Neil T. Van Dresar

Glenn Research Center, Cleveland, Ohio

Liquid Oxygen Thermodynamic Vent System 

Testing With Helium Pressurization

NASA/TP—2014-216633

April 2014
	
	
	








	



	

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Neil T. Van Dresar

Glenn Research Center, Cleveland, Ohio

Liquid Oxygen Thermodynamic Vent System 

Testing With Helium Pressurization

NASA/TP—2014-216633

April 2014


	



	


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NASA/TP—2014-216633 iii 

Contents 

Summary........................................................................................................................................................................ 1

1.0 Background ......................................................................................................................................................... 1

1.1 Previous Thermodynamic Vent System Testing ...................................................................................... 3

1.2 Modeling and Theory ............................................................................................................................... 3

2.0 Comparison of Model Predictions With Previous Test Results .......................................................................... 5

3.0 Correlation of Model Predictions ........................................................................................................................ 6

4.0 Test Overview ..................................................................................................................................................... 7

5.0 Test Apparatus and Instrumentation ................................................................................................................... 9

6.0 Thermodynamic Vent System Control Logic and Interaction With Data System ............................................ 15

7.0 Signal Conditioning .......................................................................................................................................... 16

8.0 Boiloff Data Analysis ....................................................................................................................................... 16

9.0 Boiloff Test Results .......................................................................................................................................... 17

10.0 Thermodynamic Vent System Data Analysis ................................................................................................... 17

10.1 Thermodynamic Vent System Cycle Time ............................................................................................ 18

10.2 Thermodynamic Vent System (or Vent) Duty Cycle ............................................................................. 18

10.3 Pump Duty Cycle ................................................................................................................................... 18

10.4 Turnover Time ....................................................................................................................................... 18

10.5 Vented Mass and Vent Flowrate ............................................................................................................ 19

10.6 Heat Exchanger Performance ................................................................................................................. 19

10.7 Quality Downstream of Joule-Thomson Device .................................................................................... 20

10.8 Energy Balance Applied to a Complete Thermodynamic Vent System Cycle ...................................... 20

10.9 Thermodynamic Vent System Efficiency .............................................................................................. 20

10.10 Thermodynamic Vent System Test Results ........................................................................................... 20

11.0 Discussion of Results for Each Individual Thermodynamic Vent System Test................................................ 22

11.1 Test 90A ................................................................................................................................................. 22

11.2 Test 50A ................................................................................................................................................. 24

11.3 Test 90B ................................................................................................................................................. 25

11.4 Test 50B ................................................................................................................................................. 26

11.5 Test 90C ................................................................................................................................................. 26

11.6 Test 50C ................................................................................................................................................. 28

11.7 Test 90D ................................................................................................................................................. 29

11.8 Test 50D ................................................................................................................................................. 29

11.9 Test 50E ................................................................................................................................................. 30

12.0 Observations of All Thermodynamic Vent System Test Results ........................................................................ 31

13.0 Comparison of Test Results With Thermodynamic Vent System Model .............................................................. 34

14.0 Conclusions ....................................................................................................................................................... 35

Appendix—Symbols ................................................................................................................................................... 37

References ................................................................................................................................................................... 39

 

NASA/TP—2014-216633 1 

Liquid Oxygen Thermodynamic Vent System Testing 

With Helium Pressurization 

Neil T. Van Dresar 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Summary 

This report presents the results of several thermodynamic 

vent system (TVS) tests with liquid oxygen plus a test with 

liquid nitrogen. In all tests, the liquid was heated above its 

normal boiling point to 111 K for oxygen and 100 K for 

nitrogen. The elevated temperature was representative of tank 

conditions for a candidate lunar lander ascent stage. 

An initial test series was conducted with saturated oxygen 

liquid and vapor at 0.6 MPa. The initial series was followed by 

tests where the test tank was pressurized with gaseous helium 

to 1.4 to 1.6 MPa. For these tests, the helium mole fraction in 

the ullage was quite high, about 0.57 to 0.62. 

TVS behavior is different when helium is present than when 

helium is absent. The tank pressure becomes the sum of the 

vapor pressure and the partial pressure of helium. Therefore, 

tank pressure depends not only on temperature, as is the case 

for a pure liquid-vapor system, but also on helium density 

(i.e., the mass of helium divided by the ullage volume). Thus, 

properly controlling TVS operation is more challenging with 

helium pressurization than without helium pressurization. 

When helium was present, the liquid temperature would rise 

with each successive TVS cycle if tank pressure was kept 

within a constant control band. Alternatively, if the liquid 

temperature was maintained within a constant TVS control 

band, the tank pressure would drop with each TVS cycle. The 

final test series, which was conducted with liquid nitrogen 

pressurized with helium, demonstrated simultaneous pressure 

and temperature control during TVS operation. The 

simultaneous control was achieved by systematic injection of 

additional helium during each TVS cycle. Adding helium 

maintained the helium partial pressure as the liquid volume 

decreased because of TVS operation. 

The TVS demonstrations with liquid oxygen pressurized 

with helium were conducted with three different fluid-mixer 

configurations—a submerged axial jet mixer, a pair of spray 

hoops in the tank ullage, and combined use of the axial jet and 

spray hoops. A submerged liquid pump and compact heat 

exchanger located inside the test tank were used with all the 

mixer configurations. The initial series without helium and the 

final series with liquid nitrogen both used the axial jet mixer. 

The axial jet configuration successfully demonstrated the 

ability to control tank pressure; but in the normal-gravity 

environment, the temperature in the upper tank region (ullage 

and unwetted wall) was not controlled. The spray hoops and 

axial jet combination also successfully demonstrated pressure 

control as well as temperature control of the entire tank and 

contents. The spray-hoops-only configuration was not 

expected to be a reliable means of tank mixing because there 

was no direct means to produce liquid circulation. However, 

surprisingly good results also were obtained with the sprayhoops-only configuration (i.e., performance metrics such as 

cycle-averaged vent flowrate were similar to those obtained 

with the other configurations). 

