NBSIR 75-806

PHASE EQUILIBRIUM AND FLOW-INDUCED DESORPTION DATA

FOR He-CO, He-No 0, AND He-No SYSTEMS

W. G. Steward

R. 0. Voth

J. Hord

W. R. Parrish

C. F. Sindt

J. M. Arvidson

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

March 1975

Prepard for

Kirtland Air Force Base

Albuquerque, New Mexico 87117

NBSIR 75-806

PHASE EQUILIBRIUM AND FLOW-INDUCED DESORPTION DATA

FOR He-CO, He N 2 0, AND He-N 2 SYSTEM

W. G. Steward

R. 0. Voth

J. Hord

W. R. Parrish

C. F. Sin dt

J, M. Arvidson

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

March 1975

Prepared for

Kirtland Air Force Base

Albuquerque, New Mexico 87117

U.S. DEPARTMENT OF COMMERCE. Frederick B. Dent, Secretary

NATIONAL BUREAU OF STANDARDS Richard W Roberts Director

CONTENTS

Page

Nomenclature . vii

1. Introduction . . 1

2. Phase Equilibrium Experiments 2 2.1 Experimental Method 2 2.2 Results and Discussion 4

3. Scale Model Flow- Induced Desorption Tests 20

3.1 Scaling Criteria 20

3.2 Instrumentation 22

3.3 Test Procedure 22

3.4 Results and Discussion 23

3.4.1 Head Meter Flow Coefficient 24

3.4.2 Resistive Flow 25

3.4.3 Cavitating Venturi Flow . . . 28

3.5 Experimental Observations 42

4. Conclusions 42

5. References 44

Appendix Al 45

Appendix A2 49

LIST OF FIGURES

Figure Page

1. Pseudo Henry's Law constant for the He-CO system as a function of pressure 7

2. Pseudo Henry's Law constant for the He-CO system as

a function of temperature 8

3. Enhancement factor for the He-CO system as a function

of pressure 14

4. Enhancement factor for the He-CO system as a function

of temperature 15

iii
5. Pseudo Henry's Law constant for the He-^O system as a function of pressure . 16

6. Pseudo Henry's Law constant for the He-^O system as a function of temperature 17

7. Enhancement factor for the He-^O system as a function

of pressure 18

8. Enhancement factor for the He-^O system as a function

of temperature 19

9. Flow induced desorption test apparatus 21

10. Adaptation of Martinelli and Nelson [9] correlation for

two-phase flow resistance assuming no slip and both

phases turbulent 26

11. Axial pressure gradient as a function of local pressure for

He-N 0 flow rate of 759 g/s-cm

2 (1555 lb/s-ft2

) and

initial pressure of 75.8 bars (1100 psia) 27

12. Overall pressure drop. Comparison between equilibrium

flow of saturated ^0 and pure liquid 29

13. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrogen at 80.3 K (145°R) 31

14. Flow rate indication as a function of nozzle upstream

pressure for liquid carbon monoxide at 85.3 K (153°R) ... 32

15. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 255 K (459°R) 33

16. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 266 K (479°R) 34

17. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 244 K (439°R) .... 35

18. Cavitation parameter as a function of nozzle pressure

drop parameter for liquid nitrogen at 80.3 K (145°R) ... 37

19. Cavitation parameter as a function of nozzle pressure

drop parameter for liquid carbon monoxide at 85.3 K

(153°R) 38

20. Cavitation parameter as a function of nozzle pressure

drop parameter for liquid nitrous oxide at 255 K (459°R). . 39

iv
21. Cavitation parameter as a function of nozzle pressure

drop parameter for liquid nitrous oxide at 266 K (479 °R). . 40

22. Cavitation parameter as a function of nozzle pressure drop

parameter for liquid nitrous oxide at 244 K (439°R) .... 41

LIST OF TABLES

Table Page

1. He-CO liquid phase compositions 5

2. He-CO vapor phase compositions 6

3. He-^O liquid phase compositions 11-12

4. He-^O vapor phase compositions 13

5. Homogeneous equilibrium choking flow for He-^, initial

pressure 130 bars (1890 psia) . temperature

77 K (138°R) 30

6. Reduction of flow rate due to helium saturation 36

Al. Early test results - Nitrogen 46

Early test results - Carbon Monoxide 47-48

A2. Final test results - Carbon Monoxide 50

Final test results - Nitrogen 51

Final test results - Nitrous Oxide 52-54

v
ABSTRACT

Liquid-vapor equilibrium data were obtained for the helium-carbon

monoxide and helium-nitrous oxide systems at pressures to 138 bars.

Liquid and vapor phase compositions were measured at nominal temperatures

of 80, 85, 90, 100, and 120 K (144, 153, 162, 180, and 216°R) for the

helium-carbon monoxide system, and at 235, 245, 265, and 285 K (423,

441, 477, and 513°R) for the helium-nitrous oxide system. Internal con- sistency of the data was checked by using pseudo-Henry's Law constants

and enhancement factors.

The effects on flow of helium absorption and subsequent flow-induced desorption were investigated by means of reduced scale model

experiments. Friction losses attributable to helium desorption in the

long channels proved to be negligible both in experimental measurements

and in calculations based on assumed equilibrium of liquid and gas.

Contrarily, the assumption of phase equilibrium leads to gross miscalculations of flow rates in cavitating or near cavitating nozzles or Venturis. Actual venturi mass flow rates reached in the experiments

were fifty times the theoretical choking flow rates; however, definite

mass flow rate reductions due to helium desorption were measured,

ranging from four percent for nitrogen to twelve percent for nitrous

oxide. Pertinent experiences in handling these fluids and operating the

test equipment are also discussed.

Key words: Binary mixture; cavitating venturi; enhancement factors;

experimental vapor-liquid equilibria; flow induced desorption; heliumcarbon monoxide system; helium-nitrous oxide system; Henry's Law

constants; two phase choking; two phase flow.

vi
Nomenclature

Phase equilibrium — Section 2 E = Enhancement factor, dimensionless

.

H = Pseudo Henry's Law constant, bar.

P = Total pressure, bars.

P° = Vapor pressure, bars, x = Mole fraction in the liquid phase,

y = Mole fraction in the vapor phase.

Subscripts

1 = less volatile component.

2 = more volatile component.

Flow induced desorption — Section 3 = flow meter discharge coefficient.

= cavitating venturi discharge coefficient.

f = friction factor as though the total flow were gas. tg

f ^ = friction factor as though the total flow were liquid,

f = volume fraction of vapor,

v

G = mass flow per unit area.

G = limiting or choking value of G.

max

P = pressure.

P^ = supply pressure.

?2 = pressure upstream from the cavitating venturi.

P^ = pressure downstream from the cavitating venturi.

AP, = flow meter pressure differential.

