,! j

NASA-CR-191222 ""

PENNSTATE (NASA-CR-I91222) COMBUSTION OF LOX

WITH H2(SUB 9) UNOER SU_CRITICAL,

CRITICAL, AN0 SUPERCRITICAL

CONDITIONS AND EXPERIMENTAL

O_SERVATION OF _ENSE SPRAY AND

MIXING OF IMPINGING JETS Semiannual

Report, i Feb. - 31 Oct. 1992

(Pennsylvania State Univ.) 63 p

• , /

/

/

N93-I6647

Unclas

G3/25 0137991

PROPULSION ENGINEERING RESEARCH CENTER

SEMI-ANNUAL REPORT

COMBUSTION OF LOX WITH HYDROGEN UNDER SUBCRITICAL,

CRITICAL, AND SUPERCRITICAL CONDITIONS (T,,-k 1)

and

EXPERIMENTAL OBSERVATION OF DENSE SPRAY AND MIXING OF

LIQUID JETS EMANATING FROM DOUBLET INJECTORS (T,,,k 2)

Prepared by: K. K. Kuo, W. H. Hsleh and F. B. Cheung

A. S. Yang, M. C. Kline, R. D. Woodward and J. J. Brown

TO

Mr. Klaus W. Gross, Code: EP 55

NASA: Marshall Space Flight Center

Huntsville, Alabama 35812

A UNIVERSITY SPACE ENGINEERING

RESEARCH CENTER

November 1992

106 RESEARCH BUILDING EAST

UNIVERSITY PARK, PENNSYLVANIA 16801
SEMI-ANNUAL REPORT

Combustion of LOX with H_(pO Under Subcritical 1 Critical, and

Supercritical Conditions (Task I)

and

Fx, perimental Observation of Dense Spray and

iixinl_ of Imvinl]_in2 Jets flask 2)

Submitted to

Mr. Klaus W. Gross

Code: EP 55

NASA: Marshall Space Flight Center

Huntsville, Alabama 35812

Prepared by

IC K. Kuo, W. H. Hsieh, A. S. Yang and J. J. Brown (Task 1)

K. K. Kuo, F. B. Cheung, R. D. Woodward, M. C. Kline (Task 2)

PROPULSION ENGINEERING RESEARCH CENTER

The Pennsylvania State University

University Park, PA 16802

November 1992
TASK NO. 1

TABLE OF CONTENTS

COMBUSTION OF LOX WITH H2(?0 " UNDER SUBCRITICAL,

CRITICAL, AND SUPERCRITICAL CONDITIONS

ACKNOWLEDGMENT ............................................. iii

LIST OF FIGURES - Task 1 ........................................... iv

LIST OF FIGURES - Task 2 .......................................... v

NOMENCLATURE ................................................ vi

1. INTRODUCTION ............................................. 1

2. EXPERIMENTAL WORK ...................................... 2

2.1 Evaporation Tests ........................................ 2

2.2 Combustion Tests ........................................ 8

3. THEORETICAL WORK ...................................... 11

3.1 Improvement of High Pressure Phase Equilibrium Model .......... 11

3.2 Implementation of Numerical Code for Phase Equih'brium Model .... 16

3.3 Verification of the High Pressure Phase Equilibrium Model ........ 19

3.4 Verification of Numerical Solutions from the Flow Analysis Code .... 22

4. DISCUSSION OF RESULTS ................................... 25

4.1 Experimental Results ...................................... 25

4.2 Theoretical Results ...................................... 28

OVERALL STATUS .......................................... 34

35

36

36

36

38

38

Shut-Down Procedure ......................................... 39

.

References of Task 1 ...............................................

Appendix A ......................................................

Operational Procedure for LOX/H 2 Combustion Tests ......................

Overall Setup ...............................................

Setup Checklist ..............................................

Test Procedure ..............................................
TASK NO. 2

Ii

2.

3.

4.

o

TABLE OF CONTENTS (Cont'd)

DENSE SPRAY AND MIXING OF LIQUID JETS EMANATING

FROM DOUBLET INJECTORS

INTRODUCTION ............................................ 40

PROCEDURE FOR INJECTOR CAVITATION STUDY .............. 40

DATA REDUCTION ......................................... 42

RESULTS OF THE INJECTOR CAVITATION STUDY .............. 43

4. I Observation of the Cavitating Region ......................... 43

4.2 Relationship Between the Internal and External Flow ............ 48

REMARKS ................................................. 55
°,°

IU

ACKNOWLEDGMENT

This research project has been funded by NASA-MSFC under Contract Number

NAG8-174. The authors would like to thank Mr. Klaus W. Gross of MSFC for his

encouragement and support. A part of this work was supported by the Propulsion Research

Engineering Center at The Pennsylvania State University (PSU). The interest and support

of Prof. Charles L. Merkle, Director of PERC, is also acknowledged. Valuable input from

Mr. David T. Harrje, a specialist in liquid rocket propulsion, on supercritical combustion of

LOX is deeply acknowledged.
Figure 1.1

Figure 1.2

Figure 1.3

Figure 1.4

Figure 1.5

Figure 1.6

Figure 1.7

Figure 1.8

Figure 1.9

Figure 1.10

Figure 1.11

Figure 1.12

Figure 1.13

Figure 1.14

Figure 1.15

Figure 1.16

iv

LIST OF FIGURES - Task 1

Schematic Diagram of the Apparatus Used for the Calibration of Mass

Flowmeters ............................................. 4

Calibration Curve of the Mass Flowmeter Used for the Oxygen

Consumption Rate Measurement ............................. 5

Calibration Curve of the Mass Flowmeter Used for the Helium

Consumption Rate Measurement ............................. 6

Schematic Diagram of the Apparatus Used for Oxygen Concentration

Measurements ........................................... 7

Picture of a Methane/Oxygen Diffusion Flame (Re = 300) .......... 9

Picture of a Methane/LOX Diffusion Flame (Re = 75) ........... 10

Comparison of Computed and Experimental High-Pressure Phase

Equilibria; Oxygen/Helium System ............................

20

Phase Equilibrium Calculations for the Oxygen/Helium System at

Various Pressures and Temperatures .........................

21

Comparison of Exact Solution and Calculated Axial Velocity Profile in

a Fully Developed Pipe Flow ............................... 23

Convergence Rate of Normalized Mass Residual for Code

Verification ............................................ 24

Measured Evaporation Rate vs. Pressure for LOX in Helium

Environments at Room temperature (Purge Flow Velocity = 0.2

cm/see) ............................................... 26

Measured Surface Temperature vs. Pressure for LOX in Helium

Environments at Room temperature (Purge Flow Velocity = 0.2

era/see) ............................................... 27

Measured Center-line Oxygen Mole Fraction Profiles above the LOX

Surface ............................................... 29

Temperature Contour for the Evaporation of LOX in a Helium

Surrounding at a Pressure of 30 atm ......................... 31

Oxygen Mass Fraction Contour for the Evaporation of LOX in a

Helium Surrounding at a Pressure of 30 atm ................... 32

Measured Oxygen Mole Fraction Data and Predicted Center-line

Temperature and Oxygen Concentration Profiles above the LOX

Surface ...............................................