A simple thermodynamic model was developed that 

correctly predicted the TVS behavior (temperature rise or 

pressure drop per TVS cycle) when helium was present in the 

ullage. The model predictions were correlated over a range of 

input parameters. The correlations show that temperature rise 

or pressure drop per cycle was proportional to both helium 

mole fraction and tank heat input. The response also depended 

on the tank fill fraction: the temperature rise or pressure drop 

(per TVS cycle) increased as the ullage volume decreased. 

1.0 Background 

The thermodynamic vent system (TVS) is a technology 

developed to control the tank pressure of subcritical cryogenic 

propellant tanks in low-gravity (low-Bond-number) 

environments where surface tension strongly affects liquid 

position. Because of their cold temperature and volatility, 

cryogenic liquids will inevitably absorb thermal energy from 

the surroundings and build up pressure in a closed vessel. In a 

normal- or reduced-gravity (such as on the Moon) 

environment, or if the liquid is settled by thrusting, 

conventional practice is to control pressure by venting vapor 

directly from the ullage as needed. For a given pressure 

(between the triple and critical points), a unique saturation 

temperature exists. Therefore, controlling tank pressure also 

maintains the liquid propellant in a saturated condition at the 

corresponding temperature. In a microgravity environment, it 

may not be possible to directly vent the ullage vapor with 

confidence because the liquid position within the propellant 

tank cannot be predicted with certainty. Furthermore, it may 

be undesirable to consume propellant to settle the liquid in the 

tank and enable direct ullage venting. The TVS is a proposed 

alternative method for controlling tank pressure in 

microgravity. 
NASA/TP—2014-216633 2 

With a TVS, a small quantity of sacrificial liquid is removed 

from the propellant tank: that is, it is vaporized and then vented 

from the tank. As depicted in Figure 1, liquid is withdrawn from 

the propellant tank and expanded through a Joule-Thomson 

(J–T) device to a lower pressure, becoming a two-phase mixture 

of liquid and vapor at a lower temperature. (In microgravity, a 

liquid acquisition device would be needed to extract liquid from 

the tank.) The two-phase mixture is then passed through the 

cold side of a heat exchanger, absorbing heat and vaporizing 

completely. The vapor is then vented to the external environment, generally space vacuum. Meanwhile, a second liquid flow 

is drawn from the tank and pumped through the warm side of 

the heat exchanger where thermal energy is removed from the 

flow. The cooled liquid is then returned to the tank and mixed 

with the tank’s contents. The mixing action is driven by various 

methods, such as axial jets or spray devices. 

A properly designed mixing system should at least produce 

uniform temperatures throughout the liquid region in the tank 

or perhaps the entire tank volume. Generally, a TVS is more 

efficient when operated intermittently because pump operation 

dissipates additional energy into the tank at a significantly 

greater rate than the environmental heating rate of the tank. In 

addition, when the TVS is operating, the pressure reduction 

rate is larger than the tank’s self-pressurization rate. Such a 

system with a liquid pump, known as an active TVS, is the 

subject of this report. An alternative design without a pump to 

circulate the liquid inside the tank is known as a passive TVS. 

Figure 2 illustrates typical tank pressure and temperature 

behavior during cyclic TVS operation. When the pressure is 

near the lower set point, the fluid is well mixed and isothermal. In the absence of helium pressurant gas, the minimum 

temperature corresponds to the saturation temperature at the 

lower set-point pressure. However, when the pressure is near 

the upper set point, the tank has been in a quiescent state for a 

long period and the fluid is most likely thermally stratified. 

Because the fluid temperature is not uniform, there is no 

longer a straightforward connection between the tank pressure 

and temperature. (In normal gravity, the liquid free surface is 

at the saturation temperature corresponding to the tank pressure, whereas the liquid below the free surface is subcooled 

and the vapor above is superheated.) 

A TVS also can be used to control temperature and pressure 

in a tank pressurized with gaseous helium (GHe). Two key 

differences that affect the behavior of a GHe-pressurized 

system in comparison to a tank containing only liquid propellant and its vapor should be recognized. First, with noncondensable GHe present, the total tank pressure is the sum of 

the vapor pressure of the propellant and the partial pressure of 

the GHe. Therefore, the total tank pressure and the propellant 

temperature are decoupled. The liquid is subcooled with 

respect to the total tank pressure. The second difference is that 

Figure 1.—Simplified schematic of thermodynamic 

vent system (TVS) hardware: 1—Joule-Thomson 

device, 2—Heat exchanger, 3—Liquid recirculation 

pump, 4—Axial jet (or alternatively, a spray bar). 

Figure 2.—Idealized thermodynamic vent system (TVS) 

operational control for a system of pure liquid 

propellant and its vapor. 

 
NASA/TP—2014-216633 3 

the mass of GHe in the ullage remains constant (assuming that 

no GHe is added to the tank and neglecting GHe dissolution in 

the liquid). As TVS cycles progress, liquid is removed from 

the tank and the ullage volume increases. Therefore, the GHe 

density decreases with each successive TVS cycle. The effects 

of having GHe in the ullage are discussed in greater detail in 

Section 1.1 and several other sections. 

For a tank that contains only liquid propellant and its vapor 

and that is subjected to pressure control by either ullage 

venting or TVS operation, liquid temperature can only be 

maintained at the saturation temperature corresponding to the 

tank pressure. If the tank is pressurized with GHe, ullage 

venting reduces pressure, but at the expense of some GHe 

loss. When the GHe is completely removed by venting, the 

liquid is at the saturation temperature corresponding to the 

tank pressure. A TVS allows pressure reduction and temperature control without the loss of GHe. 