4

P = saturation pressure,

sat

Re = Reynold's number,

v = specific volume.

W = mass flow rate,

x^ = axial length.

Y = mass fraction of vapor or gas.

vii
Nomenclature

Flow induced desorptlon — Section 3 (Continued)

P = two phase properties parameter.

0 = two phase flow pressure drop factor.

Pg = liquid density,

p = gas density,

(i = viscosity.

viii
PHASE EQUILIBRIUM AND FLOW-INDUCED DESORPTION

DATA FOR He-CO, He-No 0, and He-N„ SYSTEMS*

W. G. Steward, R. 0. Voth, J. Hord, W. R. Parrish,

C. F. Sindt, and J. M. Arvidson

1. Introduction

Helium gas pressurization often is preferable to the use of pumps

for transferring cryogenic liquids; however, there is concern that

helium, being soluble in these liquids, might be absorbed under pressure

and desorbed in the piping, thereby producing the high pressure drop and

choking characteristics of two-phase flow. These problems would be

anticipated if phase equilibrium prevailed throughout the flow system,

and this would probably disqualify helium pressurization as a practical

means of transferring liquids in many applications. Fortunately, perfect phase equilibrium is unlikely in dynamic systems; thus, the actual

amount of helium absorbed and desorbed is less than that expected if

equilibrium prevailed, particularly where the pressure drop is sudden

as in orifices or nozzles. Therefore, designers require not only phase

equilibrium data, but also experiments to determine the effects of flowinduced helium desorption in models of specific systems. The systems

of concern in this report involve helium pressurized liquid nitrogen,

liquid carbon monoxide, and liquid nitrous oxide; reliable phase equilibrium data are lacking for the last two of these fluids and flow induced

desorption data are unavailable for any of the fluids. These fluids are

of interest because they are candidate propellants for special purpose

gas generator systems.

In response to these requirements, this investigation supplies

needed data on phase equilibrium of helium with liquid carbon monoxide

and liquid nitrous oxide, and flow-induced desorption effects on all

three fluids of interest. The experimental program was carried out in

two phases

:

* This study was performed at the National Bureau of Standards, Boulder,

Colorado, under the sponsorship of Kirtland Air Force Base, Albuquerque,

New Mexico.
(1) Phase equilibrium experiments for helium-carbon

monoxide and helium-nitrous oxide systems.

(2) Scaled model tests to determine the effects of helium

gas absorption and desorption on pressure drop and

flow rates in pipelines and nozzles. These tests are divided into two categories : gradual pressure drop

due to friction in tubing and restrictions, and

sudden pressure drop in a flow nozzle.

2. Phase Equilibrium Experiments

Cryogenic systems containing helium have been of practical interest

for applications ranging from separation processes to pressurized systems.

However, few data exist in the literature for helium systems containing carbon monoxide or nitrous oxide. Sinor and Kurata fl] determined

liquid phase compositions of helium-carbon monoxide mixtures at four

temperatures between 77 and 128 K and at pressures up to 138 bars.

Kling [2] made an extensive experimental study of nitrous oxide systems

including the helium-nitrous oxide system; however, his data are presented in graphical form only, and the internal consistency of the data

could not be assessed. The work reported here was performed to provide

a more complete set of phase equilibrium data for these two binary

systems.

2.1 Experimental Method

The phase equilibrium apparatus used here is basically the vaporrecirculation system described by Duncan and Hiza T3]; therefore, only

pertinent details are given here.

Equilibrium pressures were measured by using either a double

revolution 0 to 20 bar Bourdon gauge or a 0 to 2000 psia Bourdon gauge.

The 0 to 20 bar gauge, calibrated against an air piston gauge, exhibited

a maximum negative deviation of 0.02 bars below 3 bars and a constant

negative deviation of 0.008 bars above 8 bars. The 2000 psia gauge,

calibrated against an oil dead weight gauge, has an estimated overall

uncertainty of 1.0 psi.

The fluid temperatures were measured with a platinum resistance

thermometer. Comparisons were made between readings of this thermometer

2
and temperatures obtained from vapor pressure measurements with pure

nitrogen and the vapor pressure data of Wagner [4]. These comparisons,

together with the specifications of the measuring instruments, indicate

the overall uncertainty of the temperature measurements is no greater

than ± 0.013 K.

Liquid phase samples were withdrawn from the bottom of the equilibrium cell through a stainless steel capillary; vapor samples were

isolated in the room temperature recirculation loop. To obtain reproducible analyses in the nitrous oxide system liquid samples were slowly

3 bled into a warm, evacuated 75 cm chamber and allowed to reach thermal

equilibrium (y 20 minutes) prior to gas analysis. The same procedure

was used for vapor sampling except that it was not necessary to wait

before analyzing the sample. In each case the sample chamber was filled

to about 2 bars gauge to allow for adequate purging of the gas analysis

equipment and to provide at least two analyses.

A gas chromatograph containing a thermal conductivity detector was used to determine mixture compositions. It was not possible to obtain

both liquid and vapor analyses during the same run because a different

carrier gas and column was required for each phase. For liquid phase

analyses argon was used as the carrier gas with a 6 m long molecular

sieve column. The chromatograph was calibrated against a 5.17 mole

percent helium in argon mixture. Helium carrier gas and a 0.5 m long

silica gel column was used for vapor phase analyses. In this case the

chromatograph was calibrated against pure carbon monoxide or nitrous

oxide, and the response in the peak area was assumed to be linear over

the composition range of interest. A previous calibration for nitrogen

verified the linearity assumption for this detector. In each case only

the peak area of one component and total sample pressure were measured.

We used a 0 to 75 psia quartz spiral Bourdon gauge to determine sample

pressure. The gauge was calibrated to better than 0.1 percent of the

reading using an air dead weight gauge. We estimate that the compositions are accurate to within ± 0.2 mole percent or ± 2 percent, whichever is greater; reproducibility, especially for the carbon monoxide

system, was much better.

3
The carbon monoxide was ultra pure (> 99.8%) and the nitrous oxide was of U.S. P. quality (> 98%) with the major impurities of both being

nitrogen and oxygen. No attempt was made to further purify the carbon

monoxide; however, after condensing nitrous oxide into the equilibrium

cell, the vapor space had to be vented three or four times before the vapor

pressure remained constant after venting. This venting also provided additional purification of the nitrous oxide.

Carbon monoxide poses a safety problem because it is toxic and not

readily detected since it is odorless, colorless and tasteless. To

prevent exposure of personnel to the gas, all vent lines were vented

directly outdoors. Before any carbon monoxide was used, the complete

system was leak-checked to 138 bars with pure helium. As a further precaution, CO detector badges were worn; these badges are reportedly

sensitive to carbon monoxide down to the ppm range.

Nitrous oxide is not considered toxic, but it is an anesthetic.