33
V

LIST OF FIGURES - Task 2

Figure 1. Schematic of Mass Flow Rate Determination ................... 41

Figure 2. Flow Variation with Increasing Reynolds Number Using the Plexiglas

Injector Channel

(a) AP = 0.0 psi., v = 0.0 m/s, Re d = 0.0

(b) AP -- 0.4 psi., v = 1.82 m/s, Re a _ 7.21 x 103 ................ 44

Figure 2 Flow Variation with Increasing Reynolds Number Using the Plexiglas

Injector Channel

(c) AP = 3.7 psi., v = 5.54 m/s, Re d _ 2.19 x 104

(d) AP = 14.7 psi., v = 11.04 m/s, Re d _ 4.37 x 104 .............. 45

Figure 2 Flow Variation with Increasing Reynolds Number Using the Plexiglas

Injector Channel

(e) _P = 30.05 psi., v = 15.79 m/s, Re d _ 6.25 x 104 ............ 46

Figure 3 Variation of the Reattachment Point Within the Orifice ........... 47

Figure 4 Visualization of the Internal Flow with Increasing Reynolds Number

within the Plexiglas Injector

(a) _d' = 0.3 psi., v = 1.47 m/s, Re d _ 2.90 x 103

(b) 61_ --- 13.5 psi., v = 9.84 m/s, Re d _ 1.95 x 104 ...............

50

Figure 4 Visulation of the International Flow with Increasing Reynolds Number

within the Plexiglas Injector

(c) _P = 19.8 psi., v = 11.92 m/s,Re a _ 2.36 x 104

(d) ZkP = 23.8 psi., v = 13.07 m/s, Re a _ 2.59 x 104 .............. 51

Figure 4 Visulation of the International Flow with Increasing Reynolds Number

within the Plexiglas Injector

(e) Ap = 26.5 psi., v = 13.79 m/s, Re a _ 2.73 x 104

(f) _P = 31.1 psi., v = 14.94 m/s, Re d _ 2.96 x 104 .............. 52

Figure 4 Visulation of the International Flow with Increasing Reynolds Number

within the Piexiglas Injector

(g) _tP -" 33.8 psi., v - 15.57/s, Re d _ 3.08 x 104

(h) h,P - 36.3 psi., v = 16.14 m/s, Re d u 3.19 x 104 .............. 53

Figure 5 Comparison of Internal Flow Characteristics within the Plexiglas

Injector.

(a) v = 1.47 m/s, Re d - 2904,

(b) v = 2.08 m/s, Re a = 4122,

(c) v - 16.14 m/s, Re a = 31949 .............................

54
NOMENCLATURE

vi

a, b = Constants in Redlich-Kwong equation of state

f = Fugacity

N -- Total number of species

ni = Number of moles of species i

P = Pressure

R,, = Universal gas constant

r = Coordinate in the transverse direction

T = Temperature

u, v = Velocity components in x and r direction

V = Total volume

W = Molecular weight

x = Coordinate in the axial direction

X = Mole fraction

Y = Mass fraction

Z = Compressibility factor

Greek Symbols

p = Density

w = Acentric Factor

Superscripts

()°

(-)

Subscripts

( )i

()o

(),

( )+

= Ideal state

= Partial molar quantity

= Species i

= Boundary conditions of surrounding purge flows

= Surface value

= Boundary conditions of the LOX surface
Task No. 1

COMBUSTION OF LOX WITH H2(g) UNDER SUBCRITICAL_

CRITICAL_ AND SUPERCRITICAL CONDITIONS

K. K. Kuo, W. H. Hsieh,A. S.Yang, and J. J. Brown

Propulsion Engineering ResearchCenter

The PennsylvaniaState University

University Park, PA 16802

1. INTRODUCTION

This semi-annuM report summarizes the progress made during the period of

February 1 to October 15, 1992, under Task 1 of the contract entitled "Combustion

of LOX with H2(g) under Subcritical, Critical, and Supercritical Conditions."

The overall objective of this investigation is to achieve a better understanding of

the combustion processes of liquid oxygen and gaseous hydrogen under subcritical,

critical, and supercritical conditions. Specific objectives of the research program

are:

1. To determine the evaporation- and burning-rate characteristics of LOX in hydrogen/helium environments under broad ranges of operating conditions;

2. To measure species concentration profile and surface temperature of LOX

employing the gas chromatography and fine-wire thermocouples under nonreacting flow situations;

3. To perform a fugacity-based multicomponent thermodynamic phase equilibrium analysis for examining the high-pressure vapor-liquid equilibrium behavior at the liquid surface of LOX;

4. To formulate and solve a theoretical model for simulating the evaporation and

combustion processes in a LOX/H2/He system; and

5. To validate the theoretical model with the measured experimental data.
2. EXPERIMENTAL WORK

The scope of the experimental work performed during this reporting period

falls into two major areas: (1) evaporation tests; and (2) combustion tests. The

progress made in each subject area is discussed below.

2.1 Evaporation Tests

During the gasification and burning processes of L0X droplets inside a liquid

rocket engine at near-critical and supercritical conditions, the effects of solubility

of ambient gases in the liquid phase and variations of thermodynamic and transport properties are important. These effects can influence both evaporation and

combustion characteristics of the droplet. Therefore, prior to the development of

a complete combustion model, it is useful to examine the above effects under nonreacting situations. In the tests, the helium is employed to pressurize the system

instead of hydrogen for safety considerations. The other reason for using helium is

due to its very similar solubility features as hydrogen.

The progress made in evaporation tests includes:

• Calibration of electronic mass flowmeters.

• Design and installation of the gas sampling system.

• Determination of evaporation rates for LOX strand vaporizing in helium environments.

• Measurements of liquid surface temperature.

• Gas composition analysis using chromatography.

2.1.1 Calibration of Electronic Mass Flowmeters

In order to calibrate electronic mass flowmeters (OMEGA, FMA-874-V) used

for mass consumption rate measurements of oxygen and helium, a calibration setup
3

was devised as shown in Fig. 1.1. The source of the calibration medium was an

air bottle. The inlet and outlet pressure of mass flowmeters were maintained at

650 and 600 psig, respectively. The condition specified here corresponds the midpoint for the test range of pressure. An aluminum case diaphragm meter (American

Meter, AL-425), with +1% calibrated accuracy in the range of 5 to 100 percent rated

flow capacity, was installed in line with the electronic mass flowmeter to measure

the standard flow rate. Between the above two flowmeters, a line regulator was

mounted to reduce the gas pressure. Results given in Figs. 1.2 and 1.3 show that

the accuracy is ±2% full scale.

2.1.2 Surface Temperature and Oxygen Concentration Measurements

The surface temperature at the outlet of the LOX feeding tube was measured

using a 75 #m Fe/Cu-Ni (J-type) fine-wire thermocouple. To reduce conduction

errors, the thermocouple lead wires were placed along the liquid oxygen surface for

a distance of over 30 wire diameters.

For the setup of species measurement shown in Fig. 1.4, a quartz micro-probe

was used to sample the gas composition along the centerline above the LOX surface.

The probe was fabricated from a 3 mm diameter quartz tube tapered down to a 25

#m orifice. For each test, a gas sampling chamber was vacuumed and then filled

with the gas sample through the probe to a pressure level slightly above atmospheric

pressure (typically, 3 psig). Subsequently, the gas sample was withdrawn from the

gas sampling chamber through a vacuum pump and analyzed by a Varian model

3700 gas chromatograph (GC) with a thermal conductivity detector. In order to

make a reproducible amount of sample and injection time, an air-actuated 10-port

switching valve (Valco Instruments, Model A410UWT) was employed to control the

injection of the gas sample into the GC column. In this study, a Supelco Molecular

Sieve 5A, 45/60 mesh 3' x 1/8" stainless-steel column was installed in the GC. The
( " 650 psig)

?

AirBottle

Mass Row Indicator

Mass Flow I Line Regulator F

Sensor p_ Safety Relief " Valve (25 psig)

( ~ 600 psig)

To Exhaust

To Exhaust

Ak-425

Aluminum case

Diaphragm Meter

Figure 1.1 Schematic Diagram of the Apparatus Used for the Calibration

of Ma_ Flowmeters
5

100 , _ , , , , I I ' I ' I ' I ' 1 ' i

Q,..

