI/RD85-312

A-68
ORIGINAL PAGE I'S

_ OUALn'Y

m

a. Square Tube Assembly

l_ _._i__ _i  ̧ ̧ ̧_

b. Test Sample

Figure 13. Liquld/Ltquld Mixing

%..,_

RI/RD85-312

A-69
PROPELLANT

SIMULANT

ELEMENT

FLOW

M IST

STATIC

I

PROPELLANT

StMULANT

INJECTOR

MODEL

TRANSPARENT

CHAMBER

PROBE

(PITOT)

POSITION

PROBE

VELOCITY

HEAD

LIQUID

SAMPLING

CONTAINER

Figure 14. Cold Flow Gas/Liquld Mixing Measurement System

RI/RD85-31 2

A-70
vapor pressure "freon" compounds, perchlorethylene (a dry-cleaning solvent), as

well as fuel-type hydrocarbon liquids. At least one past program at Rocketdyne

utilized a water/brine system, wlth the mixture ratio of the sample determined by

an electric salinity meter. Data acquisition using thls method was significantly

slower than the l_lscible fluid method, and accuracy was poor in the low mass

flow outer zones.

v

Gas-Llquld Mixing Test Methods

Gas-Llquld mixing tests are significantly more tlme consuming than the liquidliquid mixing, which probably is the reason that gas-llquld data is more

limited. A gas-llquld mixing measurement system has been utilized extensively at

Rocketdyne for hydrogen/llquld oxygen concentric elements (with the gas annulus

surrounding the liquid core). The schematic of the process Is shown in Fig. 15.

The sample element is installed at the "head end" of a transparent, pressurized

chamber, wlth a traversable probe mounted at the desired sampling plane. Water

typically is used for the liquid oxygen slmulant and a nonreactive gas simulates

the hydrogen fuel (or hydrogen-rlch hot gas in a staged combustion cycle).

Typically, the gas used Is nitrogen, sometimes diluted wlth helium to provide a

desired density. Gas density Is controlled by tank back pressure, and the mixture of gases supplied. A "base bleed" gas usually is supplied through the face

around the injection element to minimize reclrculatlon from the injected flow,

and to simulate partially the axial gas flow present in a combustion chamber. A

tracer gas (frequently oxygen) is included in this base bleed flow to allow this

local gas flow to be measured and extracted mathematically from the measured

element gas flow In each sample.

The sample Is extracted from the gas-llquld element flowfleld by the use of a

sharp edge probe that can be positioned In the desired sample area. The liquid

spray in the sample zone is collected physically by the opening in this probe,

and accumulated in a sample container over a measured time period. The gas flow

flux In the sample zone Is determined from the relationship between total and

static pressure (corrected for the liquid In the two-phase flow). The gas measurement may require a second correction for the entrained "base bleed" flow, and

the data for this correction is obtained from an "on-llne" gas analysis technique.

RI/RD85-31 2

A-Tl
As might be deducted from the preceding description, each sample requires a

sufficient tlme to stabilize the required readings and collect the llquld. When

compared to the hundreds of sample points simultaneously obtained In the

liquld-ltquld testing, the increase in test time for gas-liquid testing Is

readily apparent. Testing with concentric elements permits a reasonable

assumption of circular symmetry, allowing a reduced number of required sample

measurements. However, the more complex "fan" shapes of gas-liquid triplets and

pentads require careful study of the sample locations, and require more sample

points than for a co-ax test. Previous work with triplets and pentads In a

gas/fluldized solid system, Ref. 30, and triplets in a gas-gas system have

Indicated the shape of expected mass distribution, and show that numerous sample

points are required to characterize these element types.

COLD FLOW MIXING DATA REDUCTION

The data reduction procedures for the llquld-liquld and the gas-llquld cold flow

mixing tests are very similar. As In the testing Itself, the data reduction for

the 11quid-llquld testing is a blt more straight forward. The total sample grid

usually encompasses all the injected flow, and the grld openings usually have no

open spaces between them. Therefore, the collected totals should equal the

injected totals, thus providing a good cross-check on the data. Thls Is the

first factor computed In gas/llquld mixing tests, and is referred to as the

"collection efficiency."

Collection Efficiency

To calculate the collection efficiency of the test system, fluid input values are

compared with fluid collected values. The input values of mass flowrate is

frequently calculated theoretically by the Injector Pressure Drop Equation (g),

based on previous cold flow resistance calibration of the test model:

V

_Input = _ NO2 Cd (2pAP)

I/2

(9)

RI/RD85-31 2

A-72

wj
where

= mass flowrate

N = total number of oxidizer or fuel holes

D = diameter of orifice

Cd = dimensionless discharge coefflclent as determined

from the calibration flow test

p = density of simulant

AP = injector pressure drop

If direct flow measurement capability exists in the cold flow mixing facility,

the values from these measurements are used.

The collected values of mass flowrates are calculated from the test data; summing all of the individual sample measurements:

L_

collected = P Q

t (lO)

where:

p = density

Q = local corrected sample volume

t = collection time in seconds

Collectlon efficiency of the system Is calculated then by:

collected

. - (11)

col W input

where a value of "l" represents perfect collection efficiency. Large deviations

in the collection efficiency would indicate problems in the system or the data

for the testing. Unfortunately, collection efficiency rarely is included in

reports of mixing tests and in some cases may not even be calculated. Liquid/

liquid mixing testing is relatively simple and collectlon efflclency generally

_.__" RI/RD85-312 - A-73
is not needed or obtained. However, the much greater complexity of gas/liquid

testing requires the "check" on test methods and procedures that collection

efficiency provides.

Mixing Efficiency

The most meanlngful expression for assessing mixing efficiency is the Em (E-sub

m) value proposed several years ago by Jack Rupe at JPL. This is an expression

for the mass mixture ratio distribution of the samples based purely on the relatlonship of the samples to the overall mixture ratio with no regard to such factors as theoretical stolchlometry, etc.

This value is computed as a mass weighted summation of the mixing errors in all

the samples. In practice, it is computed as a summation of decrements based on

how far the mixture ratio of each sample deviates from the overall mixture ratio,

and weighted by the mass fraction of each of these samples. The range of this

expression is from zero to 100%, with 100% indicating all samples are the same

mixture ratlo, and zero indicating the samples are all one component or the other.

The nominal form for computation of E is expressed by:

m

R_Rs b R-Rsa ]

Em = lO0 l- Z MFsb _ + X MFsa R---T (12)

where

Em = mixing efficiency from 0 to 100%

R = overall mixture ratio as expressed by weight flow

oxidlzer/welght flow total

Rsb = mixture ratio of sample below overall mixture ratio

MFsb = mass fraction of sample below overall mixture ratio

MFsa = mass fraction of sample above overall mixture ratio

Rsa = mixture ratio of sample above overall mixture ratio

RI/RD85-312

A-74
Each local sample that Is not at the overall mixture ratio thus provides a mixing

efficiency decrement proportional to how far it is from the nominal mixture

ratio, and what mass fraction of the total flow it represents. For example, if a

sample representing 50% of the total mass has a mixture ratio fraction of 0.35

when the overall is 0.70, the total mixing loss from this sample is

I00 (0 S 0.70 - 0.35) • 0.70 = 25% loss in mixing efficiency.

This factor is much more sensitive to mixing deficiencies than combustion

efflclency-related factors, which are "rounded off" by theoretical curves and the

relationship between test mixture ratio and stolchlometrlc mixture ratio.

Mixing Limited C-Star

A frequently used parameter to describe mixing test results is mixing limited

C-star or * mix (ETA C-star mix). This can be applied only to tests for a

c

specific reactant combination, and actually only for an assumed chamber pressure. It is a prediction of the expected hot fire C-star efficiency (assuming

total vaporization). The product of vaporization efficiency and mixing limited

C-star efficiency is the predicted combustion efficiency.

At Rocketdyne, the mixing limited C-star is computed by a single stream tube

performance model technique. The computer program is provided with a theoretical

C-star function and the theoretical C-star value (M/see) is calculated for each

sample mixture ratio. Each sample collected mass is multiplied by the sample

C-star, and these products are summed for the entire sample. This answer is

divided by the total mass collected to provide the mixing limited C-star.

C_ x Mass I + C* 2 x Mass 2 .... + C_ x Mass N

Mixing Limited C-star = Total Sample Mass (13)

E-p:

RI/RD85-312

A-75
The mixing limited C-star efficiency then is determined by comparing this value

to the theoretical C-star for the overall mixture ratio:

mlxlnq limited C*

nc* = theoretical C* (14)

A C-star efficiency of one indicates that at uniform mixture ratio, E = 1,

m

mixing llmlted C-star ls equal to theoretical C-star. This parameter Is used to

make a rough estimate of performance potential for given operating conditions of

certain mixture ratio and mixing efficiency.

REVIEW OF EXISTING DATA

An extenslve literature search was conducted on past experience In determining

and evaluating mIxlng efficiency for triplet, pentad, and coaxial elements.

Numerous document references were accessed and reviewed, and a bibliography of

the pertinent reports reviewed is presented herein. The intent of this search

prlmarlly was to find reports containing quantitative cold flow mixing test data

for these injectors.

The literature search yielded fewer reports than had been anticipated, although

several valuable references were encountered. The abundance of data Involved

11quld/llquld impinging doublets followed by llquld/llquld triplets and pentads.

Gas/llquld reports were almost entirely limited to coaxial elements and presented

little data regarding gas/liquld triplets or pentads.

The data from each report was re-reduced in order to provide a uniform basis for

comparison. In each instance, the objective was to obtain as close to raw data

as possible from the information In the report. Using a computer program

designed for this task, a table of injection parameters relating to measured performance was constructed. Information from each report thus was computed In the

same consistent fashion for best comparison of results.

As expected, many important test conditions typically were omitted from the

reports, such as the distance from the injector face to the sample plane, the

RI/RD85-31 2

A-76
relatlve size of the sample grid, and the number of sample polnts In the test

plane. For gas/ltqu%d coaxial element data there had been controversy on the use

of averaged data for sample grid points, and the reports typically did not elaborate on data reduction methods. Wlth these llmltations and Eonstraints In

m%nd, the data was analyzed and reviewed for some generalized concluslons.

)

The data was extracted from all the reports that had usable m_xlng data and has

been prepared in summary chart form (Table 6). The data has been organized by

element types and propellant condition (l.e., gas-liquid triplet, ]%quld-llqu%d

pentad, etc.). A]I of the norma]]y used injector sizing and operating parameters

are displayed (If they were available or calculable from the report information). Where a report provided information on more than one element type or propellant condition combination, It has been listed in appropriate multiple locations In the charts, with cross-reference to the other elements. These charts

are intended as a summary reference source, rather than a quick graphic comparison, and a review of data comparing s_mtlar configurations can be accomplished

with minimum confusion. Most of the data also is presented elsewhere %n this

report In graphic form, w%th mlx%ng efficiency plotted against the common

injection parameters.

Triplet - Liquid/Liquid

Two documents for liquid/llquid triplets, were found each containing significant

single element data on several configurations (Ref. 5 and 7). The data was relatively consistent and Indicated a reasonable correlation with the Eiverum Morey

Factor (Fig. 15 and 16). These plots depict the Elverum Morey Factor on a

logarithmic sca]e since this factor is computed as a ratio. In both references,

It can be stated generally that maximum mlxlng efficiency occurs near the 0.66

value for the factor. Elements with near the same orlflce diameters appear to

provide the highest maximum mixing efficiency, and multiple elements reflect

lnterelement mixing with higher average values and reduced sensitivity to the

parameters. A "reverse" triplet (two oxidizer streams on a central fuel)

(Ref. 5) appears to have a significantly different characteristic as a result of

having the high density llquld on the outside. Further evaluation of these

@

RI/RD85-31 2

A-77
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RI/RD85-312

A-78

ORIGINAL PAGE IS

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ORIGINAL PAGE, IS

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PENTAD

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RI/RD85-312

A-79

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0

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X

0

0

! I I

2 4

I I ! I ! _

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Figure 15. Mixing Efficiency of Liquld/Llquld Triplets (ITUR-80-9)

Depicting Elverum-Morey Factor Influence
O0 O0

f

PO

E

L_

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13

I I I I I I I I I ] I ! I I

.2 .4 .6 .8 1 2 4 6

ELVERUM-MORET FRCTOR

I I I

8 10

Figure 16. Mixing Efficiency of Liquid/Liquld Triplets (NASA CR-R-9270)

Deplctlng E1verum-Morey Factor Influence

( (
characteristics would appear to be warranted since liquld/llquld hydrocarbon

mixtures favor a reverse triplet configuration at main chamber mixture ratios.