This fluid can explosively autodecompose at high temperature and/or

pressure [5]. However, at the operating conditions of this study the

only problem with nitrous oxide was that it caused swelling in the recirculation pump diaphragm. Thus, periodic shutdowns for pump overhaul

were required.

2.2 Results and Discussion

Tables 1 and 2 give liquid and vapor compositions as a function

of pressure for the five isotherms of the helium carbon monoxide system.

Measured vapor pressures are included because they differ from the

correlation of Hust and Stewart [6]. Figure 1 shows pseudo-Henry's Law

constants from our data and from Sinor and Kurata's data. The Henry's

Law constant is defined here as

H = (P - P°)/x

2

. (1)

Sinor and Kurata's data were reduced using vapor pressures from Hust

and Stewart. Figure 2 shows a cross-plot of Henry's Law constants as

a function of temperature for helium pseudo partial pressures (P - P?)

4
Table 1. He-CO liquid phase compositions.

Temperature

K °R

Pressure

79. 50 143. 1

bars

0. 787^

41 . 2 41

.

69. 69. 102.

103.

134.

136.

psia

11.41 = 597

602

1002

1007

1489

1500

1956

1974

Mole %

He

0.0

0.47

0.48

0. 77

0.79

1.15

1 .16

1.43

1 .46

84.71 152.5 1 .41 69. 0 69. 4 102. 9 103. 9 135. 3 136. 2

20.40 = 1000

1007

1493

1507

1962

1976

0 . 0

1 .05

1 .07

1 . 52

5285

86

90. 0 162.0 2. 37 = 35. 6 69. 4 103. 6 135.6

136.9

34. 35 = 516

1006

1 503

1967

1985

0.0

0. 72

1.41

2.03

2.53

2. 54

100. 0 180.0 5.41 = 37. 1 70. 1 104. 3 136.9

78. 5

= 538

1017

1513

1985

0. 0

1. 16

2. 27

3. 33

4. 19

120. 0 216.0 18.74=: 36. 5 69.9

105. 1 138. 0

271.

530

1014

1524

2002

0.0

1 . 74

4.78

7.67

10.03

Vapor pressure.

5
Table 2. He-CO vapor phase compositions.

Temperature Pressure Mole %

K °R bars psia CO

80.0 144.0 .830* 12.04* 100.0

11.9 172 7.96

13.7 198 6.97

28.8 417 3.64

42.3 614 2.68

56.3 817 2.13

69.8 1012 1.79

136.4 1978 1.15

84.71 152.5 1.41* 20.4* 100.0

6.9 100 12.91

7.0 101 12.87

15.2 220 10.54

28.8 418 6.05

42.8 621 4.37

56.4 818 3.51

70.9 1028 2.97

104.0 1508 2.28

90.0 162.0 2.37* 34.35* 100.0

14.6 211 18.73

26.7 387 10.82

40.1 582 7.61

55.6 806 5.87

71.2 1032 4.85

104.0 1509 3.72

137.6 1996 3.11

100.0 180.0 5.41* 78.5* 100.0

13.4 195 45.29

27. 4 398 24. 06

27. 9 405 24. 04

29.3 425 22.88

39.0 565 18.05

41.4 601 17.21

57.0 826 13.32

69.8 1012 11.37

70.2 1018 11.35

103.9 1507 8.46

131.0 1900 7.26

131.8 1912 7.25

132.7 1925 7.19

132.9 1928 7.18

120.0 216 18.74* 271.8* 100.0

29.3 425 75.99

42.5 616 59.47

56.7 822 49.18

70.5 1023 41.94

103.8 1505 32.65

137.1 1988 27.44

* Vapor pressure.

6
P-Pi » psi

0 20 40 60 80 100 120

P-Pj0

, bars

Figure 1. Pseudo Henry's Law constant for the He-CO system as a function of pressure.

7
TEMPERATURE, °R

70 80 90 100 110 120 130

TEMPERATURE, K Figure 2. Pseudo Henry's Law constant for the He-CO system as a function of

temperature.

8
of 0 and 120 bars. These figures show a fairly weak pressure dependence

for the Henry's Law constant.

A sensitive way to check the consistency of the vapor phase data

is to use the enhancement factor which is defined as

e= yjP/P°. (2)

The enhancement factor depends upon the nonidealities of both phases;

however, the liquid phase contains only small quantities of solute (He).

Therefore, vapor phase nonidealities are the most important contributions

to making the enhancement factor greater than unity. Figure 3 shows the

enhancement factor as a function of pressure for three of the helium

carbon monoxide isotherms. Figure 4 is a cross-plot showing the temperature dependence. The enhancement factor goes through a minimum between

90 and 100 K. The vapor phase data of Buzyna et al. [7] shows a similar

minimum in the enhancement factor for the helium-nitrogen system in the same temperature range. The existence of the minimum in the enhancement

factor is not fully understood. We do know that it is due to a combination of the slope of the vapor pressure curve, solvent vapor phase

nonidealities, and solute-solvent vapor phase nonidealities (He is

nearly a perfect gas at these temperatures and pressures)

.

Liquid and vapor data for the helium-nitrous oxide system are given

in tables 3 and 4. Figure 5 shows Henry's Law constants as a function

of pressure; within the scatter of the data the Henry's Law constants

are independent of pressure. Figure 6 shows a cross-plot of the Henry's

Law constant as a function of temperature. Comparisons were not made

with Kling's [2] data because the uncertainties involved in reading his

graphs make the comparisons meaningless.

Figure 7 shows the enhancement factors for the helium-nitrous oxide

system as a function of pressure for three of the isotherms and figure 8

gives the temperature dependence. This system shows no minimum in the

enhancement factor because the temperatures are too high. This conclusion
is based on a study of the enhancement factors for the helium-ethane

system (ethane and nitrous oxide have similar normal boiling points and

critical temperatures) which shows a minimum in the 160 to 190 K range

(8) . However, figure 8 shows a maximum in the enhancement factors at

the lower pressures. The maximum shown in the He-^O enhancement factor

curve is simpler to explain. It is due to the vapor pressure of ^0

rising more rapidly with temperature than the vapor phase solubility.

This also explains the disappearance of the maximum at higher pressures.

10
Table 3. He-^O liquid phase compositions

Temperature

K °R bars

Pressure Mole To He

195. 0 351 1.82*

103. 3 103. 6 136.4

137.8

26,

1498

1502

1978

1998

40* 0.0

0. 52

0. 52

0. 70

0.68

215. 0 387.0 4. 70*

49.9

51.8

86.