1.1 Previous Thermodynamic Vent System 

Testing 

Most of NASA’s documented TVS experimentation has 

been conducted at the NASA Marshall Flight Center using 

liquid hydrogen (LH2), liquid nitrogen (LN2), and liquid 

methane (LCH4) (Refs. 1 to 5). Early testing was conducted 

primarily with pure propellant vapor in the tank ullage (Refs. 1 

and 2), but later tests placed a greater emphasis on conditions 

involving GHe as a pressurant gas (Refs. 3 to 5). The test 

hardware for the Marshall tests consisted of an 18.09-m3

cylindrical test tank with elliptical ends. A combined spray bar 

and concentric tube heat exchanger assembly was oriented 

vertically within the tank near the tank’s central axis. The 

liquid recirculation pump and J–T valve(s) were located 

outside of the test tank for easier accessibility. 

In the tests with pure propellant vapor in the ullage, the 

TVS concept was demonstrated to successfully maintain tank 

pressure between the lower and upper set points, typically 

using a control bandwidth of about 3 or 6 kPa. Nominal test 

pressures were from 118 to 138 kPa. The maximum operating 

pressure for the Marshall test tank was 344 kPa. Testing with 

noncondensable pressurant gas (helium) has been limited 

mostly to relatively low GHe mole fractions in the ullage 

(estimated to be 5 to 27 percent), generally because of 

maximum test pressure limitations. 

Of particular interest to the current work is the TVS testing 

done with LCH4 and GHe pressurization (Ref. 4). In all of the 

Marshall tests, the total test pressure was limited by the test 

tank’s operational limit and generally the use of GHe 

pressurization was limited to low mole fractions of GHe in the 

ullage. However, in the LCH4 test, the vapor pressure was 

lowered (by subcooling the liquid) to 56 to 70 kPa in 

comparison to the near-normal boiling point at atmospheric 

pressure (~100 kPa) for the other tests. The low vapor pressure 

allowed for a significantly greater GHe mole fraction in the 

ullage—estimated to be 60 to 65 mol%. The liquid saturation 

pressure (i.e., temperature) rose with successive TVS cycles 

while the tank pressure was kept within the control band. 

When the control mode was changed to maintaining liquid 

saturation pressure within a prescribed band, the tank pressure 

then dropped in a saw-tooth fashion with successive cycles. 

The only previous TVS test at the NASA Glenn Research 

Center was an LN2 test using an axial jet mixer and a multipleorifice J–T device in a 1.4-m3

 flightweight tank (Ref. 6). 

Results were obtained at three fill levels at a tank pressure of 

approximately 145 kPa, and no GHe pressurant was used. The 

axial jet mixer is a simpler alternative to the spray bar concept 

tested at Marshall. Extensive testing of the mixing capabilities 

of the axial jet has been conducted in normal gravity, in drop 

towers, and on orbit (Refs. 7 and 8). The multiple-orifice J–T 

device has been proposed as an alternative to a valve. The 

device has a number of larger diameter spin chambers that 

may reduce the possibility of clogging due to contaminants, 

and it has no moving parts. Multiple-orifice J–T devices have 

been tested in LH2, LN2, and LCH4 (Refs. 9 to 11). 

1.2 Modeling and Theory 

A simple spreadsheet tool was developed to characterize 

TVS behavior for a propellant pressurized with GHe given a 

few input parameters and assumptions. Fluid properties were 

obtained using software subroutines (Ref. 12) linked to the 

spreadsheet. The model predicts the thermodynamic state of 

the propellant tank at the times in the overall TVS cycle when 

the pressure and temperature are at local minimums. These 

times occur at the end of each successive TVS cycle at the 

moment when the pump is switched off and the TVS vent is 

closed. Assumptions at these points in time follow: 

(1) The entire tank contents are isothermal (well mixed). 

(2) The binary ullage mixture consisting of propellant vapor 

and gaseous helium has a uniform mole fraction 

throughout. 

(3) The partial pressure of the vapor is equal to its 

saturation pressure at the existing temperature. 

(4) The partial pressure of GHe equals the tank pressure 

minus the partial pressure of the vapor. 

Parameters input to the model were propellant type (i.e., O2, 

CH4, H2, or N2), tank volume, initial liquid fill fraction, energy 

input to the tank over a complete TVS cycle (environmental 

heat leak plus dissipated pump power), and the initial 

temperature and/or pressure. There are three modeling cases: 

Case 1—Constant Temperature: The model determines the 

final pressure. 
NASA/TP—2014-216633 4 

Case 2—Constant Pressure: The model determines the final 

temperature. 

Case 3—Constant Temperature and Pressure: The model 

determines the mass of added GHe. 

The model is based on three conservation laws: mass 

conservation of propellant (liquid and vapor), mass 

conservation of GHe, and an energy balance applied to the 

tank contents plus propellant vented by the TVS. For Case 2, 

the energy balance could include the change in tank wall 

energy (because temperature is not constant); however, the 

change in wall energy was found to be negligible for all the 

examples considered. For Case 3, the energy balance included 

the energy of the added helium. 

Thermodynamic properties of the fluid components were 

calculated as follows (all symbols used in this document are 

also defined in the Appendix): 

(1) The tank volume Vt and fill fraction F were used to 

determine the liquid and ullage volumes (Vl

 and Vu). 

Vl

 = FVt (1) 

Vu = (1 – F)Vt (2) 

(2) The temperature T was used to determine the vapor’s 

partial pressure Pv. 

Pv = Psat (id, T) (3) 

where subscript “sat” refers to saturated and “id” is the 

propellant identification. 

(3) The partial pressure of helium PHe was calculated. 

PHe = P – Pv (4) 

where P is the pressure of the tank. 

(4) The densities were obtained from a fluid properties 

program using the proper temperatures and pressures. 

l

 = l (id, P, T) (5) 

v = g (id, T) (6) 

He = He (PHe, T) (7) 

where the subscripts l, v, and g refer to “liquid,” 

“vapor,” and “saturated vapor.” 

(5) The masses m of each fluid component were calculated. 

ml = l Vl

 (8) 

mv = v Vu (9) 

mHe = He Vu (10) 

(6) The internal energy U of each fluid component was the 

product of the component’s mass and specific internal 

energy u obtained from the fluid properties program. 