._1

U3

,.u"

r.t0

_1

II

9O

8O

7O

6O

5O

4O

5O

2O

Serial No." 9100522

Gas Used for Calibration: Air

Inlet Pressure: 6.50 psig.

10

0 i , f,l,l,l, l,I;Irl,l,l

0 I0 20 30 ¢0 50 60 70 _0 go 100

PERCENTAGE OUTPUT, %

I0

Figure 1.2 Calibration Curve of the Mass Flowmeter Used for the Oxygen

Consumption Rate Measurement
6

u_

k-

<

rF

0

.-I

i,

110

IO0

9O

80

70

6O

5O

4-O

30

20

I0

0

w

I

0

I I I ' I I ' I i ' I ' I I

Serial No." 9100557 ._

Gas Used for Calibration: Air

Inlet Pressure: 650 psig /-_

I , I i I _ I , I , I a I , I J I _ I

i 0 20 ,30 4-0 50 (50 70 80 90 100

PERCENTAGE OUTPUT, %

r

110

Figure 1.3 Calibration Curve of the Mass Flowmeter Used for the Helium

Conaumption Rate Measurement
0

I--

_:> r.._)

0

°--

._® (a.)__.ll II _ E "_

!oo

'_- _ _ :..:-_.,:-,- _' :!it!;::::.i:-_- ::i;i::::i:i:.i:: ,--__ o

i :."_.'::, .:::-:::::::::::::.L ..I

-1-
helium was employed as the carrier gas for the Molecular Sieve 5A column with the

flow rare of 22 ml/min. In addition, following the same operation procedure described above, the GC was calibrated using known-concentration calibration gases.

Details of chromatographic analysis can be found in Yeh's (1991) work.

2.2 Combustion Tests

Prior to the use of H2/LOX system for combustion studies, a number of

CH4/GO2 firing tests were carried out at atmospheric pressures to ensure the proper

operation of the experimental setup. Figure 1.5 presents a picture of a stable

methane/oxygen gaseous diffusion flame. The gaseous oxygen velocity at center

tube was 0.5 m/sec with Reynolds number of 300. In addition, a combustion flame

of LOX with methane has also been successfully established in the test chamber, as

shown in Fig. 1.6. The current operational manual for evaporation tests has been

updated to include the consideration of combustion tests. A detailed discussion of

testing results will be given in the contract report.
9

02 N2/CH4

Figug¢ 1.5 Picture of a Methane/Oxygen Diffusion Flame (Re = 300)
10

He/CH4

Fi_ 1.$ Picture of • Mcthsae/LOX Dit[mdon Fl_m¢ (Re = 75)
1I

3. THEORETICAL WORK

The progress made in theoretical work is summarized as follows:

• Improvement of the high-pressure phase equilibrium model.

• Implementation of the numerical code for phase equilibrium calculations.

• Development of the computer code for flow analysis.

• Verification of the high-pressure phase equilibrium model.

• Verification of numerical solutions obtained from the flow analysis code.

• Cold flow simulations for studying the evaporation process in a LOX/He system.

3.1 Improvement of High-Pressure Phase Equilibrium Model

To obtain the species-concentration boundary conditions for each species immediately above the liquid surface, the thermodynamic phase equilibrium is imposed

upon the gas/liquid interface. At low pressures, the dissolution of ambient gases

in the liquid and the influence of pressure on thermophysical properties are negligible. The classical Raoult's law for ideal mixtures, therefore, can well define the

phase equilibrium condition. However, in the high-pressure environments, fugacitybased multicomponent thermodynamic analysis for vapor-liquid phase equilibrium

must be conducted to fully account for ambient gas solubility and real gas effects.

The high-pressure vapor-liquid equilibria for each component can be expressed as

follows:

= ') (1)

T ('_) = T (') (2)

P(_) --P(') (3)
12

where the superscripts (v) stands for vapor and (l) for liquid. The variable f,

is the fugacity for species i, and can be integrated through the following relation

(Prausnitz et M., 1986).

R,, T In( ) = _ T, V,,j RV T dV - R,,T InZ (4)

where Xi and V are mole fraction of species i and total volume, respectively. The

above equation indicates that fi is determined by the properties of the constituent

components, the concentrations in both phases, and the temperature and pressure

of the system. To compute the integral form in Eq. (4), the Redlich-Kwong (R-K)

equation of state is employed.

RuT a

P - (5)

- b _(_ + b)T °'5

where _ is the molar specific volume. The mixture parameters a, b in cubic-type

equation of state can be expressed in terms of composition and pure component

parameters.

N N N

° = F_, x,xjo,j, b= X,b, (6)

i_l j=l i=1

The constants ai/, bi are essentially dependent on the critical properties of each

species. In addition, the quantum-gas mixing rule corrections must be made in this

study to consider mixtures comprising one or more of the quantum gases (Prausnitz

etal., 1986).

2 25

0.42748R_,T_ i

aii = pc, ,

0"42748R_T_'_5 (7)

aij -- Pc,i '

0.08664RTc_

bi =

Pci
13

where

Kij = 1 -

_

cij

Tc,j =

1 + (cl/WijT)'

P_

Pc,j = 1 + (c2/WijT)'

T;*,.j = _(1 - gij),

Zc,i R. I':, i

p* _

cij - '

VCij

-1/3 /-1/3 -1/3

v¢,_ = t,vc, + vc_ )/2,

Z_,j = 0.291 - 0.04(wi + wj),

w,, - N + N /2,

8(_,%- )1/2

-1/3 -1/3 "3"

Vci "b Vci )

(8)

Note that the values of Cl and c2 are 21.8 and 44.2, respectively. The critical properties and acentric factors of pure components employed in the analysis are given

in Table 1.1 (Reid eta/., 1987). Because of the prevailing influence of the helium

on the liquid surface behavior, only binary system containing oxygen and helium

is considered in the phase equilibrium calculations. Moreover, instead of conventional critical constants, effective critical constants are required to implement the

quantum-gas mixing rules. Table 1.2 lists the values of effective critical constants

and acentric factor for the helium (Prausnitz et a/., 1986). Applying the R-K

equation of state to Eq. (4), we have

Infi=InXi+InP+ _b (Z-1)-In(Z-B)--_[ A 2 EN=I a Xjaij bi B (9)

which is valid for both vapor and liquid phase. With further manipulations of

equation of state, the mixture compressibility factor Z can be solved from the

following equation.
14

Table 1.1 Critical Constants and Acentric Factors of Pure Components

Tc (K) Pc (bar) _c (cm3/mol) Zc w

Oxygen 154.6 50.4 73.4 0.288 0.025

Hydrogen 33.2 13.0 65.1 0.306 -0.218

Helium 5.19 2.27 57.4 0.302 -0.365

Table 1.2 Effective Critical Constants and Acentric Factor of the

Helium for the Quantum-Gas Mixing Rules

Tc (K) Pc (bar) _c (cm3/mol) w

Helium 10.47 6.76 37.5 0.0
15

Z a - Z 2 + (A- B - B2)Z - AB = 0 (10)

where

aP bP

A - B - (11)

R2 TS /2 ' R_, T

In addition, analytic characterization of critical and supercritical evaporation

and combustion is required when the LOX surface reaches the critical mixing state.

In treating critical and supercritical processes at pressures greater than the critical

pressure of oxygen, the LOX surface is remained at its mixing critical state while the

regions below and above the surface are in the state of compressed liquid (T < Tcmv)

and supercritical vapor (T > Tc,nv), respectively. Intrinsically, the LOX exhibits a

smooth state transition from liquid to vapor. Furthermore, the vanishing surface

tension and the latent heat of vaporization are expected at the interface. The

liquid- and vapor-phase compositions are identical under the critical equilibrium

conditions.