1

68. 17 = 724

751

1248

0.0

0.-29

0. 27

0. 54

235. 0 423.0 10.13*

51.2

51.6

84.8

86.2

136. 5

146. 92* 742

749

1230

1250

1980

0.0

0. 56

0. 55

0. 96

1 . 04

1.70

245. 0 441 . 0 14. 14*

71.0

103. 2 103. 7 131 . 3 133. 9 137. 6 137. 8

205. 08-- 1030

1497

1504

1904

1942

1996

1998

0.0

0. 98

1.57

1 . 5504

05

09

09

255. 0 459.0 19.22*

38. 8 38. 9 58. 2 58.3

75.6

76.0

86.

1

86. 7 87.4

101.8

103.0

103. 2 104. 5 104.7

121 . 2 121.7

133.2

135.4

136.2

278. 76- 563' 564

844

846

1096

1102

1248

1258

1267

1476

1494

1497

1515

1518

1758

1765

1932

1964

1975

0.0

0 . 42

0 .41

0. 87

0. 90

1 . 27

1 . 30

1 .45

1 .47

1. 54

86

82

1 . 80

1 . 84

1. 8509

14

4146

42

* Vapor pressure

11
Table 3. He-N-0 liquid phase compositions (continued).

Temperature Pressure Mole %

K °R bars psia He

265 0 477.0 25.38* 368* 0.0

52.5 762 0.77

52.8 766 0.75

71.6 1038 1.27

71.7 1040 1.28

72.1 1045 1.22

104.9 1522 2.15

105.5 1530 2.22

135.4 1964 2.91

136.8 1984 2.99

285 0 513.0 42. 18* 612* 0.0

81.8 1186 1.73

94.3 1367 2.38

94.6 1372 2.33

114.6 1662 1.77

116.9 1695 3.34

135.5 1965 4.14

136.1 1974 4.16

136.7 1983 4.18

Vapor pressure

12
Table 4. He-N^C vapor phase compositions,

1 emperature

K °R

?res sure

235. 0

245. 0

423.0

441 . 0

bars

10. 13*

19.7

35.6

52. 0 69.3

104. 5 132. 9

14. 14*

40.3

51.7

67. 6 98. 2 125. 9

psia

146.92=; 285

516

754

1005

1515

1928

205. 08* 585

750

980

1424

1826

Mole %

X2°

100. 0 54. 37

31.45

22.25

17.31

11.81

9.45

100.0

39.93

32.20

25. 30

17.89

14.42

255. 0 459. 0 19. 22* 50. 0 70.1

103.8

128. 5

278

739

1017

1506

1863

76* 100. 0 43. 89

33.72

23. 70

19. 58

265. 0 477.0 25. 6 41.6

72.5

98.3

99. 6 106. 0 136. 5

372= 603

1052

1426

1444

1538

: —

:

100.0

68.72

43.26

33.72

33. 30

31. 72

25. 97

285. 0 513. 0 Unable to obtain vapor phase data at this temperature due to rapid failure of circulating pump

diaphragms. (Natural rubber and butyl swelled

and PTFE, polypyromelitirri.de and a flou.ro- elastomeric material failed to maintain suitable

elastic properties).

Vapor pressure

13
PRESSURE, psi

0 20 40 60 80 100 120 140

PRESSURE, bars

Figure 3. Enhancement factor for the He-CO system as a function of pressure.

14
140

TEMPERATURE, °R

160 180

T

70 bars (1015 psi)

40 bars (580 psi)

1.2

1.0

70 80 90 100 110

TEMPERATURE, K

Figure 4. Enhancement factor for the He-CO system as a functof temperature.

15
5

4

3 -

2 -

t io

4

£ 8

D.

I

4 -

3 -

10"

500

P-Pi , psi

1000

Sr

20

1500 2000

195K (351 °R)

215K (387°R)

XT

O

235K (423°R)

-o

245K (441 °R) - 9 cP^-

1

-i

<3> &

255K (459 °R)

CP3-#

265K (477°R)

285K (513°R) —O

40 60 80 100 120

P-P]

0

, bars

6 ^ 5

r- 3

- 2

n

CL

CM X

10

b

140

Figure 5. Pseudo Henry's Law constant for the He-N

2 0 system as a function of pressure,

16
<r 10

to -0

CN u X

QI ^ 6

ii

2 -

10"

360 T

380

"1"

TEMPERATURE, ° R 400 420 440 460 480 500 520

O 20 bars (290 psi

)

A 120 bars (1740 psi)

J L 200 220 240 260

TEMPERATURE, K

280

1

- ' 6

- 4

CL

CM X

10 5 ^.

Q_

I

8 ii

Figure 6. Pseudo Henry's Law constant for the He-N^O system as a function of temperature.

17
P, psi

0 500 1000 1500 2000

pressure.

18
Figure 8. Enhancement factor for the He-N20 system as a function of

temperature

.

19
3. Scale Model Flow-Induced Desorptlon Tests

3.1 Scaling Criteria

Figure 9 is a schematic showing the final configuration of the

apparatus used at the National Bureau of Standards to measure the effect

of helium desorption on flow rate. The apparatus models those features

of the full scale systems which pertain to possible two-phase pipeline

flow resistance and two-phase choking in a nozzle. Due to the temperature, pressure, and time dependence of gas solubility the essential

characteristics of the full scale system to be duplicated in the model

are the temperatures, pressure drops, and liquid residence times. A

6.7 m (22 ft.) length of 3.2 mm OD (1/8 in.) x 0.76 mm (0.030 in.)

wall copper tubing provided the required pressure drop and residence

time for the first series of liquid nitrogen tests. Table Al in

Appendix A shows the results of the first series of tests performed

with liquid nitrogen. Runs 7 through 10 indicated that the effects of

absorbed helium on flow were almost negligible in scaled long tubes.

Consequently, we turned our attention to desorption effects in cavitating Venturis (or flow control nozzles). The pressure drop in long tubes

is gradual, but in nozzles it is quite rapid; therefore, the desorption

potential and two-phase flow characteristics could be expected to differ,

A venturi was installed along with a 91 cm (36 inch) length of 6.35 mm

OD (1/4 in.) x 1.24 mm (0.049 in.) wall tube to provide the proper

residence time; the 6.7 m x 0.76 mm tube was removed and variable

control of valve V5 provided the proper resistance. The purpose of this

change was to allow pressure upstream from the venturi to be varied

during a test. A second throttle valve (valve V23) was also added to

permit adjustment of the pressure downstream from the venturi.

To expedite the testing, the model supply and receiver reservoirs

were sized to be compatible with existing dewars, catch tank, and plant

facilities, and to give reasonable run times. The volume of the supply

3

and receiver reservoirs is 0.303 & (0.107 ft ).

20
21
3.2 Instrumentation

Temperatures at the bottom of the supply cylinder (T^ in figure 9)

,

in the stabilizing bath, T^, and in the receiver cylinder, T^, were

sensed by Chromel-Alumel thermocouples and read out on a digital voltmeter. Pressure transducers—P^ at the supply inlet, at the restrictive tubing inlet, P_ at the restrictive tubing outlet, and AP^ , the

flowmeter differential—were calibrated against secondary standard gauges

before each test run. Pressures were recorded continuously during

each test on a multichannel light beam galvanometer recorder.