Ul

 = ml ul (id, P, T) (11) 

Uv = mv ug (id, T) (12) 

UHe = mHe uHe (T, PHe) (13) 

(7) The total internal energy of the tank was the sum of the 

components: 

U = Ul

 + Uv + UHe (14) 

The conservation laws were applied to two consecutive 

TVS cycle minimums denoted as states 1 and 2. For the first 

calculation, state 1 was the initial condition. For the following 

cycles, the previous state 2 became the new state 1. 

The procedure for Case 2, pressure-controlled operation 

(P2 = P1) follows: 

(1) Determine all the properties given by Equations (1) to 

(14) for state 1. 

(2) Specify the energy Q added to the system between 

states 1 and 2. 

1 2   cycle Q Q Q t e p p     (15) 

where Qe

and Qp

are the environment heat leak rate 

into the tank and the rate of pump power dissipation 

into the tank fluid, p is the pump duty cycle (fraction 

of time that the pump is turned on), and tcycle is the 

elapsed time for a single TVS cycle.

(3) Guess the temperature at state 2, T2. 

(4) Determine the propellant and GHe partial pressures at 

state 2. 

(5) Determine the densities of liquid, vapor, and GHe at 

state 2. 

(6) Use the conservation of helium mass relation 

(mHe,1 = mHe,2) to find the fill fraction at state 2: 

Vt

m

F

He,2

He,2

2 1

 	 (16) 

(7) Use the fill fraction to determine liquid and ullage 

volumes at state 2. 

(8) Calculate the mass of the liquid, vapor, and GHe using 

the densities and volumes for each at state 2. 
NASA/TP—2014-216633 5 

(9) Use the conservation of mass relation for the propellant 

to determine how much propellant has been vented 

(denoted by subscript o) between states 1 and 2. 

,1 ,1   ,2 ,2

‡ mo  ml  mv 	 ml  mv (17) 

where ‡ indicates that the value was derived from mass 

balance. 

(10)Calculate the internal energy of the liquid, vapor, and 

GHe at state 2. 

(11)Determine the enthalpy of the vent flow leaving the 

tank, ho. For this work, the vent exit temperature was 

assumed to equal the temperature in the tank and the 

vapor was assumed to be saturated at the vent exit 

temperature. Furthermore, the vent flow enthalpy was 

assumed to be constant from state 1 to state 2. (For 

Case 2, T2 was used to obtain ho.) 

(12) Use the conservation of energy relation to determine how 

much propellant has been vented between states 1 and 2. 

o

o h

U U Q m
 1 	 2  12  (18) 

*	  		   	 	

 	"

balance. 

(13)Use the spreadsheet’s goal-seeking tool to find the 

value of T2 that makes the results of Equations (17) and 

(18) identical: 

0 ‡ 	  
 mo mo (19) 

(14)The spreadsheet now contains the correct values for 

vented mass between states 1 and 2, and the fill level 

and temperature for state 2. 

(15)Repeat for as many TVS cycles as desired by making 

the new state 1 equal to the old state 2. 

The procedure for Case 1, temperature-controlled TVS, is 

similar to that for Case 2 (just presented) except that the final 

temperature is specified (equal to the initial temperature) and 

the equations are solved for the final pressure. The procedure 

for Case 3, temperature and pressure-controlled TVS, is also 

similar, except that both the final temperature and pressure are 

equal to the initial values and that the unknown variable 

(requiring an initial guess) is the GHe mass added from state 1 

to state 2, mi. Therefore, in Step 6, mHe,2 = mHe,1 + mi. The 

temperature and pressure at the GHe inlet to the tank are 

needed to determine the enthalpy of the added GHe, hi. In this 

work, 300 K and the tank pressure were used. Equation (18) 

should be modified to include the enthalpy of the added GHe 

in the energy balance: 

o

i i o h

U U Q m h

m 	    
 1 2 12 (20) 

2.0 Comparison of Model Predictions 

With Previous Test Results 

Two reported TVS test results—the Marshall LCH4 test 

(Ref. 4) and an earlier test with LH2 (Ref. 5)—were compared 

with the model predictions. These tests had a number of 

similarities as well as significant differences. The key differences emphasized herein are the helium mole fraction and heat 

input: both were quite large for the CH4 test and very small in 

comparison for the hydrogen test. The fill fraction was about 

0.9 for both. 

Model predictions for two sets of TVS cycles from the CH4

test were generated. The required values input to the model 

were obtained from Reference 4 or estimated by hand 

digitizing the temperature and pressure plots from the figures 

in Reference 4. The first set of cycles was modeled as six 

pressure-controlled cycles (cycle minimums labeled 1 to 7 in 

Fig. 2 in Ref. 4). The second set of cycles was modeled as six 

temperature-controlled cycles (cycle minimums labeled 11 to 

17 in Fig. 3 in Ref. 4). The pump power was assumed to be 

50 W for the CH4 tests. (Because of the high reported heat 

leak (720 W) and the low pump-duty cycle, the predicted 

results were not highly sensitive to the pump power unless the 

pump power was much greater than the assumed value.) 

In Figure 3, the line is an approximate representation of the 

observed LCH4 temperature rise during a portion of the test 

that was under pressure-controlled TVS operation. The symbols indicate the predicted minimum temperatures at the end 

of each TVS cycle. The model underpredicted the experimental rate of temperature rise by about 40 percent. The GHe 

mole fraction was approximately 0.65, and the dimensionless 

heat input (discussed in Sec. 3.0) was about 1.8. 

Figure 3.—Comparison of model prediction with NASA 

Marshall Flight Center liquid methane test results for 

pressure-controlled thermodynamic vent system (TVS) 

operation (approximate test result data from Ref. 4). 
NASA/TP—2014-216633 6 

Figure 4.—Comparison of model prediction with Marshall 

liquid methane test results for temperature-controlled 

thermodynamic vent system (TVS) operation. 