Based on the above discussion, the key step in the numerical calculation is to

fix constant temperature and species-concentration boundary conditions using the

values of the mixing critical point once the LOX surface attains the critical state.

Thus, the interfaze boundary conditions are

(12)
16

3.2 Implementation of Numerical Code for Phase Equilibrium Model

For a two-component, two-phase system at a given temperature and pressure,

Eq. (1) provides two equilibrium equations. According to Eq. (9), the fugacity

can be expressed as follows (Lin, 1989).

Inf}V) =InX} v) +InP + bi (Z(V) - l)-In(Z (v) -B ('))

b(v) ',

a (v) [2 _-'7=' XJ '_)ao bi B(")'_

= lnX} t) + InP + bi (Z(O -1)-In(Z (t) - B (0)

b(t) '.

a (') 2 ET=, ""(')

B(') Aj aij bi In(1 B(t) [ a(') b(') ] + _-O J

Two more relations can be obtained from the following identities.

(13)

(14)

X} v) + X_V)= 1, X} 0 + X_/)= 1 (15)

It is noted that the unknowns in the above phase equilibrium equations are mole

fractions rY(v) _c('_) X} 0 and X_ t)) of each species. In this study, the NewtonRaphson iterative scheme was employed to solve the mole fraction distributions in

a vapor-liquid equilibrium system. In practice, the four unknowns were updated

simultaneously by their corresponding corrections during each iteration. To correlate the governing equations with unknowns, Eqs (13) through (15) were further

linearized. For Eq. (131, the linearized form is

2

/_,t n-_r 1

lnf? °' = Vio + E _kAX_V) (16/

k=l

where n + 1 represents the (n+l)th iteration.
17

Vio = In f} v)" (17a)

TV]k-- 10X_ '_) b, (Z(,) 10b (') b_ OZ(") 1 OZ (') OB (') x_) oxi_) b(_)_ )o-_ +b(o)oxi_) z(_)- B(o)(oxi_) oxi_))

1 OA (v)

B(") OX(k")

A ('J) OB (_) 2 _j _(v)

B(")' OX_ ") )( a(")"_J aO

B (')

b_)) In(i +-_-_]

B(")'_

× kB(v ) ] In

(17b)

And

( ZOO OB 0') Aj aij

B (,,) OZ (,,) (A(")_ (2 y'_j "_'(")

o E_Ej "(") "(")

.A i .Aj aij -- 2 E _ _: v'(v)

-- _ Vik'x j aij

OX_ ") i j

Ob(v) 0 _i X,(V) bi

oxi_,- oxi_, =E,_,,b,,

OA (') P Oa(")

ox_') n,.r_/_oxen)'

(18)

OB (v) P Ob(v)

oxi" - ,.r oxio,'
18

cOZ( ")

_ [(B (_) - Z(,,))OXi,, )OA(_) + (Z (v) + 2B'_)Z (_) + A (v)) _]

/(3Z(,,) 2 _ 2Z (,,) + A (_) _ B(,) _ B(.)2).

Similarly, for liquid phase

And

tl_n+l

In f_ " =

2

L,0+ _ L,,_,_xi') (19)

k=l

Lio = In f[ t)" (20a)

10X} t) bi (Z(O Ob (0 bi OZ (0 1 OZ (0 OB (0

L,_- x,(')oxl') b(,)' -_)o-_,)+b(,)oxl,) z(,)- B(,)(oxl,) oxl,))

A(O OB(t) _l 2 Y]_.i,,40 "_1 aij

B(t)' _Y)kt) ] ( a(O

B (0 "_

ox_') (') ( o_(')_ (oh(')_]

j -- a(i) ' \ ax_ ') ] + b-_ \ ax_ I) ] J

{A(O] (1 + Z(')] x \B(t)] In

(2Oh)

Z (t) ,_ OB (0

( Z(6 -+B(O ] ( 1

Z(O OX_ I)

B(') OZ(') ( A(O'_ 2E_ "(')

g(t)' OX_'))\ B(O] ( a(O'_J aij b(obi)

/._n+l

Subsequently, the equation of lnf) _' = In f} 0"+1 is imposed to meet the requirement of phase equilibrium. Thus,
19

Vio -4- Vil/_X_ v) -4- Vi2iX_ v) = iio 3v iiliX_ I) -_- Li2AX_ 1) (21)

The final equations in correction mode are given as follows.

+ Ax?) =0

Ax? + = o

(22)

v_Ax[ _ + v,_Ax__ - L_,AX_° - LI_AX_')= L_o- V,o

v_l_x}_)+ v_,x_ _)- L_,AX}_)- L_,X_')= L_o- V_o

In calculating the mole fractions of each species, the Gauss method was adopted to

solve the resulting nonlinear equations (Eq. (22)). Moreover, either Raoult's law or

experimental data can be employed to evaluate the initial mole fraction distribution

for the first iteration.

3.3 Verification of the High-Pressure Phase Equilibrium Model

With the afore-mentioned numerical implementation, calculations were conducted for the vapor-liquid equilibria of a binary system including oxygen and

helium. Figure 1.7 displays the high-pressure phase equilibrium results for the

oxygen/helium system. The predictions are in good agreement with the experimental data of Herring and Barrick (1965) over a wide range of pressure and temperature. Following the same calculation procedure, Fig. 1.8 exhibits a computed

O2/He phase equilibrium diagram at four different pressure levels (5, 30, 50, and

68 arm) which correspond the same testing pressure conditions for the companion

experimental works. It was observed that the mole fraction of helium in the liquid

oxygen (1.0 - mole fraction of LOX) increases with pressure for a fixed temperature. This tendency clearly indicates that the effect of the ambient gases solubility

in the liquid phase becomes more important at high pressures.
2O

160

140

Ld

c_

_< 120

rY

uJ

r_

LU

100

O2/He SYSTEM

Q, o, _ Experimental Data of Herring

and Barrick (1965)

Prediction

I00

50

P= 25 atm

8O

GAS PHASE LIQUID PHASE

, [ , I , I , l x

0.0 0.2 0.4 0.6 0.8

OXYGEN MOLE FRACTION

1.0

Figure 1.7 Comparison of Computed and Experimental High-Pressure

Pha_ Equilibria; Oxygen/Helium System
21

160

140

or"

E3

<<: 120

c_

uJ

O_

uJ

lOO

8O

O2/He SYSTEM

• CRITICAL MIXING POINT 68

50

30

P= 5 arm

GAS PHASE LIQUID PHASE

, I I I , I I 1 J

0.0 0.2 0.4 0.6 0.8 1.0

OXYGEN MOLE FRACTION

Figure 1.8 PhMe Equilibrium Calculations for the Oxygen/Helium System at

Various Pressures and Temperatures
22

3.4 Verification of Numerical Solutions from the Flow Analysis Code

The verification of the numerical code based on the SIMPLE scheme was carried

out by comparing the numerical result with the exact solution for the incompressible,

steady-state, fully-developed laminar pipe flow. The geometry of the calculated

domain was a circular pipe with a length of 20 m and a radius of 1 m. The flow

is forced through the pipe by certain pressure gradient. In the calculation, a 30 x

15 grid system has been adopted. The incoming axial flow velocity profile at the

inlet was set to be uniform with zero transverse velocity, while the fully-developed

condition was applied to the outlet. The assumption of negligible body forces and

constant fluid properties were employed in this test case.