The final pressure and temperature in the volume-calibrated catch

tank (initially evacuated) determined the total mass of mixture transferred

during the test. Pressure in the catch tank was measured with a quartz

bourdon tube pressure gauge of 0-1000 mm Hg range, and temperature was measured with mercury thermometers in the catch tank. The instantaneous

flowrate was indicated by a head meter at the discharge of the supply

cylinder. The meter was calibrated by comparison with the catch tank

integrated mass flow for several steady state runs of known duration

(see sec. 3.4.1). The overall uncertainty of test measurements is as follows

:

1. Temperature (thermocouples ) — ± 1.0°R (± 0.6 K)

2. Mass flow rate (head meter) — ± 1.5% variation

about stated C

d

for CO, N

2

, and N

2

0

3. Pressure (transducers) — ± 2.0% of the lowest readings

for P^, V^' and P3 (better for higher readings)

4. Pressure (quartz bourdon tube) — ± 0.45 mm Hg

5. Flow meter differential pressure (transducer) — ± 1.5%

of lowest readings for AP

6. Composition analysis — ± 0.05 mole %.

3.3 Test Procedure

The supply cylinder was filled with liquid by simply condensing the

appropriate gas. Bath temperature was maintained by 1) controlling the

liquid nitrogen bath pressure for CO and test fluids, and 2) circulating chilled ethylene glycol for ^0 test fluid. Bath temperature control

was maintained within 1 0.6 K for all experiments.

22
Flow tests were carried out with liquid nitrogen, liquid carbon monoxide, and liquid nitrous oxide. To establish the effects of dissolved

helium, each test under a given set of flow conditions was performed with

the supply liquid unsaturated and then saturated with helium gas. Saturation was accomplished by recirculating helium gas by means of a small

magnetically driven pump shown in figure 9. The helium gas was introduced

at the bottom of the supply cylinder and allowed to bubble up through the

liquid for approximately forty-five minutes before each test with saturated

fluid. The degree of saturation was verified after each test by mass spectrometer analysis of gas samples taken from the catch tank. For

unsaturated test runs the absorption of helium was kept to a minimum by

pressurizing at the top of the supply cylinder and starting the run immediately. Further preparation for a test run consisted of calibrating

the pressure transducers, evacuating the catch tank, purging the downstream tubing, and pressurizing the receiver to the required back pressure.

The purpose of this pre-pressurization was to minimize the time needed

to establish steady flow. The purpose of the receiver cylinder is to

maintain a constant back pressure during the flow desorption tests.

Heating of the receiver cylinder prevented two-phase flow through the

back pressure regulator and minimized the mass of residual fluid remaining

in the receiver cylinder following a test.

Each test was started by opening valve V5. In some cases a steady

flow was maintained throughout the entire test consuming most of the

volume of the supply cylinder. In other "burst tests" the flow was maintained only long enough to establish steady flow conditions. Operating parameters such as upstream or downstream valve positions could then

be changed and another test made without refilling the supply cylinder.

In this way a series of runs were made in rapid succession. Repetitive

tests were performed to verify the consistency of these data.

3.4 Results and Discussion

Early test results (before run 41) are slightly inconsistent showing

the consequences of undeveloped procedures. Also, lower saturation pressures were specified by the sponsor for succeeding tests. The early test

results are tabulated for completeness in the first part of Appendix A

(Table Al) ; however, they have not been plotted or processed further.

23
Instrumentation precision, calibration, and test procedures were revised

and lower saturation pressures were used after run 41. Only these later

results are discussed in this section and the data are tabulated in

Table A2.

3.4.1 Head Meter Flow Coefficient

A converging nozzle leading from the supply cylinder to the test

section inlet tubing served as the primary element of a flow head meter.

The square root of the pressure differential between the supply cylinder

and the flow nozzle throat, /APT, is proportional to the mass flow rate

W. In the analysis of the data in this report /AP^ is used rather than

W; however, a means of determining W from tabulated values of AP^ in

Appendix A is given in this section for future reference. The flow

meter was calibrated by means of seven steady state nitrogen runs and

four nitrous oxide runs. To facilitate the determination of mass flow

rate from values of AP^ for all three fluids, a discharge coefficient,

C^, has been calculated. In this case the discharge coefficient has been

defined in the most useful, simplified form as

C

d = W7/AP4

p £

. (3.1)

The discharge coefficient is a function of the Reynolds number which,

for a given flow meter, is proportional to W/u^.

The values of C. for N„ as calculated from (3.1) were corrected

d I upward by 0.8% for the flow meter throat cross sectional area reduction

due to the lower temperature of N^. Over the range of W/u^ covered in

these experiments was repeatable for each fluid to within 1.5% of the

average value. However, the average value differed for the two fluids

as follows: , \

~ 20 - 5 \ ' v

bars ' t3 <= 1 - 55(10)

"3

? Mtj 7J ) for N

2 lb

1 m ft

3

lb lb

in

2 ft3 c

d = 22 -° f /y^5

* 3

(= 1 - 61(10)

s \/7T 73 cm V m ft

} for V

It is believed that this discrepancy is due primarily to un- certainty in the density of the liquid ^0. For this reason, and

because of the similarity between ^ and CO properties, it is recom- mended that the value of C, obtained for N0 be used for CO.

d Z 24
3.4.2 Resistive Flow

To analyze the worst case, it is assumed that helium gas is absorbed to the maximum (equilibrium) concentration under supply pressure

and undergoes equilibrium desorption at the reduced pressure downstream

in the system. Desorbed vapor volume and mass fraction of vapor were

computed from the phase equilibrium data of section 2. Pressure drop

due to two phase resistive flow was calculated by means of a correlation of Martinelli and Nelson \9]. An adaptation of this correlation

was employed which is usable over the entire two phase range and reduces

properly to single phase values for pure liquid or pure gas. The twophase flow parameters used are shown in figure 10.

The pipeline axial pressure gradient is given by

f V1 * 8f (1 - Y)

1 * 8

1 2 -r~r = 9 — H -^r > where the

dx' L P g

Pg J 2D '

symbols are identified in the nomenclature of this paper.

The pressure gradient for a given flow is a function of the mass fraction of vapor which is, in turn, a function of the local pressure,

The relationship of dP/dx to P is given in Figure 11 for a model

He-^O flow rate per unit area of

G= 759 ^_ ( 1356 —*M s • cm \ s • it /

and a supply liquid pressure,

P

1 = 75.8 bars (1100 psia).