Triplet - Gas/Liquld

One report was found on LOX/hydrogen work (Ref. 34) which provided data on a

llquld/gas/llquld configured element. Suprlslngly, the penetration factor,

designed for llquld/gas/llquld elements, did not produce the desired correlation

of maximum mixing efficiency (Em) at the 0.5 theoretical optimum value

(Fig. 17). Visual aids from the report depict the gas/llquld normalized mass

flux profiles for each of three cold flow tests. Figure 18 depicts representative samples of those three tests. The sample mixture ratio is equivalent to the

overall inlet mixture ratio where the dashed lines (gas) intersect the solid

lines (liquid). It can be inferred from the distribution plots that the balance

of gas and liquid was optimum at the under-penetrated condition (penetration

factor 0.4), which contributed to the maxlmum-measured mixing efficiency. At

penetration factors greater than 0.4, the gas blowout produced by the impinging

liquid Jets was visible. This contributed to'the poorer mixing efficiency noted.

lhe Elverum-Morey criteria for this element, shown in Fig. 19, did reflect a

correlation between the 0.66 optimum value and the peak mixing efficiency. In

this test, the oxldlzer-to-fuel density ratio was over 600, markedly removed from

the design application range of 1.7. These parameters bear additional testing

since there are good designs for llquld/gas/llquld elements in LOX/ hydrocarbon

gas generators and preburners.

Pentad - LiquldlLiquld

Documents obtained with mixing data for llquld/liquld pentad elements consisted

primarily of reverse configuration* element studies (Ref. 3, 4, and lO). In these

* A reverse pentad generally is considered to have the denser liquid (oxidizer)

in the outboard streams and the less dense liquid (fuel) as the centrally located

stream.

RI/RD85-312

A-83
1

l_]:_rg TRI'PLET (R-Cj315)

i

il_ ,

i_',_'!I• _' " _,_-"

.9

.8

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m

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I lr

6

I I I

8 10

Figure 17. Mlxlng Efficiency of a Gas/Llquld Trlplet (R-9315)

Deplctlng Penetration Factor Influence

( ( (
_o

oo co

u_ u_

I

c_

t_

DEFINITION OF

F LOW SECTIONS

,y11

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I

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!

1 I

SECTION 3

1.0 0 1.0

RADIAL DISTANCE, INCH

I PENETRATION FACTOR = 0.25 I

i

2

f

SECTION3/

1.0

__ IECTION 11_

0 1.0

RADIAL DISTANCE, INCH

PENETRATION FACTOR = 0.395 !

3

12

SECTION3 SECTION 11

,"",J .F,\

,'_---' I ,_...-.--',

1.0 0 1.0

RADIAL DISTANCE, INCH

I PENETRATION FACTOR = 0.62 I

Figure 18. Normalized Cold Flow Mass Flux Profiles for LOX/GH 2,

Liquld/Gas/Liquld Triplet, 2-1nch Collection Tubes (Ref. 34)
O0 O0

G_

r_o

.9

.8

E

hi .7

Z

hi

M

u •5

tA.

4_ hi •

t._

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i I I I I I I I I I I I

• 1 .2 .4 .6 .8 1 2 4

ELVERUM-MORETPRRRMETER

I I

6

! I I

8 10

Flgure 19. Mlxlng Efficiency of a Llquld/6as/Llquld Triplet (R-9315)

Depicting Elverum-Morey Factor Influence

(
F

i

!

studies, the overall level of mixing efficiency was generally good. Single

element characteristics dld not adhere to the Elverum-Morey theoretical optimum

very well for the large element tests shown in Fig. 20 (Ref. 3), although the

multi-element tests did show peak mixing efficiency near the 2.75 optimum value

for the same experimentors. This ls either a result of secondary mixing enhancement from the multlple element configuration or ls indicative of absolute size

limitations in parameter application. Other data presented in Fig. 21 and 22

indicate some small degree of correlation wlth the 2.75 optimum parameter value.

Pentad - Gas/Liquld

The volumetric unbalance realized with gas/llquld propellant combinations frequently dictates the use of pentad (four on one) elements. Wlth the gaseous

reactant on the four outside elements, this bears some resemblance to an impingIng concentric element.

? With the gaseous component of the reaction system in the center stream, the case

resembles an extension of the llquld-gas-llquld triplet where a form of the penetration factor becomes the most likely mixing parameter.

One document was located wlth gas/llquld pentad data (Ref. 31), which includes

test data for both configurations, llquld-gas-llquld and gas-llquld-gas. This

data was replotted against three different parameters, momentum ratio, ElverumMorey ratio, and penetration factor.

Both pentad configurations showed improved mixing characteristics with increased

oxidizer (llquld) momentum (Fig. 23), regardless of the orientation of the oxidlzer stream(s). This Is not understood fully since prior experience on other

programs, such as the gas/fluldlzed-solid program (Ref. 30), indicated contrary

relationships, i.e., an increase in performance with a reduction in momentum of

the central-fluldized stream with maximum performance occurring at a relatively

high gas to liquid momentum ratio.

RI/RD85-312

A-87
PENTRD (LIg/LIQ) RFRPL-TR-66-147

1

GO Oo

Co ¢-_

I

.9

.8

,7

E

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Figure 20.

ELVERUM-MOREY FRETOR

Mixing Efficiency for Liquld/Liquld Pentads

(AFRPL-TR-66-147) Depicting E1verum-Morey

Factor Influence

0

I ! I

6

t nl

8 10
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GO CO

,.0 t._

I

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ELVERUH--MORETFACTOR

4-

I I ! ! I I

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+

.I"

I ! I ! I I

4 5 8 10

Figure 21. Mixing Efflclency for Liquid/Liquid Pentads

(NASA CR-72827 R-8415) Depicting ElverumMorey Factor Influence
_0

O

I

C_J

..a

r_J

,,

.9

.8

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x ÷x

!

2

ELVERUM-MORETFACTOR

÷

÷

X

I I I I

4 6

I I I

8 10

Figure 22. Mlxlng Efflclency for Liquid/Liquid Pentads

(AFRPL-TR-66-152) Depicting Elverum-Morey

Factor Influence

( (: (
iiliiIiI ImIIImInlInmInnm/iioiiiIiinim

1

I _u

.9

.8

E

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zn_n o 6

m

i

,-,-. 5

I,I

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X

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4"

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_

+

÷ ÷

i

m

÷

I

.3

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0

• 1 o2 .4 .6 .8 1 2 4 6 8 10

HOMENTUHRFITIO

Figure 23. Mixing Efficiency for Gas/Liquld Pentads

(NASA CR-72703 R-8361) Depicting Momentum

Ratio Influence
Extrapolating the llquld-gas-llquld test data along the Elverum-Rorey curve, Fig.

24 suggests a trend toward the 2.75 optimum value for pentads, whereas the gasllquld-gas data do not obey the parameter functions. Extrapolating the penetratlon factor data for the gas-llquld-gas element may indicate a trend toward the

0.5 optimum value (Fig. 25). The llquld-gas-llquld element apparently does not

adhere to the penetration factor function.

Concentric Coaxial Element

Several report references were obtained in the literature search containing cold

flow mixing data for coaxial elements. Some of these were from the Space Shuttle

Main Engine (SSME) Program. The mixing data from these sources were plotted

against the conventional parameters applicable to coaxial injectors, namely LOX

post recess and veloclty ratlo.

In most concentric element configuration, relatively large Improvements in mixing

are anticipated as the central tube (oxidizer post) recess ls increased to one

liquld stream diameter. Data presented In Flg. 26 (Ref. 7) depict less effect

than had been expected. The curve indicates poor overall mixing efficiency (E m

= 50 to 65%) with very llttle improvement obtained as recess is increased. However, Falk and Burlck report In their studies (Ref. 19) that cup recess does

improve mixing. This conflict needs to be resolved by additional testing, especially In the areas of hydrocarbon fuels.

V

The influence of gas-to-liquid veloclty ratio on the level of mixing efficiency

ls presented in Flg. 27 and 28, deplctlng the characteristics of SSNE LOX/

hydrogen preburner and main injector elements In cold flow test. In these flgures, mixing efficiency is conslstantly high. Propellant density matching was

achieved for these tests, which also resulted In nominal matching of hot fire

(design range) velocity and momentum ratios simultaneously.

Additional tests conducted by Rocketdyne (Ref. 31 and 7) are presented in Fig. 29

and 30, respectively, depicting the effects of velocity ratio on mixing

efficiency. The latter figure shows the Influence of gas-to-liquid density ratio

RI/RD85-31 2

A-92
CIf!!rF . LLI!I _

I-'4

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Flgure 24. Hlxlng Efflclency for Gas/Liquid Pentads

(NASA 6R-72703 R-8361) Depicting ElverumHorey Factor Inf]uence
CO

I

Go

ii i , i i i i i i i iiii ii

.8.9t ,.x I

E .7 ÷ ÷

d i "

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.1

0

• 1 .2 .4 .6 .8 1 2 4 6 8 ]a

PEIIq'RgT ION FgL'TOR

Figure 25. Nlxlng Efficiency for Gas/Liquid Pentads

(NASA CR-72703 R-8361) Depicting Penetration

Factor Influence

( ( (
u_ QO

T

°0

,8

E .7

>_-

C.) ,6

Z

i,i

m

m

m,

Z

m

X J

m

J

.|

mL.

t t

• ,I ,at

I I

J .4

I t I. t I L ! t t l 1 t I I a

,ll ,1 ,7 dl ,0 I L! I,| 1,8 1,4 III 1,1 1,7 fill J,ll II

RECESS / L.[O OIA

Figure 26. Mixing Efficiency for Coaxial Element (NASA CR-R-9270)

Depicting LOX Post Recess Influence
30

I C3

O0

0",

I

OJ

E

i.a.i

>_-

(.J

Z

LL=

m

L.Ii.:..

I.,t.I

Z

m

x

m

1

.9

.8

.7

.6

.5

.4

.3

.2

.1

a

+ ÷

- + t_x+# + +_¥

X

1

+ 55ME OPB ELEMENT I

x 5SME FPB ELEMENT I

I

! i I ! I _ _ II I l i

2 4 6 8 10 20 40

F_5 VELOCITT / LIQUID VELOCITT

I 1 ! I

60 80 100

Figure 27. Mixing Efficiency for Coaxlal Elements (IL PT 73-30)

Depicting Velocity Ratio Influence

( ( (
C!I,,

_O

mO

._j ¢jm

T

COAXIAL [GAS/LIQ] [L PT 73-37

X

GAS

2 4 6 8 ]O

VELOCITY / LIQUID VELOCITY

t_

4B 5B 80 108

Figure 28. Mixing Efficiency for Coaxial Elements (IL PT 73-3?)

Depicting Velocity Ratio Influence
1

_o Go

Go ul

I

_o

Do

E

i,I

>_-

CD

Z

m

0

m

i,

i,i

CD

Z

m

X

m

.9

.8

.7

.6

.5

.4

.3

0

J

d

1

+ A-LIO 033/ R-GAS .067 I

x :3o2/R-GAs i 89

o A-LIO .15/ A-GAS .

I I 1, i ! ! I n I

2 4 6 8 IO

GR5

AREAS (IN 2)

VELOCITT / LI_JID

+

÷

÷

1 I I

20 40

VELOCITT

X

0

n

+

X

+ +

+

X

÷

n

I ! I !

60 80 I00

Figure 29. Mixing Efficiency for Coaxlal Elements

(NASA CR-72703 R-8361) Oeplctlng

Velocity Ratio Influence

( (
I

COAX IRL (GRSIL IQ) NASA CA-R-9270

E

LLI

>:

(.J

Z

::o L.U

p-W m

"- (.j

uo co Li.

,,D u_ LL.

oJ

Z

X

.7

.6

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Ci

4"

0

0

t

8

+

- X

o

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.01:1433 DENSITY RFITIO

.01020 DEI_ITT RATIO

.02400 DENSITT RATIO

.00216 DE.NSITT RRTIO

10

! I I I I I I I 1

20 40 60 80 I00

F:,FISVELOCITT / LIQUID VELOCITT

X

200

1 I I

4OO

! I I !

600 8001000

Figure 30. Mixing Efficiency for Coaxial Elements

(NASA CA-R-9270) Depicting Velocity

Ratio for Varying Density Ratios
as well, and clearly indicates that higher gas-to-fuel density ratios produce

higher mixing efficlencies for a given veloclty ratio. This relationship

strongly suggests that a velocity-density product, such as momentum ratio, will

not peak at an optimum value, but wlll approach ideal mixing as the gas momentum

continuously increases. For this reason, an alternate parameter (Table 5. Eq. 8)

has been considered in an effort to characterize the data wlth a single

expression. The coax parameter (Ref. 19) was applied to the SSME preburner and

main injector data as shown in Fig. 31 and 32). Because of the high overall

mixing efficiency of that data, no predominant trends were evident.