Comparison between the calculated and exact solutions of the axial velocity

profile (at different radial locations) at the exit of a pipe is in good agreement, as

shown in Fig. 1.9. Figure 1.10 shows that the normalized mass residual converges to

10 -9 within 450 iterations. From the above test case, it is verified that the current

numerical scheme is capable of providing accurate solutions as well as reasonable

computational efficiency.
23

E

Z

._J

<

<

rr"

0

0

0

0

0

6

4

O

_ Exact Solution

B

o , , , I I , i , i I i , i ] f I i , , t f t , ,

0.00 0.05 0.10 0.15 0.20 0.25

AXIAL VELOCITY, u, m/s

Figure 1.9 Comparison of Exact Solution and Calculated Axial Velocity

Profzle in a Fully Developed Pipe Flow
24

_.J

¢w

t]l

-3 10

-6

10

-9

10

10 -12 i , , , , I , i , i I I I , , I I , ' ' i

O 1OO 200 300 4-00

NUMBER OF ITERATION

I f I

I

5OO

Figure 1.10 Convergence Rate of Normalized Mass Residual for Code

Verification
25

4. DISCUSSION OF RESULTS

The goal of the current investigation is to study the LOX evaporation behavior

under high-pressure conditions. Both experimental and theoretical results will be

presented and discussed in the following sections.

4.1 Experimental Results

The experimental determination of both evaporation rates and liquid surface

temperatures was conducted for LOX liquid strand vaporizing in helium environments. The chamber pressure range varied from 5 to 68 atmospheres. The flow

velocity and temperature of the ambient purge gas were maintained at 0.2 cm/sec

and at 291 K, respectively.

During the tests, an 150 mm C-mount telephoto lens with 50 mm extension

tube was installed in an industrial camera (Panasonic WV-3250) tomaximize the

viewing area around the opening of the LOX feeding tube (1/4" diameter). With

this equipment, the surface situations of liquid oxygen can be closely observed.

Generally, it was noticed that the LOX surface exhibits a slightly bulge shape at

low pressures. The bulged appearance tends to diminish as the pressure increases.

This observation was believed to be caused by the fact that the surface tension

becomes weak under high-pressure conditions.

Experimental measurements of steady liquid strand evaporation rates were carried out for a wide range of pressure, as shown in Fig. 1.11. In the pressure range

tested in this work, the change in chamber pressure shows a significant effect on the

LOX evaporation rate; the evaporation rate increases with raising pressure.

For the surface temperature measurements, the results for various pressure

conditions are presented in Fig. 1.12. The boiling point curve for pure oxygen

is also shown on the same figure for comparison purpose. In general, the liquid
26

lJ

!

E

_J

uJ

c_

Z

O

n,,

0.05

0.04

0.03

0

0.02

0.01

I !

02/He SYSTEM

I ' I ' I ' I I

i I l I i I , I i I i I I I

0 10 20 50 40 50 60 70

PRESSURE, arm

8O

Fi_rre 1.11 _ed Evaporation Rate vs. Pressure for LOX in Helium

F.mvironments at Room temperature

(Purge Flow Velocity = 0.2 cm/sec)
27

160 , , , , , , , j , ' ' I ' ' ' I ' ' ' I ' ' ' I ' ' ' I ' ' '

1,50

,y,

_< 130

_ 120

_ 110

100

gO

8O

0

Supercritica

J .,'"- I Regime

.,,'""" --- Bo ling Point for

'" Pure Oxygen

/ - -o- - Measured T_ versus P

/ Relationship for

02/He System

, , , t , , , l , , _ I , , , t = , , I , = , I _ _ _ I , = L

10 20 30 4-0 50 80 70

PRESSURE, atm

8O

Figure 1.12 _ured Surface Temperature vs. Pressure for LOX in Helium

Environments at Room temperature

(Purge Flow Velocity - 0.2 cm/sec)
28

surface temperature was observed to be slightly lower than the corresponding boiling

temperature of pure oxygen at subcritical pressures. In the supercritical regime, the

liquid surface temperature is very close to the LOX/He critical mixing temperature

calculated by the phase equilibrium model.

To study the mass-transport behavior of LOX during evaporation process, the

oxygen concentration profiles were measured using a gas chromatography. During

the gas sampling, the quartz microprobe was placed above the LOX surface and

aligned with the center line of the LOX feeding tube. Special care has been taken

to avoid disturbing the liquid surface by minimizing the sampling rates. The lower

limit of the sampling position was 0.85 mm above the LOX surface. When the

sampling position was lower than this limit, it was observed that the LOX surface

would be disturbed by the probe.

Figure 1.13 presents the center-line oxygen concentration distributions at subcritical (30 atm) and supercritical (68 atm) pressures. The data shown in this figure

represent averages of two to four test data. Typically, the values from the individual tests were within 6 % of the average. In the region near the LOX surface, the

decreasing slope of the oxygen mole fraction profile at 68 atmospheres was observed

to be steeper than that at 30 atmospheres. This phenomenon was attributed to

the reduction of convective and diffusive mass fluxes as the chamber pressure was

increased. In the downstream region (x > 20 mm), oxygen concentration profiles

were found to be almost flat.

4.2 Theoretical Results

To validate the theoretical model with the experimental data, calculations were

conducted to simulate the evaporation of LOX in a helium surrounding at a chamber

pressure of 30 atm. From measured mass flow rates of oxygen and helium, the
29

100

a, P = 68 atrn

o_ 80 - 0 P = 30 atm

6o

i,

o 40

2o

0

0 5 I0 15 2O 25 30

AXIAL DISTANCE, mm

Fi_e 1.13 Mmsured Center-line Oxygen Mole Fraction Profiles above the

LOX Surface
3O

inflow velocity abovethe LOX surfaceand the purge flow wasdeducedto be2.61 x

10-3 m/s and 2.0 x 10-3 m/s, respectively. The temperature measurementof the

LOX surfaceand the helium purge flow showed136.2K and 291K. In addition to

the known pressureand surfacetemperature as input conditions, the oxygen mass

fraction at the LOX surfacewas determined to be 0.981 through the vapor-liquid

equilibrium calculations. In this study, the computational domain involved an

axial length of 51.3mm and a radius of 38.1mm. The radius of the LOX feeding

tube was 3.175mm. The finite differencesolutions were performed using a 41 x

29 grid system. The non-uniform mesh sizewas employed,with both axially and

transversely clustered grids, to resolvesteepvariations of flow properties near the

liquid surfaceand wall. For eachcalculation, the massresidual of eachnumerical

grid convergedto 10-_; the conservationof massflow rate along the axial direction

was within 0.1%.

The calculated results are presentedin terms of temperature and oxygenmass

fraction contours, and the center-line oxygen mole concentration profile. Figures

1.14 and 1.15 present temperature and oxygen mass fraction contours with increments of 10 K and 0.05. Due to its diffusion-driven flow nature, both contours

display a mushroom cloud distribution. The closely spaced contours near the LOX

surface clearly indicate the presence of rapid variations of temperature and oxygen concentration. Figure 1.16 presents the measured oxygen mole fraction data

and the predicted center-line temperature and oxygen concentration profiles above

the LOX surface. The computed temperature distribution reveals that a significant

amount of thermal energy is transferred to the liquid phase for maintaining a steady

evaporation rate of LOX. The oxygen concentration profile was observed to drop

notably within an axial distance of 10 mm. The calculated result of oxygen mole

fraction was also compared with the experimental data. The comparison showed
31

\

\

P = 30 atm

AT = I0 K

280"

w

Z

Z

U.l

PURGE FLOW LOX

-140

Figure 1.14 Temperature Contour for the Evaporation of LOX in a Helium

Surrounding at a Pressure of 30 arm
32

\

\

\

\

P = 30 atm

/kYo= = 0.05

PURGE FLOW LOX

LU

Z

m

.J

rY

LU

FZ

i,i

L)

Figure 1.15 Oxygen Mass Fraction Contour for the Evaporation of LOX in

a Helium Surrounding at a Pressure of 30 arm
33

P = 30 arm

o Experimental Data

Prediction

300

-- 2O0 -4I"11

I"1"1

3>

--t

C

I'1"1

-- 100

0

0 10 20 ,.50 40 50 60

AXIAL DISTANCE, mm

Figure l.16 Mieasured Oxygen Mole Fraction Data and Predicted Center-line

T_nperature and Oxygen Concentration Profiles above

the LOX Surface
a reasonably good agreement.