Figure 11 shows that, in proceeding from high pressure pure

liquid to lower pressures along the equilibrium two-phase flow curve,

the pressure gradient does not rise steeply until the pressure falls

below about 27.6 bars (400 psia). Pressure drop per unit length for

the nearly incompressible unsaturated liquid flow varies little with

25
Figure 10. Adaptation of Martinelli and Nelson [9] correlation

for two-phase flow resistance assuming no slip and both

phases turbulent.

26
LiiD/saeq < X P/dP o co us coo o

*3- O

C\J o oooo

01ooco EuI 00Q>

cr> LO

hoo

<v

-(-> ros-oc • z: s

re

s_

O 0 00 4-

<u c S- 0 3 t—1

00 t—

O)

O 00 c s_ WO IT3 <T3 *r~ _Q

00 ro CL cc O

LD 0 <+- 0 O

LD

O

CO 0}

s_3 u 00 -

00 0 3 (U

i03 CL

I/) fO

(T3 +->

4-> 1

C 0 0) •I" 0 • r— m T3 -0

(0

r

s- <n CD

QJ (N

S- +J 3 4- 0 00 1 c 00 00 CM 0)

5- JD

Q. LD

T3 LO • 

LD

OX

lo ^- ro UL/LSd 'XP/dP

ID-3

2"
pressure. Since the minimum pressure, P^, in the resistive section did

not fall below 44.8 bars (650 psia) , one would not expect to see a difference between saturated and unsaturated liquids in these resistive . flow

experiments

.

Figure 12, giving overall pressure drop by integration of the

gradient, illustrates this point. To make the comparison it was necessary

to use valve 5 stem position as an independent variable since that

valve was the main resistive element and its setting varied. The ordinate

*2

of this curve is (P.^ - P

2

)/AP^, which is proportional to (P

1 - P

2

)/W , and

amounts to a flow coefficient for the resistive plumbing. This coefficient should depend upon the valve position and two-phase frictional

effects. The solid line is a smooth fit of all the unsaturated test data

obtained in runs 54 and 55. With the solid line taken as a base, corres- ponding theoretical saturated liquid-equilibrium flow values were calculated and plotted as the dashed line. It is apparent that the largest

estimated difference between the saturated flow and unsaturated flow

curves is probably within experimental precision; this is born out by

the experimental data.

3.4.3 Cavitating Venturi Flow

Pressures in the cavitating venturi are intended to drop to the

vapor pressure at the throat. As shown in section 3.4.2 the two-phase

flow resistive effects would become large at such low pressures if

phase equilibrium prevailed. Even more restrictive would be the twophase choking effect. For homogeneous, equilibrium flow the choking mass velocity is given by

C

a units conversion factor,

isentropic pressure derivative of mixture specific volume.

28
0 .2 .4 .6 .8 1.0 1.2 1.4

VALVE POSITION, ( Turns open)

Figure 12. Over all pressure ckop. Comparison between equilibrium flow

of saturated N

2 0 and pure liquid.

29
Table 5 shows the relationship between G and pressure computed from

the He-^ phase equilibrium data for an initial pressure of 130 bars

(1890 psia) and temperature of 77 K (138°R)

.

Table 5. Homogeneous equilibrium choking flow for

He-N , initial pressure 130 bars (1890 psia),

temperature 77 K (138°R)

.

p P G

TO3.X lb/s • ft

G

UlaX

g/s • i psia bars

60 4.14 670 327

100 6.89 3,400 1,660

200 13.8 11,500 5,610

300 20.7 19,000 9,280

500 34.5 32,000 15,600

800 55.2 50,000 24,400

1000 68.9 60,000 29,300

1500 103.4 84,000 41,000

1890 130.3 103,000 50,300

2

Experimental saturated nitrogen flow under these initial conditions

2 reached 21.8 g/s (0.048 lb/s), which results in G = 15,600 g/s • cm

2 (32,000 lb/s • ft ) in the venturi throat. Bernoulli's equation predicts

1.38 bars (20 psia) at the throat for pure liquid flow at this flow

2

rate, whereas table 5 shows that G < 670 lb/s • ft at 20 psia, or max

one fiftieth of the actual mass velocity. This calculation merely shows

that the phase equilibrium assumption results in a gross miscalculation

of venturi flow.

The experimental data for all three fluids do indicate a small

effect of flow-induced helium desorption in the venturi, however. In

r igures 13-17 the /APT — which, except for a constant, is equivalent to

W — is plotted against P^ the venturi upstream pressure. The maximum

reduction in flowrate due to helium saturation is shown in table 6.

30
? 2

, PSia

1000 1200 1400 1600 1800 2000 2200

4.5

4.0

"5 3.5

CL

CL

^ 3.0

Q

2.5

2.0

1.5

60

cP

O Run #43 - Saturated at 124 bars (1800 psia)

Run #44 - Unsaturated

Solid symbols indicate throttling downstream

1.2

1.1

1.0

0.9

<

0.8 5

0.7

0.6

0.5

- 0.4

80 100 120 140 160

P

2

, bars

Figure 13. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrogen at 80.3K ( 145 0 R )

.

31
1000

P

2

, psia

1200 1400 1600 1800 2000 2200

4.5

4.0

Cl oT 3.5

<

C—

)

o

3.0

2.5

2.0 -

1.5

T

60

Figure 14.

Ia

cP o °d

q

0

<9

o

O Run #42 - Saturated at 122 bars (1770 psia)

Run #41 - Unsaturated

Solid symbols indicate throttling downstream

70 80 90

P2

, bars

100

- 1.2

- 1.1

1.0

CO

fO -Q

0.9 1

i—

i

Q

0.8 g

0.7

0.6

0.5

0.4

110

Flow rate indication as a function of nozzle upstream

pressure for liquid carbon monoxide at 85. 3K (153°R).

32
P 2 > psi

a

900 1000 1100 1200 1300 1400 1500

4.5

4.0

3.5

on

5 3.0

CP S

a o

_<$>

o

_iu_

2.5

2.0

O Run =48 - Saturated at 76 bars (1100 ps1a)

Run =46 - Unsaturated

Solid symbols indicate throttling downstream

1.5

60 - 0.4

70 80 90

p 2

, bars

100 110

Figure 15. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 255K (459°R).

33
P2

> psia

900 1000 1100 1200 1300 1400 1500

4.5

4.0

3.5

1.2

O Run #53,57,59 - Saturated at 76 bars (1100 psia)

Run #52,56,58,60 - Unsaturated

Solid symbols indicate throttling downstream ~ 1.1

10Q.

5 3.0

c_>

2.5

2.0

1.5

CD

60 70 80 90

P 2

> bars

- 1.0

0.

S-

- 0.8

- 0.7 t-j0.6

0.5

-d 0.4

100 110

Figure 16. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 266K (479°R).'