1

LITERATURE SURVEY CONCLUSIONS AND RECOMMENDATIONS

As a result of the literature review and data examination, most of the initial

Impresslons regarding the state of the cold flow data have been confirmed. Large

discrepancies exist in test results noted between the various experimenters, and

there does not appear to be any proven correlatlng parameters for coaxial element

mixing efficiency. In general, the available data is insufficient to confidently

confirm or establish the optimum value of the correlating parameters for impinging elements.w

Although the gas/llquld triplet element has significant potentlal for future

llquld-oxygen/gaseous-hydrocarbon propulsion systems, very little quantitative

data exists to either support design calculations or provide correlating

expressions for combustion modeling. Most hydrocarbons considered for advanced

booster applications will be delivered to the injectors as warm or hot gas with

densities relatively high as compared to hydrogen or combustion gases used in

current concentric element injectors. This higher density favors impinging

elements rather than the concentric element. The gas annulus gap required for

the denser fuels in a coaxial element injector may approach small absolute values

that ultimately result in poor concentricity and element contamination problems.

Greater emphasis should be placed on obtaining mixing data on gas/liquld

impinging (especially triplet) elements.

RI/RD85- 312

A-lO0
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r_

i I"--

(_ r-.

,-,,.

14q

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IIHI lillllillll q Iin IIIIIIIP IEIIIIIIIIIIIMIIIlI!I II imlmllnlllliiiNIimllllNillmlll iii iiiilIIIlU|
;0

•-J O0

C_ _n

E

>.-

Z

U.l

m

U_

tll

g

m

x

m

.9

.8

.7

.6

.5

.4

.3

.2

.1

0

J

÷

X

1000

i i

÷

SSNE OPB ELEBENT

5SNE FPB ELENENT

I I 1 I

2000 4000

÷

÷

X

X +

I i I I f

600_OODDO00

C0 AX P A RAMETER

i ! I I I I

40000600_

Figure 32. Rlxlng Efficiency for Coaxial Elements (IL PT 73-30)

Depicting Falk and Burlck Parameter Influence

(i ( (
REFERENCES

1. Murthy, S.N.B.: "Turbulent Mixing In Nonreactive and Reactive Flows," Plenum

Purdue University, W. Lafayette, IN, 1974

2. Evans, David D., Howard B. Stanford, Robert W. Riebling: 'The Effect of

Injector-Element Scale on the Mixing and Combustion of Nitrogen Tetroxlde -

Hydraztne Propellants," (Technical Report 32-1178), Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA, 1 November 1967

3. Dickerson, R., K. Tate, N. Barsk: "Correlation of Spray Injector Parameters

with Rocket Engine Performance," Rocketdyne, Canoga Park, CA, June 1968

4. Study of Droplet Effects on Steady-State Combustion (AFRPL-TR-66-152-VOL 1),

Volume l: Measured Spray Parameter Analysis and Performance Correlation,

Rocketdyne, Canoga Park, CA, August 1966

5. Cheung, T.T., V. W. Jaqua: Cold Flow Investigation of Unlike Triplet Elements for Liquid Oxygen/Liquid Hydrocarbon Injectors (ITUR 80-9), 8 February

1980

6. 3oshl, P., A. Jakubowskl, J. Schetz: Effect of Injector Geometry on Penetration, Spread, and Structure of a Liquid Jet Injected Normal to a Supersonic

Air Stream, Virginia Polytechnic Institute, September 1973

7. McHale, R. M.: Nonclrcular Orifice Holes and Advanced Fabrication Techniques

for Liquid Rocket In_ectors (NASA CR-R-9270), Rocketdyne, Canoga Park, CA,

May 1974

8. Riebling, Robert W.: "Effect of Orifice Length-to-Diameter Ratio on Mixing

in the Spra_ From a Pair of Unlike Impinging Jets," Journal of Spacecraft and

Rockets Volume 7, No. 7, 3PL, Pasadena, CA, July 1970

9. Riebllng, Robert W.: "Criteria for Optimum Propellant Mixlng.ln Implnqlng -

Jet Injection Elements," Journal of Spacecraft and Rockets, Volume 4, No. 6,

JPL, Pasadena, CA, June 1967

lO. McFarland, B. L., V. Jaqua: Space Storable Thruster Investigation (NASA

CR-7287, R-8415), Rocketdyne, Canoga Park, CA, lO November 1971

II. Nagal, C. K., R. N. Gurnltz, S. D. Clapp: Cold Flow Optimization of Gaseous

Oxygen/Gaseous Hydrogen Injectors for the Space Shuttle APS Thruster AIAA/SAE

7th Propulsion Joint Specialist Conference, Rocketdyne, Canoga Park, CA, June

1971

RI/RD85-312

A-103
12. H1nde, P. T.: Basic Aspects of Gas - Mixing and Klnetlcs In Rocket Engine

Combustion Chambers, Rocket Propuls%on Establishment, Ministry of Technology,

London, February 1979

13. Falk, A.Y.: Cold Flow Mixing Characteristics of the SSME Main and Baffle

Elements, Rocketdyne Report No. PT73-37, Canoga Park, CA, 14 February 1974

14. Paster, R. D.: Hydrogen - Oxygen APS Eng%nes, Volume II: Low Pressure

Thruster (NASA CR-120806, R-8837-2), Rocketdyne, Canoga Park, CA, February

1973

15. Nurlck, W. H., R. M. McHale: Nonc%rcular Orlflce Holes and Advanced Fabrication Techniques for Llquld Rocket Injectors (NASA CR-108570, R-8224),

Rocketdyne, Canoga Park, CA, 15 September 1970

16. Yost, M.: Preburner of Staged Combustion Rocket Englne, Rocketdyne, Canoga

Park, CA, February 1978

17. 6111, G. S.: A Qualltatlve Technique for Concentric Tube Element Optlmlzatlon, Utlllzlng the Factor (Dynam%c Head Ratlo - 1), Soclete Europeenne de

Propulsion, Vernon, France, January 1978

18. Dykema, Owen W.: Llquld Mlxlng of an Implnglng Jet Uslng Invnlsclble Llqulds

(R.R. 60-28), Rocketdyne, Canoga Park, CA, 30 November 1960

19. Falk, A. Y., R. J. Burlck: Injector Deslgn Guidelines for Gas/Liquid Propellant Systems (NASA CR-120968, R-8973-3), Rocketdyne, Canoga Park, CA, May

1973

20. Buschulte, W.: Propellant Nlxlng By Injectors and %ts Efficiency Effect,

26th German Federal Republic, International Astronautical Congress of the

International Astronautical Federation, 12 November 1975

21. Falk, A. Y.: Coaxlal Spray Atomlzatlon In Accelerating Gas Stream (NASA

CR-134825, R-9753), Rocketdyne, Canoga Park, CA, June 1975

22. Hoehn, F. W., J. H. Rupe, J. G. Sotter: Liquid-Phase Mixing of Blpropellant

Doublets (Technical Report 32-1546), 15 February 1972

23. Trotter, C.: Swlrllng Hydrogen Jet N%xlng With a Coaxial Atrstream, Department of Chemlcal Englneerlng and Fuel Technology, Sheffield, England, August

1970

24. Nurick, W. H.: "Analysis of Sprays from Rocket Engine Injectors," Journal of

Spacecraft and Rockets, Volume 8, No. 7, Rocketdyne, Canoga Park, CA, July

1971

V

V

RI/RD85-312

A-104
m_

25. Burlck, R. J.: "Atomization and Mixing Characteristics of Gas/Liquid Coaxial

Injector Elements," Journal of Spacecraft and Rockets, Volume 9, No. 5,

Rocketdyne, Canoga Park, CA, May 1972

26. Frlcke, Hans O., Charles J. Schorr: "Measurement of Gaseous Mixing Downstream of Coaxial And Adjacent Orifices," Journal of Spacecraft and Rockets,

Volume 9, No. 8, Bell Aerospace Company, Buffalo, NY, August 1972

27. Falk, A. Y.: Space Storable Propellant Performance Gas/Liquid Like-Doublet

Injector Characterization (NASA CR-120935, R-8973-1), Rocketdyne, Canoga

Park, CA, October 1972

28. Analysis of Rocket Engine Injection/Combustion Processes Technical Proposal

R-g668P-1, Advanced Programs, Rocketdyne, Canoga Park, CA, 5 March 1975

29. Mehegan, P. F.: Two-Stage Blpropellant Injection System Studies (NASA

CR-72303, R-7199), Rocketdyne, Canoga Park, CA, 30 August 1967

30. Jaqua, V., T. Yu: Coal Combustion For MHD Cold Flow Mixing Studies

(ITR-76-Oll-C), Rocketdyne, Canoga Park, CA, 17 January 1977

31. Mehegan, Campbell, Scheverman: Investigation of Gas-Augmented Injectors

(NASA CR-72703 R-8361), Rocketdyne, Canoga Park, CA, September 1970

32. Huebner, A. W.: High-Pressure LOX/Hydrocarbon Preburners and Gas Generators,

(Final Report) Contract NAS8-332431, Rocketdyne, Canoga Park, CA, April 1981

33. Wheeler, D., and F. Kirby: High-Pressure LOX/Methane Injector Program

(RI/RD79-278) Contract NAS8-33206

34. Nurlck, W. H.: "Experimental Investigation of Combustor Effects on Rocket

Thruster Chamber Performance," (RI/RD 9315), September 1973

a

RI/RD85-31 2

A-105
V

V _
APPENDIX B

BASIC ATOMIZATION LITERATURE REVIEW

INTRODUCTION

Available information on liquid atomization by rocket englne-type injection elements was summarized in Appendix A, in which the literature on atomization by

llke-doublet, triplet, pentad, and coaxial injection elements was summarized,

discussed, and assessed. The general conclusions of that summary were that

reported atomization data are largely empirical and ad hoc, only qualitatively

understood, and of little general validity or utility.

Part II of the review, reported herein, covers the literature on the more basic

or theoretical aspects of liquid atomization. This effort is primarily directed

toward studies related to droplet deformation, drag, and breakup, as these processes tend to influence the ultimate size and motion of droplets and are of

great importance in efforts to model sprays. Some of the more basic and general

atomization studies for alrblast atomizers are also included. The discussions of

droplet distribution functions, definitions of average droplet diameters in

sprays, droplet-slze measurement techniques, and problems of spatial versus temporal droplet distributions which are presented in Appendix A of this atomization

literature review (Ref. B-22, also see Ref. B-8g) are, of course, equally applicable to this alrblast atomizer research.

The importance of the atomization process, particularly in combustion appllcatlons, has resulted in the publication of hundreds of papers and reviews concerned with various aspects of this subject. A selection of these studies,

representing classical and current procedures, results, and theories related to

liquid atomization, are briefly summarized in this report. This summary, together

with that in Ref. B-22, provide a complete description of the state of the art of

atomization as it applies to liquid rocket engines. It should serve as a useful

reference to those familiar with this area and as a basic introduction for those

entering this field of study.

RI/RD85-312

B-1
DISCUSSION

Conversion of a volume of liquid to a spray of droplets can be accomplished by a

variety of methods, most of which function by the imposition of a high relative

velocity between the liquid and its surrounding gas. In "pressure" atomizers, a'

hlgh-veloclty Jet of liquid Is discharged through an orifice Into a Iow-veloclty

gas; In "alrblast" atomizers, a Iow-veloclty liquid is exposed to a hlgh-veloclty

gas stream. Various mechanical means (e.g., impingement of two or more liquid

Jets on each other, impingement of a Jet on a solid surface, or added swirl to

the liquid and/or gas) are frequently employed to augment or enhance the rate or

extent of atomization.

Thls review wlll be concerned primarily wlth the breakup of the large drops Inltially formed In the atomization process. The discussion wlll be presented In

terms of four interrelated aspects of the process:

.

2.

3.

4.

Criteria and requisite tlme for breakup of liquid drops

Drop deformation and drag coefficient as functions of time

Sizes of droplets formed in breakup of large drops

Effects of system parameters on the atomization process

v

Criteria and Requisite Time for Breakup of Liquid Drops

If a large liquid drop is exposed to a gas whose relative velocity is sufficiently hlgh to overcome the restoring force of its surface tension, the drop

wlll disintegrate into a cloud of daughter droplets (secondary breakup). Two

basically different modes of drop breakup have been observed, the "bag" type and

the "shear" (or "stripping") type; these are described in the section of this

report entitled "Droplet Deformation and Drag Coefficient".

Early Drop Breakup Studies

General reviews of the early work (to about 1965) on liquid particle breakup in

gas streams have been published by Forsnes (Ref. B-24), Lapple et al (Ref. B-57),

and Luna (Ref. B-61).