34

5. OVERALL STATUS

The progress made during this reporting period is summarized as follows. Experimentally, evaporation rate and surface temperature were determined for LOX

liquid strand vaporizing in helium environments at pressures ranging from 5 to 68

arm. The gas chromatography analysis was also employed to profile the gas composition above the LOX surface. These data were used to validate the theoretical

model under cold-flow conditions. To date, combustion tests have been initiated.

Preliminary results showed the establishment of a steady LOX/CH4/He diffusion

flame above the LOX surface. Combustion test results will also be used for model

validation in the near future.

Theoretically, a comprehensive theoretical model, consisting of the conservation equations of mass, momentum, energy, and species concentrations for a multicomponent system, has been formulated and solved numerically to study the highpressure transport behavior associated with the vaporization process. Important

features including the solubility of ambient gases as well as variable thermophysical

properties have been considered in the model. In addition, a fugacity-based multicomponent thermodynamic phase equilibrium analysis with quantum-gas mixing

rules has been established and solved numerically to treat the ambient gases solubility at the interface of a LOX/He system. As a future work, the current computer

code will be further extended to simulate the LOX/H2(g) diffusion flame structure.
35

REFERENCES OF TASK 1

Herring, R. N. and Barrick, P. L. (1965). Advances in Cryogenic Engineering, 10,

151-159.

Lin, J. C. (1989). Thermodynamic and Numerical Analyses of Multicomponent

Fluid Phase Equilibria at High Pressures, M.S. Paper, The Pennsylvania State

University.

Prausnitz, J. M., Lichtenthaler, R. N., and de Azevedo, E. G. (1986). Molecular Thermodynamics of Fluid-Phase Equilibria, Prentice-Hall Inc., Englewood

Cliffs, New Jersey.

Reid, R. C., Prausnitz, J. M. and Poling, B. E. (1987). The Properties of Gases

and Liquids, 4th Ed., McGraw-Hill, New York.

Yeh, C. L. (1991). Measurement of Ignition and Surface Temperatures of Boron,

M.S. Paper, The Pennsylvania State University.
36

APPENDIX A

OPERATIONAL PROCEDURE FOR LOX/H2 COMBUSTION TESTS

OVERALL SETUP

Power System

1. Plug the two power strips on the test rig into a shielded extension cable and

plug the extension cable into the electric outlet.

2. Turn on the two power strips which are mounted on each side of the test setup.

This will activate a hydrogen leak detector and provide power to the solenoid

valve, the carrier demodulator for a Validyne pressure transducer, the mass

flow meters, and four thermocouple readouts.

Gas

°

.

,

Supply System

Connect one end of the helium purge line to the high pressure regulator on a

helium bottle and the other end to the inlet purge tube.

Connect one end of the hydrogen purge line to the high pressure regulator on

a hydrogen bottle and the other end to the hydrogen inlet tube.

Connect one end of the oxygen line to the oxygen bottle pressure regulator and

the other end to the oxygen inlet tube located on the test rig. The flow rate of

this line can be well controlled by a micro-metering valve. After introducing

gaseous oxygen into the system, the oxygen will be condensed as it passes

through the liquid nitrogen bath, and enters the test chamber as a liquid.

Liquid Nitrogen Supply System

Caution: Extreme care must be exercised in use of Hquid nitrogen. Due to

its cryogenic nature, LN2 will cause severe burns when in contact with exposed skin. It is recommended that the operator takes the appropriate safety

measures, namely, the use of gloves and goggles when handling liquid nitrogen

related hardware during the fili operation of the nitrogen bath.

1. Connect the flexible hose to the liquid nitrogen dewar and insert the dispensing

end into the nitrogen tank.

Electronic Control/Monitor Systems

1. Calibrate the Validyne pressure transducer and the carrier demodulator.

2. Connect the demodulator to the pressure transducer and mount the transducer

to the test chamber.
37

3. Connect one end of the BNC cable to the back of demodulator and the other

end to the IBM PC/AT computer.

4. Turn on the demodulator.

5. Attach the solenoid control cable to the computer.

6. Turn on the computer.

7. Logon to the pressure control program and perform a function check on the

solenoid valve.

8. Turn on oxygen deficiency monitor (the reading should be a value of 20.9).

9. Check hydrogen leak detector (a green light should be activated).

10. Connect each thermocouple to its respective readout if not already connected.

11. Turn on the mass flow meter power supply/readout box.

Camera

1. Position the video camera in front of the test chamber window.

2. Connect the camera to the VCR recorder and the recorder to a video monitor.

The monitor should be visible to an operator standing in front of the control

panel.

3. Insert a labeled video cassette into the recorder and turn on the recorder.

4. Turn on the camera and monitor.

5. Set the test information (i.e. title, date, test number, etc) on the picture for

future reference. This is done on the camera.

Control Panel

1. Prior to start, open the helium needle valve and leave the hydrogen needle

valve closed for evaporation tests.

2. Make sure all pressure gages read zero and thermocouple readouts indicate a

realistic room temperature.

3. Close the purge valve.

Ignition System

1. Install the nichrome-wire ignitor assembly and check its electric resistance (typically 3.5 ohm).

2. Set the switch of the power supply to OFF position and then connect ignitor

assembly to the power supply.
38

SETUP CHECKLIST

1. Both power strips are on.

2. Thermocouple readouts give temperature readings (TCs are working properly).

3. Oxygen, hydrogen and helium lines are securely connected to their inlets on

the setup.

4. Liquid nitrogen bath is ready to be filled in LN2.

5. Computer interface is complete; i.e. data lines, and BNC cables are connected.

6. Demodulator is on.

7. Mass flowmeter indicator is on.

8. Camera is in position and engaged.

9. Proper needle valve settings on the control panel have been chosen.

10. Ignitor is in place.

TEST PROCEDURE

1. Enter the pressure level into the IBM PC/AT pressure control program and

initiate the solenoid control sequence. The pressure level is entered as a predetermined demodulator voltage.

2. Briefly open the oxygen regulator and purge the oxygen feed line with a pressure

of 50 psig.

3. Close the purge valve.

4. Gradually fill the cooling tank with liquid nitrogen until the liquid reaches

a level of 1-2 inches from the top. During this operation the temperature

readouts for LOX inlets 1 and 2 should drop to -196.5°C. This provides a

means by which to judge whether the oxygen line is submerged by the liquid

nitrogen bath.

5. When the nitrogen bath is full, check the oxygen regulator to ensure the oxygen

bottle pressure exceeds that of the chamber pressure by about two times.

6. Open oxygen regulator and adjust the pressure until the desired mass flow

conditions are met. The micro-metering valve on the test setup should be

wide open to accelerate the condensation of 02 in the feeding tube.

7. Check the high pressure helium regulator to ensure the delivery pressure of the

helium exceeds the chamber pressure by about two times.

8. Pressurize the chamber and adjust the mass flow rate of He with the helium

needle valve located on the control panel.

9. Turn on the camera and begin recording.
39

10. When desired chamber pressure is achieved, and liquid oxygen is flowing from

the feeding tube, decrease the LOX mass flow by adjusting the micro-metering

valve on the oxygen inlet line until a stable, stationary level of LOX is maintained at the exit of the feeding tube.

11. Exit test cell and close the door. All personnel should be outside the cell

during ignition sequence.

12. Activate ignitor by turning on power supply. Adjust the output current of

power supply until nichrome wire is white hot.