Figure 4 shows an approximation of the tank pressure cycles 

(only the minimum and maximums were digitized) for a part 

of the LCH4 test with temperature-controlled TVS operation 

as well as the cycle minimum points predicted by the model. 

The model overpredicted the experimental rate of pressure 

drop by about 40 percent. For the temperature-controlled test 

segment shown, the GHe mole fraction was approximately 

0.55 and the dimensionless heat input was about 4.4. 

Considering the simplicity of the model and the uncertainty 

in some of the input values, the model did a reasonably good 

job of matching the experimental trends: the key point is that 

pressure-controlled operation caused an increase in liquid 

temperature, whereas temperature-controlled operation caused 

the tank pressure to drop. 

This TVS behavior (pressure drop or temperature rise) 

observed during the Marshall methane test had not been 

observed in previous TVS tests. A test from Reference 5 for 

LH2 pressurized with GHe at 90-percent fill was selected for 

comparison with the model. In the selected LH2 case, the 

helium mole fraction was 0.037. After a long transient period 

involving mixing without venting and heating the liquid, about 

20 consecutive TVS cycles were conducted using pressurecontrol logic. The tank pressure cycled within the lower and 

upper control points. The liquid temperature appears to be 

steady in Figure 6 in Reference 5. The model for this test 

predicted a temperature rise over the 20 cycles of 0.01 K—a 

very minor increase that would be difficult to detect. The 

model assumed a pump power of 25 W, which—combined 

with the reported values of tank volume, environmental 

heating rate, duty cycle, and cycle time—resulted in a 

relatively low dimensionless (*

) heat input value of Q* = 0.41. 

Therefore, the helium mole fraction and the heat input in this 

test appear to have been sufficiently low to result in the 

unnoticeable temperature rise during pressure-controlled 

operation. 

Another finding from modeling the Marshall LH2 TVS test 

was that there was substantial loss of GHe from the ullage 

because of solubility in the liquid. In the test, the amount of 

GHe added to the ullage was metered. The amount was 0.34 kg. 

However, on the basis of the test pressure, temperature data, and 

modeling assumptions, the GHe in the ullage should have been 

0.25 kg. If both values are correct, then 0.09 kg of GHe must 

have dissolved in the liquid. Given sufficient spray-bar operating time, the calculated amount of dissolved GHe is possible 

because of the vigorous liquid-ullage interaction resulting from 

liquid droplets passing through the ullage during spray-bar 

operation. The maximum amount of GHe that can be dissolved 

can be predicted using the correlation for GHe and hydrogen 

given in Reference 13. At the test pressure of 0.166 MPa and 

liquid temperature of 21.9 K, the correlation predicts a mole 

fraction of GHe dissolved into the liquid of 0.00013, which 

converts to 0.29±0.02 kg of GHe dissolved into the liquid. 

Therefore, 0.09 kg of GHe lost to solubility is feasible. Loss of 

GHe from the ullage due to solubility in the liquid is very likely 

in other TVSs as well, such as the Marshall CH4 test and the 

present work. 

3.0 Correlation of Model Predictions 

Although the model is quite simple, the manner in which the 

various input values affect the behavior of the TVS cycles is 

not readily apparent. Therefore, two correlations were developed to capture the major trends predicted by the model. First, 

the model was used to predict pressure loss per TVS cycle for 

a range of GHe mole fractions, tank heat values, and tank fill 

levels for liquid oxygen (LO2), LCH4, and LH2 when the TVS 

was operated in the temperature-control mode. The results 

were then correlated to show the significance of these factors. 

The following dimensionless groups were used:

 GHe mole fraction: P

P

x He He  (21) 

Dimensionless heat input: 1000 cycle *   Vt h fg

Qt

Q l



 (22) 

where Q Qe Qp   TVS , tcycle is the time for one cycle, and 

hfg is the latent heat of vaporization. 

 Liquid fill fraction: F

 Dimensionless pressure drop per TVS cycle (in percent): 

 100 *  	  int

int fin

P

P P

P (23) 

where int indicates “initial” and fin indicates “final.” 
NASA/TP—2014-216633 7 

 Dimensionless temperature: 

cp tp

tp

T T

T T

T 	

	  † (24) 

where the subscripts cp and tp indicate the critical point and 

triple point and † indicates the dimensionless quantity in 

correlation. 

The model was run for F from 0.03 to 0.95, using 10 values 

of F for T† = 0.2, 0.4, and 0.6 for each propellant. Values 

for xHe and Q*

 were randomly generated within the ranges 

0.01 < xHe < 0.7 and 0.1 < Q* < 6 for each pair of (F, T†

). The 

ranges of mole fraction, heat input, and temperature covered 

the experimental values for the current LO2 test as well as the 

Marshall LH2 and LCH4 tests. Values for xHe and F are for the 

start of a TVS cycle. The correlated predictions are shown in 

Figure 5 in the form of the following power law fit: 

F

x Q P 	  1

0.1 * * He (25) 

The correlation is quite good. Several trends are evident. 

First, the pressure drop per cycle is proportional to both the 

GHe mole fraction and the dimensionless heat input. Second, 

the pressure drop per cycle is larger at high fill levels (small 

ullage volume). Third, the correlation is independent of the 

dimensionless group T†

. Equation (25) is not reliable for 

values of T†

, xHe, and F above the upper limits stated earlier in 

this section. 

A second similar correlation of model predictions was 

developed for the pressure-controlled mode. The dimensionless group P*

 was replaced with T*, which was defined as the 

dimensionless temperature rise per TVS cycle (in percent): 

100 *  	 

int

fin int

T

T T

T (26)

The same ranges of F, T†

, xHe, and Q* were used. The 

resulting correlation for T*

 was 

 0.7

* * He

1

0.02

F

x Q T 	  (27) 

Note that Equation (27) is also independent of the 

dimensionless group T†

. The T* correlation, as shown in 

Figure 6, is not nearly as tight of a fit as the P*

 correlation, but 

otherwise it exhibits similar trends. That is, the temperature 

rise per cycle is proportional to the GHe mole fraction and the 

dimensionless heat input. Also, the dimensionless temperature 

rise per cycle increases with fill level, but it is a lesser function 

of fill level than is the dimensionless pressure drop per cycle. 