RI/RD85-312

B-2
A theoretical analysls of drop breakup was presented by Hlnze (Ref. B-37). He

described the deformation of a liquid sphere caused by the normal pressure distribution resulting from a flow of gas over its surface. Closed-form solutions

of the llnearlzed hydrodynamic equations were obtained for very high and very low

liquid viscosity. Hlnze also postulated the existence of a critical Weber number

as a criterion for drop breakup (Ref. B-36):

We = PG dL _21OL (B-l)

m

Where PG Is the gas density, d L is the original drop diameter, AV is the

relative gas/llquld velocity, and _L is the surface tension. An analytical

derivation of a criterion for the critical breakup condition of drops of low and

high viscosity exposed to a gradually increasing gas flow and for hlgh-vlscoslty

drops suddenly exposed to a constant-veloclty gas flow gave:

(6/r)ma x = 0.095 (We)ma x

(B-Z)

where 6 Is the radial deformation of the drop and r is the original drop

radius. For Iow-vlscoslty drops suddenly exposed to a constant-veloclty gas

flow,

(a/r)ma x = 0.17 (We)ma x

(B-3)

The analysis also yielded equations for the breakup time of drops suddenly

exposed to a constant flow:

<:lox] 0.5

(low viscosity) (B-4)

~ I0 (high viscosity) (B-S)

tb ~ PG _2 max

where PL and UL are the liquid density and viscosity, respectively.

RI/RD85-312

B-3
The llmlted experimental data then available for the case of gradually increasing

gas velocity (Ref. 8-66) indicated the following approximate values:

(a/r)crlt = 1 (We)crlt = 10 (low viscosity)

(6/r)crlt _ 2 (We)crlt _ 20 (high viscosity)

The difference between the two cases was ascribed to the extra time delay In drop

deformation resulting from high viscosity, which gives the drop time to accelerate and thus lowers the relative velocity below the critical value.

Note that for low viscosity llqulds, Eq. B-4 reduces to

tb -N (dL/2 a-V) (pL/PG)0"5 (B-6)

Thls Is the form used by subsequent investigators, although some experimental

values were up to ten times longer than predicted (Ref. B-13 and B-?l).

Taylor (Ref. B-90) presented a crude boundary layer analysls of the shear breakup

mode. Boundary layer stripping results from tangential friction, which was not

considered by Hinze (Ref. B-36), who assumed zero tangential stress. Taylor

solved for the boundary layers of both gas and liquid, assuming two-dlmensional

uniform flow, and obtained the velocity profiles of both fluids. The rate of

loss of liquid from a drop by tangential drag is given by:

dV = -K d2 vL_] (B-7)

where V Is the drop volume (or mass), K is a constant on the order of unity, d is

the drop diameter, u L Is the liquid kinematic viscosity, and b Is the boundary layer thickness.

RI/RD85-312

B-4
From experimental data, Taylor established a value of 2.7 for the critical Weber

number in steady, Iong-duratlon gas flow, where the Weber number was defined as

We = pg r _212, OL (B-8)

He also pointed out that the relative velocity required for drop breakup in transient gas flow is different from that in steady flow, and that

(We)crlt - transient flow

(We)crlt - steady flow

l

= _- (B-9)

as long as the drop oscillation period (0.258 rI'5

approximately two times the gas flow duration.

for water) does not exceed

Lane (Ref. B-56) studied the breakup of water drops in transient and steady

streams of air. In steady flow, bag mode breakup was observed, wlth the following correlation of the experimental data for the critical condition:

_-_2A crlt d = K = 612 (B-lO)

A theoretical analysis of the steady-flow case, treating the critical condition

as the point at which the drag force equals the surface tension, gave an expression for the drag coefficient:

Co = (16 _L)/(D PG _2) (B-ll)

With an assumed drag coefficient of 0.4, the value of K in Eq. B-lO was about

1200, twice the experimental value. The dlfferencewas ascribed to deformation

of the drop into a nonspherlcal shape before breakup. Lane observed that about

70% of the original mass of a drop undergoing bag mode breakup was in the heavy

rim; after the bag burst into a mist, the rim broke into larger droplets.

RI/RD85-312

B-S
Lane observed shear mode breakup in transient flow tests and concluded that this

L

was a combination of boundary layer stripping and the cresting of capillary waves

formed on the drop surface. He found that Taylor's estlmate of the ratio of critical velocities in transient and steady flow was fairly accurate. Lane also

concluded that the daughter droplet mass mean diameter (MMD) decreased wlth

increasing gas velocity, but only up to a limiting velocity, beyond which no further decrease occurred. The lower limit of the MMD was about 15 microns. Another

observation was that since the disintegrating drop accelerates and decreases the

relative velocity, the latter stages of breakup should produce larger droplets.

V

Following Lane's work, Hinze (Ref. B-38) modified his original Weber number criterion for drop breakup by including the effect of viscosity:

(We)crlt = C [I ÷ _ Nvl] (B-12)

where C is We

crlt

coslty and

for negllglble-vlscoslty liquids, _ is a function of vlsNvl = UL/(PG _L d)O'5 (B-13)

The critical Weber number was not only dependent on llquld viscosity but also on

the tlme-varlatlon of the relative gas/llquld velocity. It was smaller for a

suddenly applied gas velocity and larger for higher viscosity liquids. Experimental data with Iow-vlscoslty liquids gave a critical Weber number of approximately 13 for the transient case and about 22 for steady gas flow.

Gordon (Ref. B-30) supplemented the investigations of Hinze by considering the

cases of intermediate drop viscosity and surface tension. For bag-type drop

breakup, he derived a theoretically based critical diameter:

160

dcrlt °G _2 (B-I 4)

RI/RD85-312

B-6
=

For drops larger than the critical slze and with negligible velocity, the breakup

time for drops with low viscosity and surface tension is given by

2 dcrlt IPLGLI "5 tb - __ (B-I 5)

AV

For drops with higher viscosity and low surface tension

32 _L

tb = (B-16)

PG

the indicated independence of breakup tlme and drop diameter is probably unreal,

but comparison wlth then-avallable data showed that the estimated breakup times

were off by less than a factor of 2.

i

i

Morrell (Ref. B-67) postulated that the breakup mode is controlled by the action

tlme (flow duration) of the gas flow on the drop. If the action time is greater

than the natural period of the drop, the drop disintegrates by the bag mode.

When the action time is less than the oscillation period, the shear mode occurs.

From the results of experiments on the breakup of a liquid Jet by a transverse

shock wave, Morrell later concluded (Ref. B-68) that the flow duration merely

affects the extent of drop breakup.

Shock Tube Studies of Drop Breakup

Shock tubes provide a convenient method for studying the atomization of single

drops in a stream of high velocity gas. The drop, suspended from a thln wire or

released from a support and allowed to hang free, is exposed to a shock wave,

whose intensity controls the gas properties in Its wake. Hlgh-speed motion plctures of the subsequent atomization permit close examination of the details of

the process.

RI/RD85-31 2

B-7
Such studies were carried out by a number of investigators. Their results indicate that the influence of the incident shock, the internal motion of the liquid,

and the increased vaporization caused by the temperature increase across the

shock are negllglble and that the only effect of the shock is to induce a high

relative llquld/gas velocity in its wake. With these conditions, dimensional

analysis shows that the drop breakup time is a function of the Weber, Reynolds,

and Mach numbers and the gas/llquld density and viscosity ratios. The actual

functional relationships must be determined experimentally and are independent of

the particular cause of the relative gas/llquld velocity.

Hanson et al (Ref. B-33) studied the aerodynamic shattering of drops in a shock

tube, over a wlde range of liquid viscosity. In contrast with the finding of

Lane (Ref. B-56), they observed bag mode breakup wlth transient gas flow, providing evidence that the breakup mode is not a function of the type of flow, steady

or translent. Hanson et al also found that the bag mode occurred near the critical breakup velocity, and the stripping mode occurred at higher velocities. A

transitional mode was also reported (the "bag-stamen" mode), in which the bag

develops a re-entrant portion resembling the stamen of a flower. The "stamen"

increases in length and stands alone as the rim and the remaining portion of the

bag blow downstream and disintegrate. The authors suggest that Lane's expression

for the critical breakup condition (Eq. B-lO) has the correct form

A-_crit d = K (B-17)

but that the values of the constants m and K must be experimentally determined

for particular liquids and flow conditions. The effect of viscosity on the critical breakup velocity was found to be negligible for viscosities less than

approximately lO centistokes; above lO centistokes, increasing viscosity (at constant drop diameter) raises the critical velocity.

For water, the critical Weber number ranged from 3.6 (D = 600_, A-V = 84

c

ft/sec) to 6.6 (D = 120u, AVc = 239 ft/sec); for methanol, the range was

6.0 (D = 625u, _c = 60 ft/sec) to 8.4 (D = I18_, _ = 157 ft/sec).

RI/RD85-312

B-8
Viscous liquids required higher critical Weber numbers for drop breakup. At Weber

numbers slightly greater than critical, the drops disintegrated in the bag mode,

In which surface tension is important. For weber numbers substantially greater

than critical, shear-type breakup, in which surface tension is not important, was

observed. The experimental data also indicated that the critical Weber number

was not constant for liquids of about the same viscosity. Instead, it varied

inversely wlth drop diameter.

With the product of Weber and Reynolds numbers as correlating factor, the critical velocity was expressed by

PG d2 m3

(We Re)crlt - 2 a vG aVcrlt

(B-18)

whence, if PG' a, and vG are constant

1/3 (B-19)

AVcrlt _ a

which fit the test data.

I

Engel (Ref. B-21) studied the shear mode breakup of water drops in a shock tube

and gave detailed descriptions of her observations. However, the only variables

used were drop size and shock strength; liquid properties were not changed.

Rabln and Lawhead (Ref. B-78) observed both bag and shear-type breakup of burning

and nonburnlng drops of fuel in shock tubes. The breakup mechanism and the crltlcal velocity were dependent on the duration and velocity of the flow plateau

following the shock front' Their photographs showed that shear-type breakup

always occurred at gas velocities considerably hlgher than critical and that the

drops were shattered by the flow behind the shock wave, not by the shock front

itself.

V

RI/RD85-k312

B-9
A later report by these authors (Ref. B-79) summarized their experimental study

of the breakup of fuel drops by weak shock waves. Duration of the gas flow

behind the shock wave was varied by changing the iength of the pressure section

In the shock tube. Effects of gas flow velocity and duration, chamber pressure,

and liquld surface tension on the shattering of burning and nonburnlng drops were

Investigated.

The solenoidal retraction of the wire upon which the drop was suspended usually

resulted in the formation of a "primary" drop (500 to 1600 micron diameter) and a

"satellite" drop (50 to 300 micron diameter). Usually, the larger drops exhibited shear breakup, while the smaller drops exhibited bag breakup, at the same

gas velocity and flow duration. It was again verified that the drop was not

shattered by passage of the shock front; It was the flow that followed which

caused the breakup. Although no theoretical explanation of the choice of breakup

mode was presented, It was confirmed that gas velocities much higher than critical always resulted In shear breakup. The critical velocltles In these experiments were comparatively low (60 to lO0 ft/sec at one atmosphere; lO to 15 ft/sec

at 34 atmospheres). Flow durations were l.O to 2.5 msec.

At atmospheric pressure, both types of drop breakup were observed; at elevated

pressures, only the shear mode occurred. No significant differences In breakup

characteristics were seen between burning and nonburnlng drops, aside from a

slightly lower critical velocity for burning than for nonburning because of lower

surface tension in the former.

The test data could not be correlated in terms of the drop Weber number alone.

Instead, It was postulated that shear breakup (which Is more pertinent than the

bag mode In rocket engine combustor applications) occurs when the tangential

component of the aerodynamic forces on the drop Is greater than the surface tension forces and

-0.5

(We) (Re) = C (B-20)

The value of C for nonburnlng drops was found to be 0.5. For burning drops, the C

value was not constant, probably because of the unknown surface tension.

RI/RD85-312

B-lO
L

Shock tube photographic studies of 1800-mlcron RP-1 drops in shear breakup were

made by Rojec (Ref. B-BS). After exposure of a drop to high velocity gas, no

changes were observed for 20 to lO0 microseconds. At that time, small ripples,

or capillary waves, appeared on the drop surface, with wavelength inversely proportional to the gas velocity, and the downstream side of the drop was deformed

into a truncated cone. The surface waves increased in amplitude with time until

11gaments and droplets began to be shed. This shedding of daughter droplets

began from about 40 microseconds after exposure to the gas flow (at a velocity of

1250 ft/sec)to about ll5 microseconds after exposure (at 320 ft/sec).

The time required for complete breakup of the parent drops could not be determined directly because they broke off from their supports and were carried out of

the field of view. However, by extrapolation of the available data, the time for

complete breakup could be estimated: at VG = 200 ft/sec, tb = 800 microseconds; at VG = 500 ft/sec, tb = 500 microseconds.