13. Introduce hydrogen into the chamber (wait for 5 seconds).

14. If ignition is successful, turn off power supply, and begin data acquisition.

15. If no visible ignition occurs within 5 seconds, remain outside test cell, turn

ignitor off and close hydrogen supply while allowing test chamber to be purged

with helium for 2 minutes.

16. Make proper needle valveadjustments to increase hydrogen mass flow.

17. Repeat procedure 11 through 15 until successful ignition occurs.

SHUT-DOWN PROCEDURE

1. Close the hydrogen supply (flame will extinguish).

2. Close the oxygen supply and purge the test chamber with helium for 2 minutes.

3. Open the exhaust valve to depressurize the system.

4. Disconnect ignitor assembly from the power supply.

5. Turn off camera.

7. Stop the computer program for pressure control.

8. Turn off Demodulator.

9. Turn off mass flowmeter indicator.

10. Turn off oxygen deficiency monitor.

11. Turn off two power strips on the test rig.

12. Inspect and clean the chamber components.
40

Task No. 2

DENSE SPRAY AND MIXING OF LIQUID JETS EMANATING FROM

DOUBLET INJECTORS

IC K. Kuo, F. B. Cheung, R. D. Woodward and M. C. Kline

Propulsion Engineering Research Center

The Pennsylvania State University

University Park, PA 16802

.

t

INTRODUCTION

Observation of a like-on-like injector element in the near-injector region

performed in the previous phase of this project has identified the existence of a high

Reynolds number regime in which the pre-impingement jets are fully turbulent and

undergoing surface breakup. The new spray regime, which has not been observed

by previous investigators, is characterized by the presence of many fine droplets and

the disappearance of the well-defined liquid breakup wave pattern in the postimpingement region. It is speculated that a cavitating region may be present within

the orifice so that it could induce strong turbulence, leading to an onset of

atomization of the jets prior to impingement.

To further investigate the dense spray behavior of the impinging jets in the

high Reynolds number regime, experiments have been conducted using Plexiglas

injector components for direct internal flow observation. The main objective is to

determine under what conditions a cavitating region would form and whether or not

the cavitation is respons_le for the development of the high Reynolds number spray

regime. This part of the Semi-Annual Report describes the procedure and major

findings of the injector cavitation study.

PROCEDURE FOR INJECTOR CAVITATION STUDY

The major parameter considered in this investigation is the injection flow rate

or jet velocity. Figure 1 shows one setup for determining the mass flow rates. A

hand value is used to regulate the flow of liquid from the city-water supply line. The

pressure is measured using a pressure transducer connected to a PC-based data

acquisition system. A shutter is held over a pre-weighed collection bottle until the

flow is at the desired pressure drop. The sutter is then removed while a hand held
41

O

ICD

I,E

crj

r_

a)

0.

ft_

II

LL

c_O

o_

E

cl.

CI)

.0

.__

LL
42

stopwatch measures the time for collection. At the end of the duration, the shutter

is moved back into the original position and the time is recorded. The collection

bottle is dried with a towel and then the mass of the injected liquid is determined.

By knowing the amount of liquid collected, collection duration, pressure drop,

properties of the working fluid, and geometry of the injector orifice, the mass flow

rate, mean liquid exit velocity and discharge coefficient can be determined.

Alternatively, the mass flow rate or velocity at the exit of injector can be obtained

from the injection pressure by applying the existing pressure-velocity correlation

determined previously for the same injector.

The internal flow within the Plexiglas channel is observed while the behavior

of the flow, extent of cavitation, etc., is regulated by controlling the hand valve. This

is important since the characteristics of the resulting spray depend greatly on the flow

conditions before exiting. Specifically, separated flow induced cavitation can affect

the exit flow and possibly its instabilities. One major objective is to visualize the

internal flow and relate it to the characteristics of the external flow just after exiting.

Another major objective is to make quantitative measurements of the progression and

development of the cavitating flow.

The internal flow was visualized by one of two methods using the laser-sheet

setup: direct photography with the 35 mm camera or recording of the image to

either a S-VHS tape via a Dage/MTI video camera or directly to a Macintosh-based

processing system image. The light source for the photographs was either a

handlamp placed behind the injector system yielding a shadow graph-type image of

the cavitating region, or for better quality photographs the laser-sheet centered on

the Plexiglas injector.

3. DATA REDUCTION

The photographs are digitized by a scanner and transferred to the Macintosh

for proce,_ing. Images taken with the Dage/MTI camera can be directly stored to

the image processing system. Also, the images may be replayed in real time with the

S-VHS tape or viewed frame by frame with the image processor. This enables the

visualization or measurement of flow characteristics.

A digitized background image of the injector channel filled with water was

subtracted from the store images for each test run in which cavitation was observed.
43

This allows for the removal of any minor flaws in the Plexiglas-channel; otherwise

these flaws could inhibit certain portions of the field of view. Since the injector

channel was present in both photographs being subtracted, the channel must be

redrawn with a gray level corresponding to one that would be visible after the

processing of the image. A gray level of the cavitation region was chosen. The

isophote analysis was used to deduce the cavitation region. The cavitating region

appeared as a gray area approximately between 65 and 100 of the 256 gray scales.

Finally, the processed image contains mainly the orifice channel and the cavitation

region. Measurement of the extent of cavitation or reattachment location (x/L) was

then made.

4. RESULTS OF THE INJECTOR CAVITATION STUDY

The cavitation study consists of two parts. The first was a preliminary

assessment to determine if a cavitation region existed within the injector channel.

Tests were conducted with a Plexiglas channel similar in characteristics to the

previous injector only doubled in length and diameter for visualization purposes. The

second part was a study to relate the internal flow characteristics to the external flow

behavior. This was conducted using a Plexiglas injector with the same (original)

dimensions as the 60 degree metal injector previously studied.

4.1 Observation of the Cavitating Region

A photographic study was conducted to observe the internal flow of the

channel, using a handlamp to illuminate the Plexiglas test section from behind.

Figures 2 a to 2 e show the flow variation with Reynolds number, ranging

from Re d = 0 in Figure 2 a to Re d = 6.25 x 104 in Figure 2 e. There is no

flow, i.e., ZXP = 0, in Fig. 2 a. Figure 2 b (ZH' = 0.4 psi) depicts a laminar

flow throughout the channel and at the exit. Figure 2 c (_P - 3.7 psi) shows

the inlet to be free of separation (attached flow within the channel), a glassy

surface at the orifice exit. The fourth photograph in the series, Figure 2 d (ZIP

= 14.7 psi) shows a light gray region near the orifice inlet, signifying the start

or onset of cavitation has been reached. Figure 2 • (_LP = 30.05 psi) shows

a broad dark region along the lower wall of the channel. A gray region also

appears near the center portion of the port. The broadening of the dark

region and the appearance of the gray region signifies that the separation

region or cavitation region has developed to a fuller extent. The spray, upon
44

Figure 2, "

b°

Flow Variation with Increasing Reynolds Number Using the Plexiglas

Injector Channel.

(a) AP = 0.0 psi., v = 0.0 m/s, Re d = 0.0

(b) AP = 0.4 psi., v = 1.82 m/s, Re d = 7.21 xl03
45

Co

do

Figure 2 (Cont'd).

(c) AP = 3.7 psi., v = 5.54 m/s, Red== 2.19 xl04

(d) AP = 14.7 psi., v= 11.04 m/s, Red=, 4.37 xl04
46

e.

Figure 2 (Cont'd).

(e) AP = 30.05 psi., v = 15.79 m/s, Re d = 6.25 x104
47

I

i i i ) I ! i I | I 1 1 i 1 1 i

I

i

I _\ i i 1

........................ i----.-:-----_ .................................................. _........................

|

l ° I I "

'......................... L._..:._.......................... i........................