The key point of the correlation development was not so 

much to obtain the correlations, but rather to better understand 

Figure 5.—Correlated model predictions for dimensionless 

pressure drop per thermodynamic vent system (TVS) cycle; 

P*, dimensionless pressure, xHe, mass fraction of helium, 

Q*

, dimensionless heat transfer, and F, liquid fill fraction. 

Figure 6.—Correlated model predictions for dimensionless 

temperature rise per thermodynamic vent system (TVS) 

cycle; T*, dimensionless temperature, xHe, mass fraction of 

helium, Q*, dimensionless heat transfer, and F, liquid fill 

fraction. 

the TVS cycle behavior’s dependence on GHe mole fraction 

and heat input. The correlations show that pressure drop (in 

temperature-control mode) and temperature rise (in pressurecontrol mode) were both directly proportional to both GHe 

mole fraction and heat input. 

4.0 Test Overview

The test results presented herein are from the first TVS 

demonstration using LO2 conducted by NASA. For most tests, 

the liquid was heated about 21 K above its normal boiling 

point and pressurized with helium to about 1.6 MPa. These 

conditions were chosen to represent (as closely as possible 

given test facility limitations) operating conditions for a lunar 

lander ascent-stage oxidizer tank. The resulting GHe mole 
NASA/TP—2014-216633 8 

fraction (~60 mol%) was comparable to that of the Marshall

TVS test with subcooled LCH4. A benefit of having an 

elevated liquid temperature relative to the normal boiling point 

was that the TVS could be vented to the atmosphere with the 

J–T expansion providing a sufficient temperature drop in the 

vent flow for successful TVS heat exchanger performance. 

Three different fluid mixing configurations were tested: 

(1) a submerged axial jet with vertical upflow, (2) liquid 

droplets sprayed directly into the ullage using a pair of spray 

hoops near the top of the tank, and (3) combined flow using the 

axial jet and spray hoops simultaneously. Figures 7 and 8 show 

simple diagrams of the axial jet and spray hoop concepts, 

respectively. Optimization of the pump operation duration prior 

to venting was not possible because of time constraints. The 

time delay between pump startup and vent valve opening 

Figure 7.—Axial jet. 

did not appear to be beneficial in the tests with GHe 

pressurization, so the delay time was reduced for these tests. 

Two special tests also were conducted: one without GHe 

pressurization and another with LN2 as the test liquid. The test 

without GHe was primarily an initial checkout to verify TVS 

operation and to test automated control using two modes of 

control logic. For this test only, the pump was briefly operated 

prior to venting in each TVS cycle to destratify the liquid. The 

LN2 test allowed further testing with a second test fluid and 

was also used to demonstrate more advanced control logic 

using systematic helium repressurization. All of the special 

tests were performed with the axial jet configuration. Boiloff 

testing was done to quantify the environmental heating rate of 

the test. Table I summarizes the conditions for each test series 

performed: Test Series A through E. 

Figure 8.—Spray hoops.

TABLE I.—SUMMARY OF COMPLETED TESTS AND CONDITIONSa

Test 

series 

Hardware 

configuration 

Test 

fluidb

Approximate test pressure, 

MPa 

Boiloff test at 

~95-percent fill 

TVSc test at 

~90-percent fill 

TVSc test at 

~50-percent fill 

Boiloff test at 

40-percent fill 

A Axial jet O2 only 0.6 Yes Yes Yes No 

B Axial jet O2/GHe 1.5 to 1.6 No Yes 

C Axial jet and spray hoops O2/GHe 1.3 to 1.6 No No 

D Spray hoops O2/GHe 1.4 to 1.6 Yes No 

E Axial jet N2/GHe 1.6 No No No a

Pressures and fill levels are approximate. See test result sections for detailed information. 

b

O2, oxygen; GHe, gaseous helium; N2, nitrogen. c

TVS, thermodynamic vent system. 
NASA/TP—2014-216633 9 

5.0 Test Apparatus and 

Instrumentation 

The test tank was cylindrical with domed ends plus a 

smaller upper cylindrical section with a domed lid. It had a 

main diameter of 1.22 m and an overall height of 1.91 m. The 

upper section had a diameter of 0.48 m and a height of 0.34 m. 

The tank was made of stainless steel and had a wall thickness 

of 0.95 cm. Total tank volume was approximately 1.64 m3

. 

The tank and tank lid were bolted together at matching flanges 

at a location corresponding to a fill level of approximately 

98 percent. The lid had seven ports for plumbing and electrical 

feedthroughs. Four additional ports were located at the tank 

bottom. Figure 9 shows the bare test tank prior to installation 

of the multilayer insulation (MLI). The tank is shown 

suspended from the vacuum chamber lid, which is supported 

by the blue framework. 

Eighteen thermocouples were attached to the exterior 

surface of the test tank at various locations to monitor the tank 

wall and lid temperatures. A 432-W band heater was clamped 

around the tank at the 27-percent fill level. The heater was 

used only to heat the liquid in the tank to the desired test 

temperature. The tank was insulated with three layers of MLI. 

Figure 9.—Test tank prior to installation 

of multilayer insulation. 

The inner and outer layers were made of an aluminized tearresistant material, and the center layer was double-aluminized 

polyethylene terephthalate (PET) film. The layers were 

separated by polyester netting. Figure 10 is a photograph of 

the test tank after installation of the MLI. The purpose of the 

MLI was not so much to reduce heat leak: it was necessary to 

allow the test tank to absorb heat quickly so that a suitable 

number of TVS cycles could be performed in a relatively short 

time. Instead, a minimal number of MLI layers were used to 

provide at least some resemblance to flight hardware. 