This photographic evidence supports the assumption that shear mode atomization by

a hlgh-veloclty gas stream proceeds by the growth, cresting, and disintegration

of capillary surface waves. These capillary waves are characterized by very

small wavelengths, on the order of lO0 microns, which result in small-radlus surface curvatures. Since the pressure exerted by surface tension forces is

Inversely proportional to the radius of curvature, the llquld-gas Interfaclal

surface tension is important in the analysis of capillary wave dynamics.

An analysis for the case of plane liquid surfaces (Ref. B-64 and -65) derived an

expression For the mean diameter of the generated droplets. It was postulated

that when a capillary wave reaches its critical amplitude, it erodes to form a

ligament, from which droplets of a diameter proportional to the wave length (x)

are formed:

d = F), (B-21)

where the dimensionless parameter F was assumed to be independent of _ and was

found to be nearly independent of the fluid characteristics.

RI/RD85-312

B-ll
A similar analysis for drop atomization would be excessively complicated because

of surface curvature, divergent propagation of capillary waves from the forward

stagnation point, and the magnitude of the deformation that occurs when liquid

drops are subjected to hlgh velocity gas flows. However, an experimental study

was conducted of the rate of mass loss from liquid drops subjected to a high relatlve gas velocity (Ref. B-15 and -16). A shock tube was employed, in conjunction with hlgh-speed motion pictures from streak and framing cameras. The liquid

was RP-I (1400-mlcron drops) and the driver gas was nitrogen. Analysis of the

streak data gave the mass loss history of the drops and the diameters of the product droplets. The correlating expression was

2.8 -0.42

PL /

(B-22)

where m is the mass loss rate of the drop and Sd is the surface area of a

sphere with equivalent mass. Equation B-22 applies only to a quasi-steady-state

process.

The atomization rate of a drop is zero from the instant of drop exposure to the

gas flow until the capillary waves reach sufficient amplitude to crest. It is

therefore important to define the "drop breakup time" as the time elapsed before

the drop begins to disintegrate, the time required for the disintegration itself,

or the sum of these. Most publlshed reports are ambiguous In this regard, which

increases the difficulty of comparing different results.

Buzukov (Ref. B-5) investigated drop breakup in a shock tube. A dimensional

analysis using the Navier-Stokes equation without viscous effects gave the

breakup time of a drop as

AV

(B-23)

RI/RD85- 312

B-12
S

V

which Is Identical in form to Eq. B-6 and B-15, derived by Htnze (Ref. B-36) and

Gordon (Ref. 8-30). Buzukov also observed bag mode breakup near the critical

velocity and shear mode at higher velocities. He concluded that the Weber number

was the controlling parameter and that the breakup of capillary wave crests was

the basic mechanism of the shear mode.

Several other shock tube investigations (Ref. B-21, -50, -79, -80, -85, and -97)

have indicated the same relation for drop breakup time In the shear mode as Eq.

B-23:

N_ -- or t ~d

b~ v \PG/ b~

0.5

(B-24)

where q Is the dynamic pressure. Defining B as the gas/liquid density ratio and

Tb as a dlmenslonless droplet breakup time, [_tb/d ], then

T b _ Tb _'= constant (B-25)

This equation shows that the Weber number, Reynolds number, and gas/llquld viscosity ratio are of little importance in shear-mode breakup time. Similar results

have been obtained for liquid Jets (Ref. B-? and B-69). Reported values of Tb

are 3.5 (Ref. B-?2), 4.5 (Ref. B-50), and ranging from 4 at low subsonic gas velocity to 5.5 at sonic velocity (Ref. B-80).

Fishburn (Ref. B-23) reported the results of a boundary layer analysis of drop

undergoing constant acceleration and deformation in hlgh-veloclty gas stream in

which We >> Re0"5 He concluded that drops deform and fragment at the same

dimensionless time regardless of initial size.

Several approximate theoretical analyses (Ref. B-5, -7, -67, -80, and -97) using

varied approaches have led to expressions similar to Eq. B-24. However, an analysis (Ref. B-14) that treated shear breakup as stemming from the generation and

shedding of capillary waves did not lead to Eq. B-24 or the expected limiting

behavior, but did appear to give some agreement wlth expermental data.

RI/RD85-312

B-13
In a departure from other drop breakup theories, Wolfe and Anderson (Ref. B-97)

postulated that the breakup is a rate process. They theorized that the customary

analysls, which equates the maximum force tending to break up a drop to the surface tension force, Is valid only for small rates of stress loading and not for

shock processes nor for situations in which the stress tending to break up the

drop undergoes a change In less time than that required for the breakup. Their

approach applied kinetic theory to the breakup process, In addition to the hydrodynamic/mechanlcal aspects.

V

Aerodynamic pressure drag and aerodynamic friction drag were considered to be the

factors responsible for bag and shear drop breakup, respectively. A qualitative

theoretical derivation, using rate process theory to relate drop deformation to

these aerodynamic forces, gave the following expressions relating drop breakup

time to the gas and llquld flow parameters:

d

tb = (B-26)

(A 2 ÷ B p)O.5 - A

where

A = 16UL/d PL

B = 2/p L

k = constant (Reflects drop curvature and breakup mode).

For liquids with negligible vlscoslty and surface tension (and C D = l), Eq.

B-26 becomes

RI/RD85-312

B-14

V
Again, this is the same relationship, Eq. B-23, as that suggested by many other

investigators.

Breakup time in Eq. B-26 and B-27 is the time from inception of the aerodynamic

flow around the drop to complete disintegration of the drop. If the frictional

drag on a drop is twice the pressure drag, "as generally assumed, then the pressure factor in Eq. B-26 becomes

Pb = -2 - d (B-28)

Ps G Ks a = - d (B-29)

for the bag and shear modes, respectively. Constants Kb and Ks reflect the

effect of surface tension in each type of breakup. Experimental data indicated

that Kb = 4 and Kf = 2, and that the droplet sizes produced by the bag and

shear modes of drop breakup were essentially the same.

From capillary wave theory, Wolfe and Anderson developed the following expression

for the mass mean diameter of the droplets resulting from shear mode breakup of a

drop:

I/3

36 _L 03/2 dIvv4/31

MMD = 2 I/2 (B-30)

PG PL

This equation is stated to be valid for cases in which the dynamic pressure

forces are much greater than the viscous or surface tension forces. Agreement of

their shock tube data with both the breakup tlme and mean diameter equations was

fairly good.

RI/RD85-312

B-15
Ranger and Nlcholls (Ref. B-80) studied shear-mode droplet break at high Weber

numbers in a shock tube. Drop diameters were in the range 750 to 4400 microns.

At lower gas velocities, the drop deformed into an elllpsoidal configuration,

followed rapidly by stripping of the surface to a micromlst.

At high gas velocities, these processes occurred essentially simultaneously. The

mlcromlst was observed to follow the gas streamlines, indicating that it was composed of very small droplets.

Ranger and Nicholls derived the same breakup time equation as other investigators

(Eq. B-23). Reported values of the constant K in thls expression are:

K = 0.57

K =2.0

K = 0.20

K =I

K=I

K =I

Hinze (Ref. B-36)

Gordon (Ref. B-30)

Ranger & Nicholls (Ref. B-B0)

Wolfe & Anderson (Ref. B-97)

Clark (Ref. B-?)

Buzukov (Ref. B-5)

The value of the drop drag coefficient was estimated as CD N 3, the breakup

distance was on the order of 25 drop diameters, and the size of the daughter

droplets was estimated to be approximately lO microns.

A review published in 1967 (Ref. B-61) summarized the state of the art at that

time:

l , The best criterion for drop breakup is a critical Weber number, which is

a function of liquid viscosity and the variation of gas flow wlth time

(i.e., steady vs. transient).

0.5

2. For low vlscoslty liquids (Ohnesorge number, (_/p _ d) < I),

the critical Weber number is between l and lO.

go as high as 30.

,

For Oh > I, Wecrlt may

Drop breakup requires sufficient time of exposure to the gas stream, so

Wecrlt is a necessary but not sufficient requirement.

RI/RD85-312

B-16
. Near the critical velocity, a bag mode drop breakup occurs. At higher

gas velocltles, a shear or stripping mode occurs, consisting of boundary

layer stripping and cresting and breakup of unstable surface waves. The

diameter of the droplets formed is on the order of magnitude of the most

unstable wavelength, approximated by:

3 if _

k - (B-3I)

PG

(For water/air with a relattve velocity of 500 ft/sec, this wavelength

ls 23 microns).

Miscellaneous Studies

%j;

Harper et al (Ref. B-34) analyzed an accelerating liquid sphere as a boundary

value problem to determine the conditions under which small surface waves would

become unstable and grow. The study concluded that a liquid sphere would be subJect to surface instability If the product of the bond number and drop acceleration exceeded a critical value:

(PL- PG) d2 (d _\

\dt ) >Y (B-32)

i

where (d AV/dt) IS the acceleration of the drop relative to the gas.

Kreczkowski (Ref. B-55), in a recent investigation of drop breakup In a wind tunnel, observed the bag, transition, and stripping modes. He considered the

breakup to be controlled by three dimensionless groups: Weber number, Laplace

number, and the llquld/gas viscosity ratio, where the Laplace number is:

PL o d

La = 2 (B-33)

gL

r •

RI/R085-312

B-17
The crltJcal Weber number increased with llquld viscosity and the effect of viscosity on breakup time was minimal. Variation In viscosity by a factor of 1000

resulted In a breakup tlme increase of only a factor of 2.

Volynskly and Lipatov (Ref. B-91) derived a critical Weber number of 5.4 (Eq.

B-l) for the disintegration of a drop of Iow-vlscoslty fluid by elllpsoldal

deformation at low Reynolds numbers. At hlgh Reynolds numbers, a critical Weber

number of 15 tO 22 is suggested by the data of Korsunov and Tlshln (Ref. B-54).

The capillary wave analysis by Mayer (Ref. B-64) discussed above, which assumes

that waves formed on the llquld surface amplify, crest, and shed ligaments that

rapidly disintegrate to droplets, was extended by Adelberg (Ref. B-I and -2). He

postulated a region of high aerodynamic forces, In which the product of the Weber

and Reynolds numbers Is greater than lO6, where another type of breakup, called

acceleration wave breakup, occurs. Thls breakup mode occurs when the surface

waves are accelerated by the hlgh-veloclty gas stream and requires freestream

dynamic pressure above about 300 Ib/ft 2. The acceleration wave mechanism of

Jet breakup Is probably negligible for subsonic flows (Ref. B-2). Experimental

studies of acceleration wave Jet breakup are discussed In a subsequent section of

this report ("Effect of System Parameters on the Atomization Process").

Williams (Ref. B-96) reviewed earlier studies of fuel jet breakup in subsonic and

supersonic gas flows, and described the three mechanisms of Jet and drop atomization: shear stripping, capillary wave, and acceleration wave.

Borlsov et al (Ref. B-3) presented crlterla for drop breakup by several mechanisms. Defining the Weber number as one-half the customary value,

V

PG _2 d

We - 2 _ (B-34)

they gave We > 3 as the requirement for drop breakup.

RI/RD85-312

B-18
V

The observed breakup modes Included the bag and shear types, as well as Intermediate modes, and conformed to the following criteria:

Mode Criteria

Simple division 1

Parachute type (bag)

Chaotic (intermediate)

4<We<20 m

O.l < He Re-0"5 < 0.8

In these modes, sizes of the secondary drops were

of the same order of magnitude as the original drop.

Stripping (shear)

Yields mist of flne droplets lO<We<lO 4

0.5 < He Re-0.5 <lO

Explosive 103 < We < lO5

lO < We Re-0"5 < lO2

In this hlgh-gas-veloclty regime, droplet breakup is related to a Raylelgh-Lamb-Taylor instability and the Bond number (Bo = PL d2/_)"

Breakup times were given as follows:

tI =

time for drop deformation to the critical stage, when shear breakup

begins

(B-35)

RI/RD85-312

B-19
t 2

= characterlstlc drop acceleratlon time = tlme for AV to decrease to

one-half Its original value

(8-36)

t 3 = characteristic boundary layer establishment time

(8-37)

t 4 = drop destruction time in the explosive breakup mode

t 4 = lO We-0"25 t 1 (8-38)

Craig (Ref. B-lO) recently reported the results of an experimental investigation

of the aerodynamic breakup of liquid drops (e.g., water, ethanol, and mercury).

Holographic and laser veloctmetry were used for droplet-size measurements. With

original water and ethanol drop slzes In the range of lO0 to 600 microns, the

observed critical Weber numbers (on the order of lO0) were up to an order of magnitude hlgher than those previously considered to be necessary for breakup. Only

the bag mode of drop breakup was observed. However, other recent studies gave

values of the critical Weber number (based on original drop dlameter) in general

agreement with the results of the earlier investigators: 6 to lO (Ref. B-12), 12

(Ref. B-19), lO (Ref. 8-28), and 4 (Ref. B-35).