I I I

[,.\i

i.......................... [......................... T....... _°_ ......................... ! i ! ,

i _ ! . ! \

t i i \! o\

i i ! .'i.....................

..... , !,,, 1°,',,

I !

0

0 0 0 0 0 0

0

_t

0

O.

v

Q.

0

b--

O.

0

'r'-

Q

0

(.-

E

.C

.c_

0

O.

E

E

o

(II

rr

.c

0

.9

r_

IJ.

(41_ue7 ooUuO lo %) lu!od luew4oe_eel:l
4.2

48

exiting the orifice, appears gray and fuzzy as compared to the other

photographs.

To determine the reattachment point within the orifice, the laser sheet

is also used in a different series of tests to illuminate the cavitating flow within

the Plexiglas injector using the Dage/MTI camera with an output line

connected to the image processor and video cassette recorder. The result was

a stored frame or a motion picture of the cavitation region. The data were

processed by the background subtraction method mentioned earlier. The

results of the analysis are presented in Figure 3 along with the data for the

Plexiglas injector discussed in the next section. The data is plotted as a

function of the measured reattachment point in terms of x/L, versus injector

pressure drop. A linear relationship was obtained for the cavitation region

variation with pressure drop after the point of inception of cavitation.

The cavitation region is formed by the separation region at the channel

inlet. The external flow may not immediately be affected by the start of the

cavitation. Only when the pressure drop across the injector is beyond a

critical level, then the cavitation at the orifice inlet can grow enough in size

to influence the external flow. The effect of cavity growth is to increase the

turbulence level, which perturbs the flow into atomizing at the exit (jet

divergence and loss of glassy appearance).

Relationship Between the Internal and External Flow

After confirming the existence of the cavitating flow within the injector

channel of the scaled mode, the Plexiglas injector was fabricated. The flow

within the injector could then be related to the external flow. The injector

was studied by three methods: (1) visual observations, (2) still camera

photographs, and (3) video images recorded to the image processor.

The photographs and video images were taken with the same

procedure used to study the existence of cavitation region inside the injector

channel except the back-lighting technique was not used. The pertinent data

can be taken with the methods using the laser sheet as the illumination source

providing greater detail in the images. The photographs are similar to the

images of the injector channel. The point of inception of cavitation was within
49

__.1.0 psi for the two different Plexiglas injector units. The pressure for

inception was slightly higher for the injector unit. This implies that the

diameter effect was not an important factor for the inception of cavitation.

Logically, the pressure for cavitation to occur for this injector geometry

strongly depends on the turning angle of the flow. As the impingement angle

decreases, the pressure drop required for the flow to cavitate would increase.

Figure 4 shows the sequence of the progression of cavitation from a

low pressure drop case, zXP = 0.3 psi, to a high pressure case, Ap = 36.3 psi.

During the analysis using the Dage/MTI camera, an interesting flow appeared

within the orifice channel. Figure 5 shows three digitized images of the

injector orifice at a velocity of 1.47 m/s, (Re d = 2,904), 2.08 m/s (Re d =

4,122), and 16.14 rn/s (Re d -- 31,949). Unlike the cavitating regions seen in

previous figures, where the flow looked "fuzzy" or gassy, the flow within the

low velocity bubble was very clear (Fig. 5 b). This condition was a rare

occurrence. There may be a couple of suitable explanations, poss_ly, there

may have been a pocket of vapor (or air) trapped in the injector that found

its way to the low pressure separation region and remained there stably as

shown. Another possibility is that the bubble could actually be a cavitation

region that remains intact due to the low turbulence level of the liquid flowing

over the separation region. In line with this explanation, it seems reasonable

that there would be no turbulence enhancement due to cavitation. The

bubble did not appear to interfere or disrupt the downstream flow, whereas

the typical cavitating region of such extent definitely affected the external flow.

The images were processed in the same manner as before, and the

extent of cavitation was measured for different pressure drops. The results

were presented in Figure 3 along with the injector channel data. It was

observed that the start of the atomization/cavitation regime, corresponding to

the appearance of the diverging jet at the exit, occurs for the Plexiglas injector

when the cavitation region is between 20 to 30 percent of the length of the

orifice. The importance of the cavitation region on the external flow field

should not be neglected in this injector design.

The onset of cavitation for the 60 degree Plexiglas injector was seen to

be within the fully developed spray regime. Once the cavitation progressed

to 20 to 30 percent of the channel length, the exiting flow diverged. The
5O

a°

Figure

b.

Visualization of the Internal Flow with Increasing Reynolds Number

within the Plexiglas Injector.

(a) AP = 0.3 psi., v - 1.47 m/s, Re d ==2.90 xl03

(b) AP = 13.5 psi., v = 9.84 m/s, Re d = 1.95 x104
51

Co

d.

Figure 4 (Cont'd).

(c) AP = 19.8 psi., v = 11.92 m/s, Re d ==2.36 xlO 4

(d) AP = 23.8 psi., v = 13.07 m/s, Re d ,= 2.59 xl04
52

e=

Figure 4 (Cont'd).

(e) AP = 26.5 psi., v= 13.79 m/s, Red,= 2.73 xl04

(f) AP = 31.1 psi., v = 14.94 m/s, Re d ,=2.96 xl0 4
53

g=

hi

Figure 4 (Cont'd).

(g) AP = 33.8 psi., v= 15.57 m/s, Red,= 3.08 xl04

(h) AP = 36.3 psi., v = 16.14 m/s, Red== 3.19 xl04
54

,,E

0

Cr)

II

0

n"

II

vO)

x

n _

--C II

°_

-r"

• It

m II

I1)

0

¢.,

0

,ee_

E

0

E

U.
55

cavitation was then able to disrupt the flow by increasing the turbulence level of the

stream prior to the point of impingement, causing the distinct change in the resultant

spray pattern.

5. REMARKS

A straight unangled orifice (un-inclined with respect to inlet and exit plans)

exhibits the typical vena-contracta formation with an axisymmetric cavitation region

surrounding the contraction when the critical pressure drop is exceeded. This can

lead to hydraulic flip which is characterized by the cavitation vapor exiting the orifice

asymmetrically at a location that oscillates in time.

An orifice drilled at an angle to the inlet and exit planes (impingement half

angle) causes the flow to separate from the wall adjacent at the sharpest flow turning

angle. At sufficient pressure drops for cavitation to occur, the cavitation region

remains stably on the same side of the orifice passage without the oscillation

characteristic of hydraulic flip. The cavitation, after sufficient development, causes

the exiting jets to prematurely breakup before impingement. The premature breakup

could be beneficial, especially for a like-on-like injector, for a few reasons. A wider

more thoroughly atomized spray fan results after the impingement of the streams

compared to the case with no cavitation. Finer droplets are present and more evenly

distn'buted, thus allowing the possibility of increased mixing with a neighboring spray

fan. Also the well developed wave pattern, corresponding to a specific breakup

frequency, seen in the previously defined "fully-developed" spray regime, is broken

down by the introduction of turbulence at a wide range of frequencies. This loss of

a dominant characteristic frequency could be helpful in avoiding combustion

instabilities in liquid rocket engines.
The Pennsylvania State Uuivcrsily. in compliance v_,ilh li__dcral and M;./lC l_lw,'.;, is, committed to lhc policy thai all persons shall have

equal :.lcccs <-,t() pro_raltls, +idilli'-,sion, l.itld ctI/ploymcllt witilouI rc'_arc] to race, rcli,gioll, sex, national origin, hatMicap, a_c or status +is

a disabled or Vietnam era vclcr,tin, l)ircci all affirlllaiivc ilclioil itlquirics to the Af|imlativc Actiorl Office, 201 Willard P,tiildirig,

lhfiversityPark, PA 16802; (814)863-(i471.

U. l';d. F.NG 91 79