34
4.5

4.0

3.5

00

O- ,<

3.0

2.5

2.0

1.5

900 1000 1100

P

2

» PSia

1200 1300 1400 1500

Run #54 - Saturated at 80 bars (1160 psia)

Run #55 - Unsaturated

Solid symbols indicate throttling downstream

m

cP

a*

o 8°

^

60 70 80 90 100

P

2

. bars

1.2

1.1

1.0

0.9

0.8

Q

2:

0.7

0.6

0.5

- 0.4

110

Figure 17. Flow rate indication as a function of nozzle upstream

pressure for liquid nitrous oxide at 244K (439°R).

35
Table 6. Reduction of flow rate due to helium saturation.

Fluid Runs No

.

K

Nominal

T - sat

K

Maximum Flow Rate

Reduction, %

4J, 44 on o oU . L

/, 4 41 , 4Z ico Ijj Q C T. OJ . J 4

40 , 40 45y 0 c. c 9

N 0

2 52 ^3 56-60 A80 H O VJ N

2 0 54,55 439 244. 9

Experimental data points involving cavitation can best be distinguished from noncavitating data points in figures 18-22. The rationale

behind the coordinates used in these figures is as follows:

Non-cavitating flowrate through the venturi is approximately proportional to the square root of the pressure drop across the venturi,

w <* /p

2 -p

3

, and for the flow meter,

W cc /ap . 4

In non-cavitating flow a reduction in the pressure P^ downstream from the

venturi (by opening valve V23) increases increases the flowrate, and

correspondingly increases /AP^ . Normalization of the quantities /P^-P^ and

/AP^ by d ividing each of them by /P 0 -P does not change this trend; i.e.,

/VP

3 for non-cavitating flow an increase in \ / ——- is accompanied by an

2 sat

increase m

When the flow through the venturi is initially non-cavitating a reduction in P^ reduces the pressure in the venturi throat until cavitation

begins. In the case of a neat fluid cavitation occurs at or slightly

below the vapor pressure, and in a helium saturated fluid cavitation may

begin when the helium gas begins to desorb. A further reduction in P^ does

not correspondingly increase the flowrate, but rather increases the

extent of cavitation while the venturi throat pressure r remains near P

sat

or a slightly higher desorption pressure. In this cavitating condition

the flowrate is affected very little by changes in P^ and the flowrate is

36
0.13

0.12

I

0-11

C

0)

4-1 J 0.10

I

Q.

0.09

0.08

0.2

O Run #43 - Saturated at 124 bars (1800 psia)

Run #44 - Unsaturated

0.3

O

0.4 0.5 0.6

P 2 -P3 P 2

_ P sat

, dimensi onl ess

0.7

Figure 18. Cavitation parameter as a function of nozzle pressure drop parameter

for liquid nitrogen at 80. 3K ( 145° R)

.

37
0.13

0.12

QJ £ o.n

c

Q.

<J

0.10

>

0.09

0.08

0.2

O Run #42 - Saturated at 122 bars (1770 psia)

Run #41 - Unsaturated

0.3

qgtDD g Q

oc§R$>

on

0.4 0.5 0.6

P 2 "P

pr , dimensionl ess

sat

0.7

Figure 19. Cavitation parameter as a function of nozzle pressure drop parameter

n for liquid carbon monoxide at 85. 3K (153 R).

38
0.13

0.12

in

</) <DC

o

•r-C

0.11

0.10

<

0.09

0.08

0.3

O - Run =48 - Saturated at 76 bars (1100 psia

- Run =46 - Unsaturated

0.4 0.5 0.6 0.7

, dimensi onl ess

0.8 0.9

Figure ZO. Cavitation parameter as a function of nozzle pressure drop parameter

for liquid nitrous oxide at 255K (459

: R).

39
n. 13

0.12

1/1 £ o.n

o

c

E

0. 10

Dl

I cn

0.09

0.08

0.3

O- Run ^53 ,57,59 - Saturated at 76 bars (H00 psia)

- Run #52,56,58,60 - Unsaturated

0.4 0.5 0.6 0.7 0.8

P2 - P ? pr

, dimensionless

sat

0.9

Figure 21. Cavitation parameter as a function of nozzle pressure drop parameter

for liquid nitrous oxide at 266K (479°R).

40
0.13

0. 12

1/5 to - 0.11

toc

=

^0.10

0.09

0.08

0.3

O Run #54 - Saturated at 80 bars (1160 psia)

Run #55 - Unsaturated

ma

ooo

^

O QP O

0.4 0.5 0.6 0.7

'2 " ^3

'2 - rS at

•, dimensionless

0.9

Figure 22. Cavitation parameter as a function of nozzle pressure drop parameter

for liquid nitrous oxide at 244K (439°R).

41
approximately, W /P_-P ; therefore while the quantity \ / —— is

Z sat V P —

P

2 sat

/VP

3 relatively constant, the quantity y p _p maY vary widely.

2 sat

The onset of cavitation may then be identified in figures 18-22 as the

region in which \

Pa

_ p

becomes relatively flat. The effect of helium

* 2 sat

( K \

saturation in reducing slightly the flowrate l"\ /—— I at which cavitaP -P

2 sat

tion occurs may also be seen in figures 18-22.

3.5 Experimental Observations

Two phenomena observed during the testing of the CO and ^0

systems are noteworthy: (1) CO pressure oscillations in the transmitting

tubes leading to the pressure transducers, and (2) the formation of slush

when the ^0 system was vented rapidly. The oscillations were eliminated

without significantly affecting the pressure response time by inserting

wires into the instrument lines for frictional damping, and by adding a small expansion volume in the transmitting tubing. The latter was

located adjacent to the transducer and sized to eliminate oscillations.

The mixture of liquid-solid ^0 (slush N^O) formed when the system was vented rapidly because Boulder, Colorado atmospheric pressure is below

the ^0 triple point pressure. It is important to note this occurrence

since it pertains to the possible venting of ^0 systems in subatmospheric environments. Such formation of slush or solid may cause partial

or total bleed line blockage, flow disruptions, and inability to drain

the ^0 tank. Gradual venting and/or addition of heat during venting

prevents this slush formation.

4. Conclusions

Liquid and vapor phase compositions have been measured for the

helium-carbon monoxide system between 80 and 120 K and for the heliumnitrous oxide system between 235 and 285 K. Using enhancement factors

and Henry's Law constants as a guide, the data for each phase appear

internally consistent and have the proper qualitative behavior.