Fox and Dabora (Ref. B-25) studied the breakup of drops In a spray. The criterion for stripping mode breakup was found to be We > 20. The breakup times for

this mode were lower for drops in the wake of other drops than for free drops.

This was attributed to the effect of an increase in the effective gas density,

caused by the presence In the gas of droplets sheared from preceding drops.

RI/RD85-312

B-20
Gel'fand and Borlsov (Ref. B-28) emphasized the necessity of considering the

effects of drop breakup in models of spray combustion. They presented calculations of drop Weber and Reynolds numbers and breakup times as functions of chamber length, for several experimental conditions and showed that there was quite

sufficient chamber dwell time for drop breakup to occur. The predominant mode of

droplet breakup is stripping.

An important consideration in studies of liquid drops and sprays is the effect of

drop aggregations on the behavior of individual drops. This factor was examined

by O'Rourke (Ref. B-75) in a theoretical investigation of the dynamics an evaporation of drops in sprays. He defines three spray regimes:

"Very thin sprays": Total mass and volume of liquid are negligible in comparison

with those of the gas. Therefore, the gas is an infinite sink as far as the drops

are concerned, and collisions between drops need not be considered in analyses.

"Thin sprays": The drops have negligible volume but significant mass in comparison with the gas. Because of their small volume, interaction between drops is

not important, but because the mass of liquid In a given volume of the spray

field is of the same order as the mass of the gas, evaporation from the drops

affects the gas phase properties.

"Thick sprays": The drops occupy a significant volume fraction of the spray

field, which nevertheless consists of discrete drops In a continuous gas phase.

In this regime, drop interactions, such as collisions and modifications of drop

drag and vaporization rates resulting from close spacing, are important. Drop

collisions may be so significant that the size distributions produced by the

atomization process may be completely altered by subsequent coalescence and

shattering.

Petela (Ref. B-76) also recognized that there are different types of sprays (or

different regions within a single spray) in which drops of given diameter and

velocity may not behave in the same manner. He proposed modeling of the atomlzation process by computation of minimum stable drop sizes for each of three

breakup mechanisms: turbulent breakup at the orifice exit, aerodynamic breakup,

and breakup by mlcroexploslons.

RI/RO85-31 2

B-21
The changing character of drop sizes and drop-sIze distributions w%thin a spray

undergoing evaporation, combustion, or acceleration was originally pointed out by

Shapiro and ErIckson (Ref. B-86).

Drop Deformation and Drag Coefficient as Functions of Time

Drop Deformation and Breakup Mechanisms

A spectrum of drop breakup modes exists, ranging from the "bag" type at one

extreme to the "shear" type at the other. In bag-type breakup, the aerodynamic

force of the relative liquid/gas velocity deforms the drop lnto an ellipsoldal

shape, with its major axls perpendicular to the flow direction. This deformation

has been called Md%sk-shaped", "saucer-shaped", and "toroldal-shaped" by varlous

authors. As the deformation continues, the center of the drop opens like a bag

In the direction of the flow and appears as a th%n film of liquid anchored to a

heavler r%m around the drop perimeter and "stretched" In the flow direction until

the bag Is several times larger than the original drop or the circumferent%al

rlng of l%quid. When a crit%cal condition ls reached, the bag breaks Into a

shower of small droplets and the rlm dlslntegrates %nto several large droplets.

Photographs of bag mode breakup are shown in Ref. B-33, -55, -56, -62, -79, and

-97.

In the shear breakup mode, capillary waves formed on the surface of the liquid

are stripped off as ligaments that rapidly break Into droplets much smaller than

the parent drop. When the relative liquid/gas veloclty ls sufficiently high, the

stripping act%on appears as a shower of droplets being torn from the surface of

the drop. Typical photographs of shear mode breakup are shown In Ref. B-21, -55,

-56, -79, -80, -85, and -97.

In addition to the two extreme breakup modes, combined bag and shear breakup

(called the "bag-stamen" and "bag-shear" modes), have been observed photographically (Ref. B-33, -55, and -97.).

RI/RD85-31 2

B-22
v

The bag-type drop breakup process occurs at a lower Weber number than the shear

type and requires that the disrupting aerodynamic force be imposed for substantially larger times. The shear-type breakup mode occurs when the flow field

imposes shear forces on the drop surface faster than the drop inertia will permit

It to distort as a single mass, resultlng in stripping of a spray of daughter

droplets.

The shock tube tests of Rabln et al (Ref. B-79) showed a decrease In the critical

velocity of a given drop size as the flow duration was increased. Hence, a crltlcal drop diameter was postulated for a stated flow duration.

The time required for a drop to deform sufficiently from its original spherical

shape to one which induces breakup ("deformation time") was found to be inversely

proportional to the drop diameter. The deformation tlme was assumed to be

Inversely proportional to the gas flow velocity and may be the same as the prebreakup time previously mentioned.

v

An analytical model of drop deformation was presented by Chlu (Ref. B-6), based

on small perturbations and drop vibrations.

Collins and Charwat (Ref. B-9) constructed a fairly complex model that calculates

drop stripping, drag coefficient, deformation, and velocity as functions of time.

Model predictions showed good agreement with the experimental data then available.

Drop Drag Coefficients

A number of empirical correlations for the estimation of drop drag coefficients

have been proposed, most of which are related to the Reynolds number, defined by

d PG AV

Re - (B-39)

gG

RI/RD85-312

B-23
Time variation of the drag coefficient would therefore follow the corresponding

variation of Reynolds number.

• 611bert et al (Ref. B-29) reported the following empirical expression

for the drag coefficient of a single spherical particle•

CD = 0.48 + 0.28 Re-0"85 (B-40)

This value was also used by Ishlkawa and Murakalno (Ref. B-49) in a

study of sprays generated by diesel engine type orifice atomizers•

•

.

Ingebo (Ref. B-41) obtained the following equation for accelerating

groups of solid spheres and drops in the Reynolds number range 0.5 to 78:

Co = 27/Re 0.84 (B-41)

The reliability of this correlation has been questioned by Cllft and

Gauvln (Ref. B-B):

Rabln et al (Ref. B-TB and -79) measured drag coefficients experimentally In their shock-tube study of slngle-drop breakup. For drops

smaller than about lO0 microns, the drag coefficients agreed roughly

with those of Ingebo (Ref. B-41). For larger drops, however, there were

considerable differences, probably because smaller drops deform only

slightly from spherical shape at gas velocities below critical, while

larger drops deform into disk shape even at subcrltlcal velocities.

Their flndlngs indicate an expression of the form:

CD = 0.386 Re0"177

for droplet Reynolds numbers between lO2 and I04.

(B-42)

RI/RD85-312

B-24
. Hughes and Gllllland (Ref. B-39) proposed a modification of the standard

Stokes solid sphere drag coefficient (C d = 24/Re) to fit experimental

data In the Reynolds number range 0.5 to 200:

24 _/I 10-2 Re C = _-_ 1 + .II x

where _ = 3.75 x 105 (Re We)

(_ + V_-1 (B-43)

We = pG A-V2 d/_

. Crowe et al (Ref. B-ll) studied the effects of burning and acceleration

on the drag coefficients of particles suspended and accelerating In gas

streams at Reynolds numbers 250 to 1600. Boundary layer analyses indicated that burning and acceleration tend to reduce the drag coefficient.

Shock tube experiments were carried out on burning and nonburnlng solid

particles to measure the drag coefficients as functions of gas density

and relative velocity The values measured for burning particles were

imprecise, probably because of nonuniform burning rates. The following

correlation of the experimental data for nonburning particles was

reported for "low" gas velocities:

loglo C D = 2.586 - 1.705 loglo Re + 0.25 (loglo Re) 2 (B-44)

(200 < Re < 1600)

. Yuen and Chen (Ref. B-98) suggest that the equation for the drag coefficient for solid spheres can be applied to liquid drops In a gas stream

If the Reynolds number Is based on the following combination of liquid

and gas viscosities:

= gL + 0.33 (_G - _L ) (B-45)

RI/RD85-312

B-25
They note that the drag coefficient of a drop tn a spray cloud ts lower

than that of an isolated drop of the same diameter because the aggregate

of drops causes motion of the gas and changes the relative velocity.

V

. For drops undergoing shear mode breakup--and therefore at comparatively

high Reynolds numbers--the following values of the drag coefficient

(based on initial drop size and flow conditions) have been suggested:

3.0 (Ref. 8-80), 2.5 (Ref. 8-50), and 2 (Ref. 8-79).

Groeneweg (Ref. 8-32) constructed a mode] of drop motion that included

drag and evaporation as part of a study of the statistical description

of a spray in terms of drop velocity, size, and position.

The critical shortcoming In experimental measurements of drop drag coefficients

Is the difficulty of determining the effects of such factors as deformation,

acceleration, vaporization, and turbulence. Considerable variation between data

reported by various investigators is therefore not unexpected.

Sizes of Droplets Formed In Breakup of Large Drops

The following expression was derived (Ref. 8-97) for the "mean" droplet size

resulting from the breakup of a drop by aerodynamic forces that are much larger

than either the viscous or surface tension forces (i.e., at high Weber numbers):

I13

136 L cr

Omean = -- -- 1/2 (B-46)

L PG PL _'4

The photographic studies of the shock-lnduced breakup of an 1800-micron drop by

the shear-mode mechanism, which were previously discussed (Ref. 8-85) showed that

the daughter droplets are very much smaller than the parent drop, on the order of

SO microns. No breakup of the daughter droplets was observed In these

experiments.

RI/RD85-31 2

B-26
Effects of System Parameters on the Atomization Process

=

V

Typical available information on the effects of system parameters--particularly

liquid propertles--on the degree of atomization Is summarized in this section.

Appendix A is specifically limited to llquld rocket atomizers whlle the injectors

described here are designed for other appiicattons. They may, however, have some

applicability and utility in assessing the atomization characteristics of liquid

rocket injectors.

Alrblast Atomization

When the velocity of the gas stream is substantially higher than that of the

liquid, the process is referred to as "atrblast" atomization. Two general types

of alrblast atomizers have been studied. One uses the "preftlming" concept, in

which the liquid is spread out Into a thin continuous sheet and is then subjected

to the atomizing action of high velocity air*. To be fully effective, this system requires both sides of the llquld sheet to be exposed to the gas, which complicates design, since it entails two separate gas flows through the atomizer.

In the second, or "plain-Jet" concept, the liquid is injected into the highvelocity gas stream in the form of discrete Jets. The objective in both cases is

the same: to use the available energy of the flowing gas to achieve the maximum

degree of atomization. A recent review of alrblast atomization by Lefebvre (Ref.

B-58) was the source of many of the references and much of the discussion in this

section.

Drop-Slze Correlations

The drop-slze correlations developed for alrblast atomization, like those for

liquid rocket-type injector atomization, are highly empirical. Their applicability Is limited as a result of problems in drop-slze measurement techniques,

incomplete or no variation of fluid properties, and Improper development of data

Most of the alrblast atomization studies have used air; any gas might be used

as well.

RI/RD85-31 2

B-27
correlations. In addition, various investigators use different average drop

sizes and, without data on size distributions, conversion to a common basis for

comparison is difficult**.

Nevertheless, sufficient information is available to permit some generalized concluslons as to the effects of fluid properties and flow parameters on the degree

of atomization produced by atrblast atomizers.

For convenience of reference, the most significant of the various correlations

proposed for the mean drop sizes produced by all types of airblast atomizers are

summarized in Ref. B-58. It is strongly recommended that these correlations be

very carefully considered before applying them to any type of liquid rocket, gas/

liquid injectors. These are all empirical relationships with little or no theoretical basis. Their application will generally require the extrapolation of

these correlations far beyond the range of test conditions used in their development. Also, a great variety of airblast atomizers is available, and plain-Jet

airblast atomizers come In many different shapes and forms. It is very important

that these factors be carefully considered prior to application of these correlations. As shown in Section V, Atomization Study, even the relatively minor

extrapolation of one of these correlations (Lorenzetto and Lefebvre--Ref. B-82)

to the conditions of the tests performed In this program, can result In very

major errors. The nomenclature used in this section is listed In Table B-1.