The test tank was suspended from three load cells used to 

weigh the tank to measure liquid fill levels accurately. Each 

load cell had a full-scale range of 8900 N and a static error 

band (accuracy) of ±0.03-percent of full scale. The load cells 

were designed as calibration reference standards. A calibration 

system was used to calibrate the tank-weighing system, which 

is shown with its calibration hardware in Figure 11. The 

calibration system consisted of a rigid beam that could rotate 

about its midpoint position in a bearing assembly anchored to 

the floor underneath the vacuum chamber. The end of the 

beam directly below the center of the vacuum chamber could 

exert a downward force on the test tank via a linkage that 

passed through a port with a flexible bellows in the bottom of 

Figure 10.—Test tank with multilayer insulation 

installed. 
NASA/TP—2014-216633 10 

Figure 11.—Weighing system and calibration hardware. 

the vacuum chamber. The linkage included a calibration load 

cell. The bellows design eliminated any sensitivity of the 

calibration load on the external-to-internal pressure difference, 

and the linkage was decoupled during actual testing. The 

opposite end of the rigid beam was loaded with an upward 

applied force from a hand-activated pneumatic jack. Therefore, 

the applied load acted on the weighing system and could be 

directly measured with the calibration load cell. 

The tank vent and fill lines were constructed with multiple 

flex hose sections to minimize additional load forces acting on 

the test tank. Other physical connections between the vacuum 

chamber lid and the test tank, such as the fluid-sampling tubes, 

tank pressurization line, and electrical leads, had greater 

compliance and minor impact on the weighing of the tank. The 

fill and vent lines had a small impact on weighing system 

performance because the flex lines stiffened and moved when 

pressurized. This effect was accounted for in the calibration of 

the weighing system. The calibration equation for the fluid 

mass (liquid plus vapor) follows: 






     

3

j 1

m fl Rlc P

j (28) 

where the first term on the right is the sum of the three 

individual load cell readings . j Rlc The second term is a small 

correction for the effects of tank pressure and the third term is 

approximately equal to the tare weight of the test tank. For the 

series of tests reported herein, the values for the coefficients 

were 

 = 0.10222 kg/N  = –4.1076 kg/MPa  = –773.66 kg (29) 

The calibration fitted 330 data points of simulated fluid 

weight to within 0.25 percent of the tank full-scale weight 

(LO2 at 111 K and 1.59 MPa), and the calibration error band

accounts for the repeatability, hysteresis, and nonlinearity of 

the system. The weighing system was used primarily to 

determine liquid fill level, but because it could sense small 

changes in level, it was useful as a secondary method of 

measuring outflow rates during boiloff or TVS testing. 

A submersible, centrifugal liquid oxygen pump that was 

designed for 37.9 liters/min at 2920 rpm with a 6.1-m head 

was located inside and near the bottom of the test tank. The 

direct-current (DC) pump motor power was 100 W at the 

maximum speed of 3500 rpm (120 percent of the design 

speed). The pump was operated at the design speed for all 

tests. It could also be operated in LN2, but at an off-design 

head condition. 

The TVS heat exchanger was attached at the pump inlet as 

shown in Figure 12. The heat exchanger is a coiled-loop design 

where the TVS vent flow flowed through 22 loops of tubing 

located between cylindrical inner and outer shells. On the hot 

side of the heat exchanger, liquid flowed through the annulus 

over the external surface of the coiled tube and then into the 

pump inlet. Further details about the modeling and design of the 

heat exchanger are available elsewhere (Ref. 14). The discharge 

from the pump was flowed through piping to hand valves 

mounted outside the tank where the flow could be routed to 

either a submerged axial jet or a pair of spray hoops near the top 

of the tank via separate lines to each that reentered the tank. 

Flow for the cold side of the heat exchanger was drawn from a 

well formed by one of the bottom tank ports and was routed 

through a Visco Jet (Lee Company, Westbrook, CT) (Ref. 15) 

with a flow resistance rating of 5000 Lohm.1

 The Visco Jet was 

a multiple-orifice flow restrictor used as a J–T expansion 

device. After expanding to a lower pressure and temperature, 

the cold-side vent flow was vaporized as it flowed through the 

coiled loops of the heat exchanger and then routed out of the 

tank bottom and through the vacuum chamber lid to an 

atmospheric vent system containing vent gas flowmeters 

(discussed later in this section). Figure 13 shows the internal 

hardware and instrumentation suspended from the test tank lid, 

and Figure 14 shows the internal hardware and instrumentation 

installed in the bottom of the test tank. Figure 15 is a 

photograph of the TVS plumbing on the outside of the test tank. 

In a previous TVS experiment at Glenn (Ref. 6), it was 

observed that the submerged portion of the TVS vent line 

would fill with liquid under no-flow conditions. Each time 

venting was initiated, a liquid slug would pass through the 

vent line without effective vaporization. The corresponding 

loss of cooling resulted in reduced TVS efficiency. For the 

present series of tests, a cryogenic coaxial solenoid (shutoff) 

 

1

A measure of fluid resistance (http://www.theleeco.com/ 

engineering/lohm-definition.cfm Accessed Feb. 2014). 
NASA/TP—2014-216633 11 

Figure 12.—Thermodynamic vent system (TVS) heat exchanger and pump assembly. 

valve was installed in the inlet line upstream of the J–T device. 

This valve could not be placed inside the tank, however, so the 

inlet flow was routed out of a bottom port, through the solenoid 

valve, and back into the tank to the J–T device. The solenoid 

valve generated heat when energized, and because it was operated in a vacuum environment, an aluminum thermal strap was 

fabricated to transfer heat to the vent line outside the test tank. 

As an additional precaution, the thermal strap was cooled with 

LN2 from an external source. 