42
The effects of helium absorption and flow induced desorption were

investigated by means of scale model experiments. Calculations by a two phase flow correlation, using the phase equilibrium data of this study,

indicate that the flow resistance effect of desorbed helium in long

pipelines would be almost negligible until the pressure in the system

dropped below about 27 bars (400 psia) . Since this calculation was based on the worst case of assumed phase equilibrium the resistive

effects in the actual flow stream, in which the pressure dropped no

lower than 46 bars (675 psia), were expected to be even less. The

experimental data verified this calculation and reasoning and showed no appreciable two-phase resistive effects. Contrarily, phase equilibrium

calculations for the flow nozzle indicated that two phase choking would

occur at extremely low flow rates with helium saturated liquids. However, actual experimental flow rates were fifty times as great as the

calculated equilibrium choking flow rates. A definite effect of helium

desorption was detected in the nozzle flow tests; though smaller than

theoretical, this effect ranged from a four percent flow reduction for

N2 and CO to a twelve percent reduction in N^O. From these data we

may conclude that the assumption of phase equilibrium in a flow nozzle

or venturi leads to gross miscalculations.

43
5. References

1. Sinor, J. E. and Kurata, F. , J. Chem. Eng. Data 11, 537 (1966).

2. Kling, G. , Chem. Ingr. Tech. 38, 1080 (1966).

3. Duncan, A. G. and Hiza, M. J., Advances In Cryogenic Engineering,

Edited by K. D. Timmerhaus, 15, 42, (Plenum Press, New York, N.Y.

1970)

.

4. Wagner, W. , Cryogenics 13, 470 (1973).

5. Lewis, G. , Pratt Whitney Aircraft, West Palm Beach, Fla. , private

communication, (1973) .

6. Hust, J. G. and Stewart, R. B., Nat. Bur. Stand. (U.S.) Tech. Note

No. 202, (1963).

7. Buzyna, G., Macries, R. A. and Ellington, R. T. , Chem. Eng. Progr.

Symp. Ser. 59 (44), 101 (1963).

8. Hiza, M. J. and Duncan, A. G., Advances in Cryogenic Engineering

Edited by K. D. Timmerhaus, 14, 30, (Plenum Press, New York, N.Y.,

1969)

.

9. Martinelli, R. C. and Nelson, D. B., ASME Transactions 7^0: 695-716

(1948)

.

44
Appendix Al

The objective of runs 7 to 10 was to determine the effect (if any)

of desorbed helium gas on the flow resistance in a 6 m (22 ft) length

of 1.65 mm (0.065 in) ID tubing.

The venturi was first installed for runs beginning with no. 16.

The objective of runs 16 through 40 was to determine the downstream

pressure (varied by valve V23 setting) at the transition between non- cavitating and cavitating venturi flow. In runs 34 to 40 a 0.61 mm

(0.024 in) throttling orifice was installed upstream of V5 to provide

the desired pressure drop. A 91 cm (36 inch) length of 6.35 mm (1/4 in)

OD x 1.24 mm (0.049 in) wall tubing provided the required residence

time and replaced the 22 ft tube used in tests 7 to 33. Several of the

tests in Table Al (starting with run 15) were "burst tests" in which

valve V23 was varied during the test while all other test parameters

were held constant. The objective of these tests was to investigate

the effects (on flow) of varying pressure downstream of the venturi

for neat and helium-saturated fluids.

i

45
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cm

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to C

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O O 00 Oo o o o

CM O CM

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in in LTl in

—I ,-. O

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X X

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1 i i i oooooooooi i i i i £ i i i

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ca

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1 i i i i i

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1 i i i

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1

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< < < <

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< < < <

2 2 2 2

a c x Z

n. -w

Q, CD O °-

co £

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0* *

CJ PJ PJ PJ PJ

pj pj cj pj pj X X X X X

o o o c o

nO •£> vC <2 vO vfl ^ \C O

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Appendix A2

The throttling orifice upstream of V5 was removed for all runs

in Table A2, and the required pressure drop upstream from the venturi

was provided by throttling through valve V5. All of the runs in

Table A2 except 56-59 were "burst tests" in which either valve V5 or V23 was varied during the test while all other test parameters were

held constant. The objective of these tests was to investigate the

effects (on flow) of varying the pressure upstream and downstream from

the venturi for neat fluids and fluids saturated with helium.

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54
bS 14A (REV. 7-73)

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. NBSIR 75-806

2. Gov't Accession

No.

3. Recipient's Accession No.

4. TITLE AND SUBTITLE

PHA^F FOTTTT TRP.TTTM AND FT OU— TNTlTTrF'n nFQnPPTTntJ DATA

FOR He-CO, He-N

20, and He-N

2 SYSTEMS

5. Publication Date February 1975

6. Performing Organization Code

7. author(S)w. G. Steward, R.O. Voth, J. Hord, W.R. Parrish,

C. F. Sindt, J.M. Arvidson

8. Performing Organ. Report No.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No. 2750559

11. Contract/Grant No. AF/Kirtland No.

AFWL 75-177

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Kirtland Air Force Base

Albuquerque, New Mexico 87117

13. Type of Report & Period Covered

Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

16. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.) Liquid-vapor equilibrium data were obtained for the helium-carbon

monoxide and helium-nitrous oxide systems at pressures to 138 bars.

Liquid and vapor phase compositions were measured at nominal temperatures

of 80, 85, 90, 100, and 120 K (144, 153, 162, 180, and 216°R) for the

helium-carbon monoxide system, and at 235, 245, 265, and 285 K (423,

441, 477, and 513 °R) for the helium-nitrous oxide system. Internal con- sistency of the data was checked by using pseudo-Henry's law constants

and enhancement factors.

The effects on flow of helium absorption and subsequent flow-induced desorption were investigated by means of reduced scale model

experiments. Friction losses attributable to helium desorption in the

long channels proved to be negligible both in experimental measurements

and in calculations based on assumed equilibrium of liquid and gas.

Contrarily, the assumption of phase equilibrium leads to gross miscalculations of flow rates in cavitating or near cavitating nozzles or Venturis. Actual venturi mass flow rates reached in the experiments were fifty times the theoretical choking flow rates; however, definite

mass flow rate reductions due to helium desorption were measured,

ranging from four percent for nitrogen to twelve percent for nitrous

oxide. Pertinent experiences in handling these fluids and operating the

test pgnipmprif arp alan Hi gpusgpH . _

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons) Binary mixture; cavitating venturi; enhancement factors;

experimental vapor-liquid equilibria; flow induced desorption; heliumcarbon monoxide system; helium-nitrous oxide system; Henry's law

constants; two phase choking; two phase flow.

18. AVAILABILITY (j£ Unlimited 19. SECURITY CLASS

(THIS REPORT)

21. NO. OF PAGES

I For Official Distribution. Do Not Release to NTIS

UNCLASSIFIED

Order From Sup. of Doc, U.S. Government Printing Office Washington. D.C. 20402. SD Cat. No. C13

20. SECURITY CLASS

(THIS PAGE)

22. Price

^ Order From National Technical Information Service (NTIS)

Springfield, Virginia 22151 UNCLASSIFIED

ISCOMM • ERL USCOMM-DC 29042-P74