-- A

Effects of System Variables on Mean Drop Size

Liquid Properties. The liquid properties of importance In alrblast atomization

are viscosity, density, and surface tension. The adverse effect on spray quality

of an increase in viscosity at varying levels of gas velocity and constant liquid

Note, however, that the ratio of mass mean diameter to Sauter mean diameter

has been reported to be 1.20 ± 0.006 (Ref. B-33)

RI/RD85-312

B-2B
TABLEB-l. NOMENCLATURE FORAIRBLASTATOMIZATION

A, B

D

D

0

DL

Dp

L

Lc

P

Q

Re

U

W

We

P

o

P

MMD

SMD

SUBSCRIPTS

EXPERIMENTAL CONSTANTS

DROP SIZE, m

LIQUID ORIFICE DIAMETER; INITIAL 3ET DIAMETER, m

DIAMETER OF ATOMIZER CUP (AT LIP), m

DIAMETER OF PREFILMER LIP, m

LENGTH, m

CHARACTERISTIC ATOMIZER DIMENSION, m

PRESSURE, Pa

VOLUMETRIC FLOW RATE, llters/sec

REYNOLDS NUMBER, ULp/p

VELOCITY, m/sec

MASS FLOW RATE, kg/sec

WEBER NUMBER, U2Lo/o

DENSITY, kg/m 3

SURFACE TENSION, kg/sec 2 OR N/m

DYNAMIC VISCOSITY, kg/m/sec OR N-sec/m 2

KINEMATIC VISCOSITY, m2/sec

MASS MEAN DIAMETER, m

SAUTER MEAN DIAMETER, m

GAS

LIQUID

GAS RELATIVE TO LIQUID

k.j

RI/RD85-312

B-29
flow rate is shown for a prefllmlng and a plaln Jet atomizer in Fig. B-I (Ref.

B-83) and Fig. B-2 (Ref. B-60), respectively. Viscosity forces tend to suppress

the formation of capillary waves on the liquid surface, which precedes atomlzatlon, and also resist deformation of subsequently formed ligaments into drops.

200L f2)

2201 , , , I , t I

200_- wL'OOz5kQ/s _o J

4C),_ J T,, 296°K

_L: 0 015 kg/s

_,_ _ ,25 ,. _ joI _ '_ _o:_o,_ ,_o- o, ,'o_'o_' " " ,'o,'_

ABSOLUTE VISCOSITY, kg/ms. I0_ ABSOLUTE VISCOS:TY, kglm$ • 0 3

Figure B-1. Variation of Mean Drop

Size Wlth Liquid Viscosity

for a Prefilmlng Atomizer

(Ref. B-83)

Figure B-2. Variation of Mean Drop

Size With Liquid Viscosity

for a Plaln-Jet Atomizer

(Ref. B-60)

Surface tension forces tend to impede atomization by resisting disturbances or

distortions of the liquid surface, thereby opposing the creation of surface waves

and delaying the onset of ligament formation. The effect of surface tension on

drop slze is illustrated in Fig. 3 (Ref. B-83) and Flg. 4 (Ref. B-60) for a prefilming and a plain Jet atomizer, respectively.

Liquid density affects droplet size in a complex manner, the net result of which

is that the influence of density is minor. For prefllmlng atomizers, drop slze

increases slightly wlth density, while the opposite occurs wlth plaln-Jet nozzles, as shown In Flg. 5 (Ref. B-82) and Fig. 6 (Ref. B-60), respectively.

RI/RD85-312

B-30
v

In

L)

[E 80

S

:}

U)

4O

I_M:)| ' I I I ] I

L

uguIDrl.oW_mE,_WL

• 0030 kg./_

120 x 0025

• 0 020 "

o 0 015 "

OOlO "

I I

t

l I I I I 1 ,I

IO 20 30 40 50 60 70 80

SURFACE TENSION, kg/s 2, 103

Figure B-3.

140

120

IOO

(E 6O

u3

4O

20--

0

OB

Figure B-5.

Variation of Mean Drop

Slze With Surface Tension

for a Prefllmlng Atomizer

(Ref. B-83)

• 105N/'m 2

"TA=296°K

-- U,,, m/$

549

_ " _

i_..._.--_ ''_ 914 x

_ 103 6,_

v--

1219 'I"

I 1 I I 1

I0 12 14 16 18

LIOUID DENSGY, K_/rn _. 103

Variation of Mean Drop

Slze With Liquid Density

for a Prefilmlng Atomizer

(Ref. B-82)

'601

/ ,

m ._1 40 1307)

l I I I i I I

0 20 30 40 50 60 70 80

LIQUID SURFACE TENSION, Nlm • I0 _

Figure 8-4. Variation of Mean Drop

Size With Surface Tension

for a Plain-3et Atomizer

(Ref. B-60)

120

I00

8O

u_ZO

I

_' RATIO Pro' ' lOSN"me /

I "'I I l l''Z? TM ] I

08 IO r2 14 _6 J8 20

LIOUID DENSITY, kg/n_. 1(73

Figure B-6. Variation of Mean Drop

Size With Liquid Density

for a Plaln-3et Atomizer

(Ref. B-60)

Gas Properties. Gas velocity is the most important factor contro111ng mean

i

drop slze in alrblast atomizers, as shown in Fig. B-l, -2, -4, -5, and -6. For

typical Iow-vlscoslty liquids, the Sauter mean diameter (SMD) is approximately

inversely proportional to alr velocity. A second important factor is the air/

liquid mass fiow-rate ratio, as shown for prefilming atomizers In Fig. 7 (Ref.

B-84) and Fig. 8 (Ref. B-20). Atomization quality starts to decline when the

RI/RD85-31 2

B-31
I0O

8c

{.}

_.6c

4C

i l I | I I ] ! i l I

ATOmiZING VELOCIT Y,U A

l 0 54 g _/J

• 6? I "

-- x "t93 "

II • 914 "

_, ,". 1036 "

¢ iill g "

, ] , , , , ,5

2.0 4.0 60 80 I00 120

ATOMIZINGAIR/LIOUID MASS RATIO

Figure B-7. Effect on Mean Drop Size

of Air/Liquld Mass Ratio

(Ref. B-84)

40 - KEROSINE o/O_)60

Dp,1905ram .._ /70

P..,e_,m, <# j

V'29s'_ t

35- 9°

JO0

:_ 1t0

 o177///

25L/ /i t i 1

0 02 04 06 08 JO

LOG(_._Z%)

Figure B-8. Graphs Illustrating

Relationship Between Mean

Drop Slze and Alr/Liquld

Ratio (Ref. B-20)

alr/llquld ratio Is below about 4 and deteriorates rapidly below a ratio of

approximately 2. Increasing the air/liquid ratio beyond about 5 gives only marglnal drop-size decreases.

The effects of air temperature and pressure on mean drop sizes are shown In Flg.

9 and lO (Ref. B-84).

300,

200-

C'}

100 -

$2

u_ 50 --

40--

30--

2%00

Figure B-9.

..... WA'TER

KEROSINE

_, IOS_,',..2

-,\

"% _lU i • SOm,'$

(x,O.72)

_...UA ,60m_

%.'_ _,o73)

1 I I I

soo¢oo300 2oo loo

AIR TEMPERATURE°K

Influence of Alr Temperature on Mean Drop Slze

(Ref. B-B4)

2°01 0 WATER

x KEROSINE

I00

__6o

d 40

_ 3O

2o

iO • 1 I_LI I1111

2 3 4 ,5 678910

-AtR PRESSURE (N/m2 • I05)

Figure B-lO. Influence of Alr Pressure

on Mean Drop Size

(Ref. B-84)

z

V

RI/RD85-312

B-32
V

Summary of Property Effects. The effects of the major flow parameters on

mean drop size, for Iow-vlscoslty liquids, are summarized in Ref. B-83, together

with the ranges of test conditions that were Covered in the experimental investigations, the liquids used, methods of measurement, and drop-slze data.

Interestingly, the various expressions proposed for predicting the mean droplet

sizes produced by prefilmlng alrblast atomizers show some consistency. Thus, for

Iow-vlscoslty liquids, the effects of the major variables on mean drop size may

be expressed as power dependencies, with exponents in fairly narrow ranges:

Air velocity

Air density

Liquid density

Surface tension

(l ÷ WL/W A)

Linear scale

-l.O to -I.2

-0.6 to -0.7

0 to -0.25

0.5 to 0.6

0.85 to l .0

0.4 to 0.5

The dimensionally correct expression for mean drop sizes produced by prefilmlng

alrblast atomizers that best satisfies experimental data is (Ref. B-20):

SMD = + WAA .073 (PA_UA2 \TA/ OPO"4 + (8-47)

°sI

where D

P

is the prefilmer diameter.

This comparatively recent correlation was developed for the atomization of water

and kerosene; drop sizes were measured by a Fraunhofer diffraction llght-scattering technique. The variables used in this study were atomizer size (prefilmer

llp diameter), and liquid density, viscosity, and surface tension. Agreement

between measured SMD values and those predicted by Eq. B-47 was good (Fig. B-ll).

RI/RD85-31 2

B-33
140 KEROSzNE end WA'I"ER DATA

UA' 60" 150 m/s 411

120 AFR =O.5-50 •

I

20 40 60 (30 K:)O 120

S MD, ME _SURED-MICRONS

Figure B-11. Comparison of SMD Values, Measured

vs Predicted From Eq. B-47 (Ref. B-58)

The most recent and reliable correlation for plain-Jet atomization was reported

by OasuJa (Ref. B-52), who varied liquid properties and alr pressure. A Fraunhofer diffraction laser light-scattering technique was used for drop size measurements. The experlmental data were falrly well correlated (±30%) by the following expression:

÷ 1.43 x 10 -3

\_ PL/ + WG/ (B-48)

This correlation does not include geometric (i.e., liquid orifice diameter)

effects, since atomizer slze was not varied In the investigation. A comparison

of the performance of plaln-Jet and prefilmlng alrblast atomizers is presented in

Ref. B-42, which indicates that despite the simpler design of the former, Its

experimentally measured SMDs are only slightly larger, particularly for Iow-vlscosity liquids.

Significant differences occur In the power dependencies in the drop size equatlons for prefllmlng and plaln-Jet atomizers. For example, liquid density

appears to have opposite effects on the SMDs of the two types. This may be

RI/RD85-312

B-34
_J

caused by different mechanisms of bulk liquid conversion Into a droplet spray, so

0.1

that the SMD is proportional to (pL/PA) in prefllmlng types and to

-0.3

(pL/PA) In plain-Jet units.

The experimental data obtained from both prefilmlng and plaln-Jet alrblast atomizers with Iow-viscoslty liquids (water or Eerosene) indicate that the main factors governing product drop size are the liquid surface tension, and the air velocity and density; with hlgh-vlscoslty liquids, the effects of air properties are

less significant, and the SMD is more dependent on the liquid properties, especially viscosity. The observation that the effects of liquid and gas properties

are separate (Ref. B-20, -51, -53, -60, -73, and -83) suggests an expression for

SMD consisting of the sum of two terms:

SMD = (SMD) l + (SMD) 2

(B-4g)

where (SMD) 1 Is dominated by air velocity and density and (SMD)2• by liquid

viscosity. The two different but complementary mechanisms of alrblast atomlzatlon--one governed by the ratio of aerodynamic to surface tension forces, as

2 D/a), and the other by viscous embodied in the Weber number (PG UG

forces, as expressed in the Z number (_/aPLD)--support_ the idea that

any expression for SMD should consist of two terms to represent these effects.

This provides a theoretical basis for the SMD equations for alrblast atomizers.

In actual application of the equations to different atomizer designs, however, it

is necessary to divide the calculated value of SMD by a factor, ¢, which represents the overall efficiency of the atomizer, compared to the one for which the

equation was developed, as well as to account for different methods of drop sampling and drop-slze measurement. For prefilmlng atomizers, values of ¢ were

found to be as low as 0.6 and were sensitive to minor changes in atomizer geometry. With the right values of ¢, data from various experlmentors agree fairly

well wlth Eq. B-47, supporting its general validity to prefllmlng alrblast atomizers. Unfortunately, the available data were obtained under conditions in which

(SMD) l as very large compared to (SMD) 2. Tests in which the (SMD)2/(SMD) l

ratio is maximized are required for accurately checking the general appilcablllty

of the correlation.

RI/RD85-312

B-35
The most acceptable expression for the SMDobtained From plaln-Jet alrblast atomizers (Eq. B-48) has the sameform as that for the prefilmlng type (i.e., it consists of the sum of a term dominated by air velocity and density and a term domInated by liquid viscosity). In fact, the similarity of the two correlatlons is

strong evldldence that both types of atomizer function by the same fundamental

processes.

It has been suggested (Ref. B-58) that at least two different mechanisms are

involved in alrblast atomization; their relative importance depends on the level

of liquid viscosity. For Iow-vlscoslty liquids injected into a comparatively

1ow-veloclty gas stream, capillary waves are produced on the liquid surface.

These waves grow, become unstable, and break off as ligaments that disintegrate

into droplets. Increased gas velocity causes earlier formation of the ligaments,

which are thinner and shorter, and disintegrate into smaller drops. With liquids

of high viscosity, the caplllary wave mechanism may not be valid. Instead, the

liquid is drawn out in the form of long ligaments that break up relatively

slowly, in regions of lower gas velocity, into larger-slze droplets.

An investigation of an external mixing* atomizer using kerosene/alr (Ref. B-18)

resulted in a drop-slze correlation that was not of the two-term type, nor do the

property exponents follow those of other empirical correlation

V

SMD = (l x lO-3)Do (Re)

-0.29

0.39(We)0.18 /WG_

where

m

Re = PL Do AV/_L

We = PL Do _21°

*The prefllm and plaln-Jet alrblast atomizers previously discussed are "internal

mixing" types, in which the gas/llquld interaction occurs within the atomizer

body. In external mixing, the interaction occurs outside the atomizer.

RI/RD85-312

B-36

V
7

v

i

The characteristics of a Y-shaped alrblast atomizer, in which the liquid Jet is

atomized by the air In an internal chamber, were studied by Prasad (Ref. B-77).

p

His drop-slze data conformed in general to the empirical correlation of Wigg

(Ref. B-95):

+ hO.l o0.2

wG/ (85 >

where h is a characteristic atomizer dimension. The data of Mullinger and Chtgter

(Ref. B-70) were also in fair agreement with this expression.

Jet Atomization

Several investigators have studied the atomization of single Jets of liquid in

low- or moderate-veloclty gas flow fields. Although these may be considered as

types of plaln-Jet airblast atomizers, they are not generally so classified and

will therefore be discussed separately.

Ingebo has worked extensively In this field. In Ref. B-43, he reported drop size

data obtained for the conditions VG = O, VG > VL, and VG increasing with

constant acceleration. Maximum measured drop sizes were correlated in terms of

several dimensionless numbers by the following expression:

(DoBO_0"33 (ReG)0"5 [

Do - 0.64 +

Dmax \-_-t / (ReL)O.l(Bo)O.07 0.044 +

We L

(WeG)O'2(Bo) 0-09

1.25 x 10-6 )

+ (3.5 (Ac)0"5_

x

(B-52)

RI/RO85-312

B-37
where

D

0

Dt

D_x

Bo

ReG

ReL

WeG

WeL

Ac

= 11quid orlflce dlameter

= gas tip dlameter

= maxlmum droplet slze

= bond number = PL D_/a

= gas Reynolds number = PG Oo VG/PG

= llquld Reynolds number = PL Do VL/PL

= gas Weber number = PG Do _2/a

= llquld Weber number = PL Do _2/_

= aerodynamic acceleration = PG 02o a/o

= gas stream acceleration (constant)

The maximum measured droplet diameter may be considered to be associated with the

critical Weber number at the specified flow conditions and to constitute an

experimental determination of this factor.

In an Investigation of the effect of gas velocity on mean droplet sizes (Ref.

B-44), Ingebo found that D32 was proportional to VG0"75 with a subsonic alr flow

atomizer and to the product (VG)-O'TS(FN)0"4 In a pressure atomizer, where FN is

the flow number:

FN = --q- (liters/hour)

0.5 (B-S3)

(N/m2)

Results of a study of the atomization of water Jets (V L = 5 and 23 m/sec) by a

moderate-velocity air flow (V G = 40 - 150 m/sec), swlrled and unswlrled, at

wind tunnel pressures between one and two atmospheres, were reported in

Ref. B-45. Swirl of the llquid Jet reduced droplet sizes, compared to unswlrled

Jets, and llttle difference was found between upstream, downstream, or crossstream injection.

RI/RD85-312

B-38
-- =

An experimental investigation of the acceleration wave breakup mode of liquid

Jets was reported by Ingebo (Ref. B-46). Water Jets injected into stagnant air

(V G = O) start to disintegrate at a distance of approximately 4.5 times the Jet

diameter (from Rayletgh analysis). Jets injected tnto high-velocity air streams

break up by the capillary wave or acceleration wave mode, depending on the air

velocity, with mean droplet sizes given by:

0.25

Do/DIo = 0.21 (WeGReL) (B-54)

for (WeGReL) < lO6 (capillary wave breakup mode)

and

Do/DIo = 0.27 (WeGReL)0"4

for (WeGReL) > lO6 (acceleration wave breakup mode)

(B-55)

The data showed that In acceleration wave breakup (very high gas velocity and/or

density) the effect of relative gas/llquid velocity is magnified and the effect

of orifice diameter is minimized, compared to the capillary wave breakup mode.

Further experimental work on acceleration wave breakup for various types of water

injection into air streams (Ref. B-47) gave the following minor modifications of

Eq. B-55:

(i) For downstream injection, nonswlrllng air flow,

0.4

Do/DIo = 0.23 ..(WeGReL) (B-56)

(2) For downstream injection, swirling airflow, Eq. B-55 is valid.

(3) For upstream injection, nonswirllng air flow,

0.5

Do/DIo = 0.0045 (WeGReL) (B-57)

RI/RD85-312

B-39
Ingebo (Ref. B-48) recently reported results of an experimental study of the

atomization of liquid sheets In quiescent and hlgh-veloctty air. Emplrtcal correlations were made in terms of the liquid Reynolds number (Re L = DoVLPL/UL), the

alr stream Reynolds number; (Re 6 = DoVGPG/_G); and the relative velocity Reynolds

number (Re R = DoA'Vp6/_6).

For like-doublet impinging Jets*,

Do/D m = 0.023 (ReL)0"5 ÷ 0.002 (ReR) (B-58)

where D° ls the orifice diameter and Dm is a "mean" droplet diameter, slmllar

to D32. (A nearly Identical expression was derived by Ingebo for the heptane

Impinging Jet data reported In Ref. B-8.)

For splash plate Injectors,

Do/D m = 2.9 x I0-4 (ReL) _ 2.4 x lO-3 (ReG) (B-5g)

For atomizers producing swirling, hollow-cone sheets,

Do/D m = (Do/Dm, h) ÷ 2.2 x 10-3 (Re G - Rec) (B-60)

where Dm, h and Rec are constants, defined as the hydrodynamic mean drop diameter and the critical Reynolds number for aerodynamic breakup, respectively. The

same type of simplex swirl atomizer was recently used to spray Jet and diesel

fuels Into ambient and heated alr at pressures of 20 to 80 psla (Ref. B-17).

Measured droplet sizes showed a strong effect of alr density:

-0.53

SMD _ PA (B-61)

V

*Impinging Jets form a liquid sheet before breakup.

RI/RD85-312

B-40
i

i

I

i

i 7

it Is of interest that Eq. B-58, B-S9, and B-60 are two-term expressions, one

related to hydrodynamic forces (ReL) and the other related to aerodynamic

forces (Re R) or (ReG). This may be compared to the two-term correlations for

alrblast atomizers previously discussed.

In another study of the breakup Of llquld Jets by cross-flowing alr streams,

Husseln et al (Ref. 8-40) constructed a model of stream breakup. Drop-size measurements made to verify the model showed that the SMD varied stgntflcantly over

an axial distance of 5 cm from the injection point. Beyond 5 cm, the sizes were

relatively constant and showed the following dependencies: with air velocity,

" DO.5S SMD = V 1.34 wlth orifice diameter, SMD = • and wlth liquid

' 0 '

jet velocity, no appreciable effect.

Matta (Ref. B-63) conducted an experimental investigation of the breakup behavior

of viscoelastic liquids in hlgh-veloclty alrstreams. He modeled the process as

simply the breaking off of ligaments from the Jet which then neck into a series

of drops, with no significant interaction with the gas flow. For the fluids

tested (e.g., glycerine, dlethylmalonate, and various polymers), the "average"

drop size was a function of the Ohnesorge number:

Dmean = 1.88 DO (I + 3Z) I/6 (B-62)

where

Z = ULI(DoPL(_)I12

Simmons and Harding (Ref. B-88) measured the drop sizes in sprays produced by

simplex pressure-atomlzlng nozzles. They found that water and kerosene SMD

values showed differing dependencies on surface tension. However, by defining a

Weber number based on the liquid sheet thickness, t, (We = PGV_t/a),_

A

they could correlate their data by

0.51

SMD = o (We < l),

0.16

SMD = o (We > l),

SMD _ AP-0"275 We-0"4

RI/RD85-312

B-41
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m_--.

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V

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Morrell, G., and F.P. Povinelll, "Breakup of Various Liquid Jets by Shock

Naves and Application to Resonant Combustion," NASA TN-D-2423 (]964)

Mulllnger, P.H., and N.A. Chtgler, "The Design and Performance of Internal

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Nicholson, J.E., and A.F. Hill, "Rain Erosion on Splke-Protected Supersonic Radomes," Report No. MC-61-6-R3, Mlthas, Inc., Cambridge, Massachusetts (1963)

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in an Airstream," Trans. Society of Mechanical Engineers, Japan, Vol. 5,

pp. 68-75 (1939)

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LA-9069-T, Los Alamos National Laboratory, Los Alamos, New Mexico (1981)

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=

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RI/RD85-312

B-49

F

APPENDIX C

DISTRIBUTION LIST FOR FINAL REPORT

ATOMIZATION AND MIXING STUDY

NAS8-34504

NAME

National Aeronautics & Space Administration

Washington, DC 20546

Attn: RP/F.W. Stephenson

George C. Marshall Space Flight Center

National Aeronautics & Space Administration

Marshall Space Flight Center, AL 35812

Attn: AS24D

AlOl

EM12-34

EP23/C.R. Bailey

EP22/F.W. Braam

EL24/K.W. Gross

Lewis Research Center

National Aeronautics & Space Administration

21000 Brookpark Road

Cleveland, OH 44135

Attn: MS 60-3/Library

MS 500-219/S. Gorland

MS 500-200/L. Diehl

AMES Research Center

National Aeronautics & Space Administration

Moffett Fleld, CA 94035

Attn: Library

Goddard Space Flight Center

Natlonal Aeronautics & Space Administration

Greenbelt, MD 20771

Attn: Library

John F. Kennedy Space Center

National Aeronautics & Space Administration

Kennedy Space Center, FL 92899

Attn: Library

NUMBER OF COPIES

1

2

l

RI/RD85-312

C-1
Lyndon B. Johnson Space Center

National Aeronautics & Space Administration

Houston, TX 77058

Attn: Library

EP2/C.R. Gibson

Langley Research Center

National Aeronautics & Space Administration

Langley Statlon

Hampton, VA 23665

Attn: Library

Jet Propulsion Laboratory

National Aeronautics & Space Administration

4800 Oak Grove Drive

Pasadena, CA 91109

Attn: Library

NASA-Sclentlflc & Technical Information Facility

P.O. Box 8757

Baltimore-Washington International Airport

Baltimore, MD 21240

Attn: Accessing Department

Defense Documentation Center

Cameron Station, Building #5

5010 Duke Street

Alexandria, VA 22314

Attn: TISIA

Air Force Rocket Propulsion Laboratory

Edwards, CA 93523

Attn: Library

Arnold Engineering Development Center

Air Force Systems Command

Tullahoma, TN 3?388

Attn: Library

U.S. Naval Research Laboratory

Washington, DC 20390

Attn: Library

V

RI/RD85-312

C-2

V
i_--o

T_,.I

U.S. Army Missile Command

Redstone Scientific Information Center

Redstone Arsenal, AL 35808

Attn: Documents Section

U.S. Naval Missile Center

Point Mugu, CA 93041

Attn: Technical Library

Battelle Memorial Institute

505 King Avenue

Columbus, OH 43201

Attn: Library

Bell Aerosystems, Inc.

Box 1

Buffalo, NY 14240

Attn: Library

Boeing Company

Space Division

P.O. Box 868

Seattle, WA 98124

Attn: Library

General Dynamlcs/Convalr

P.O. Box I128

San Diego, CA 92112

Attn: Library

Grumman Aerospace Corporation

Bethpage, Long Island, NY I1714

Attn: Library

Lockheed Missiles & Space Co.

P.O. Box 504

Sunnyvale, CA 94088

Attn: Library

2

RI/RD85-312

C-3
Marquardt Corporation

16555 Satlcoy Street

Box 2013 South Annex

Van Nuys, CA 91409

Attn: Library

Martln-Marletta Corporation

P.O. Box 179

Denver, CO 80201

Attn: Library

McDonnell Douglas Astronautics

5301 Bolsa Avenue

Huntington Beach, CA 9254?

Attn: Library

Rocket Research Corporation

Willow Roadat ll6th Street

Redmond,WA 98052

Attn: Library

TRW Systems, Inc.

l Space Park

Redondo Beach, CA 902?8

Attn: Library

United Technologies Research Center

East Hartford, CT 06108

Attn: Library

United Technologies Pratt & Whitney

Government Products Division

P.O. Box 2691

West Palm Beach, FL 33402

Attn: Library

AeroJet Tech Systems Company

P.O. Box 13222

Sacramento, CA g5Bl3

Attn: Library

RI/RDBS-312

C-4