The axial jet was constructed of stainless steel tubing with 

an inner diameter of 1.65 cm. The discharge orifice in the 

axial jet’s end cap had a diameter of 1.02 cm and was located 

at a height corresponding to the 34-percent fill level. The axial 

jet was offset from the tank’s vertical axis by approximately 

0.2 m as shown in Figure 14. The upper and lower spray 

hoops were both constructed of 1.27-cm-diameter stainless 

steel tubing. Both hoop diameters were 0.43 m, and the hoops 

were centered about the tank’s vertical axis along with 

the tank fill/drain tube. The upper hoop had 36 discharge 

holes directed upward or upward/inward. The lower hoop 

had 102 discharge holes angled either upward/outward, 

upward/inward, or slightly downward/inward. All spray hoop 

discharge holes had diameters of 1.5 mm. The vertical supply 

pipe from the pump to the hoops had a 1.9-cm diameter. The 

upper spray hoop was located at a height corresponding to the 

99-percent fill level, and the lower hoop was located at a 

height corresponding to the 94-percent fill level. 

Instrumentation inside the test tank included two vertical 

temperature sensor rakes mounted near the tank’s central axis. 

Rake A contained 11 silicon-diode temperature sensors positioned at 5-, 10-, 20-, 30-, 40-, 50-, 60-, 70-, 80-, 90-, and 

95-percent fill level locations. Rake B contained 10 silicondiode sensors at the 15-, 25-, 35-, 45-, 65-, 75-, 85-, 96-, 97-, 

and 99-percent fill level locations plus two close-spaced diode 

clusters with sensors located at 1-percent intervals from 46- to 

55-percent fill and 86- to 95-percent fill. One purpose of the 

diode clusters was to measure temperature gradients near the 

liquid surface at the two nominal fill levels of 50 and 

90 percent. The second function of the diode clusters was to 

sense liquid level by operating the diodes in a high current 

mode to determine which sensors were submerged and which 

ones were exposed to the ullage. Only the diodes in the two 

clusters were set up to sense liquid level. Other silicon diode 

temperature sensors were located inside the piping to measure 

the temperature at the entrance to the upper and lower spray 

hoops, near the axial jet, at the pump discharge, upstream and 

downstream of the J–T device, and in the flow downstream of 

the heat exchanger on the vent side (two sensors). 
NASA/TP—2014-216633 12 

Figure 13.—Internal hardware and instrumentation 

suspended from the test tank lid. (The upper 

spray hoop is hidden from view.) 

The remaining piece of instrumentation inside the tank was 

a venturi meter located downstream of the pump and upstream 

of the flow control valves to the axial jet and spray hoops. The 

only significant instrumentation failure during the test was a 

cryogenic differential-pressure sensor attached to the venturi 

meter. Two sensing lines were routed out of the test tank 

bottom with short tubing extensions to connect to the 

differential-pressure sensor. These lines and the transducer 

body, but not the electronics module, were insulated with MLI 

insulation. The output from the differential-pressure sensor 

was not consistent, and the zero-flow reading had substantial 

shifts from day to day. The venturi meter results were deemed 

to be unreliable and are not reported. 

Outside of the test tank, there were three internal flowsensing thermocouples located along both the fill and vent 

lines to measure fluid temperatures. Each of these lines also 

had two additional thermocouples attached to the external 

walls. The TVS vent line had a wall-attached thermocouple on 

each side of the thermal strap attachment point plus another 

near the vacuum chamber lid. Pressure inside the test tank was 

sensed by a pressure sensor attached to the ullage vent line. 

Other external data recorded included the test facility indoor 

temperature, outdoor temperature, and barometric pressure. 

Two digital thermal-type mass flowmeters were installed in 

the test tank’s vent line. A 500-slpm meter was used2

 to 

measure flowrates during TVS venting. A 20-slpm meter was 

intended to measure steady-state boiloff rates. The manufacturer’s reported accuracy specifications were ±1.0 percent of the 

full-scale rating (FS) above 20 percent of FS and ±0.2 percent 

of FS below 20 percent of FS. Furthermore, the meters had a 

50:1 turndown ratio (1/50 is 2 percent of FS), below which the 

readings were unreliable. Both meters were calibrated by 

Glenn’s Calibration Laboratory at the conclusion of the test 

program. The large meter was calibrated with air, and the small 

meter with nitrogen gas. The calibration results were fitted with 

a linear equation having a zero intercept. The calibration factors 

were 1.019 for the large meter and 1.0026 for the small meter. 

The manufacturer’s conversion factors were 0.998 for air, 

1.000 for nitrogen, and 0.988 for oxygen. Therefore, the 

conversions used for oxygen flowrates were 0.988/0.998 = 

0.990 for the large meter and 0.988 for the small meter. 

Similarly for nitrogen TVS flowrates, the conversion for the 

large flowmeter was 1.000/0.998 = 1.002. (The small meter was 

not used with nitrogen because there was no nitrogen boiloff 

test.) For the conversion of volumetric flowrate to mass 

flowrate, the appropriate densities for oxygen or nitrogen had to 

be used. Density was based on the referenced standard pressure 

and temperature at calibration. Both meters were calibrated 

using a standard pressure of 101.3 kPa (1 atm), but the reference 

temperatures were different: 21.1 °C for the large meter and 

0 oC for the small meter. The corrected measured flowrate was 

calculated as follows: 

mo  ref Rfm Ccal Cfluid (30) 

where C is the flowmeter coefficient and the subscripts ref, fm, 

and cal indicate reference, flowmeter, and flowmeter 

calibration. 

 

2

Standard liters per minute. 
NASA/TP—2014-216633 13 

Figure 14.—Internal hardware and instrumentation in the bottom of the test tank; TVS, thermodynamic 

vent system. 
NASA/TP—2014-216633 14 

Figure 15.—Thermodynamic vent system (TVS) plumbing outside the test tank. 
NASA/TP—2014-216633 15 

6.0 Thermodynamic Vent System 

Control Logic and Interaction With 

Data System 

The control logic developed for the present series of tests 

allowed automatic and unattended operation of the initial 

heating and self-pressurization portion of the test as well as 

TVS operation. Human intervention was only required to start 

the test, switch TVS control between pressure or temperature 

control modes, adjust the tank pressure (by adding GHe 

pressurant), and change the tank fill level. Various input 

parameters could be changed while a test was in progress, but 

generally, these parameters were not modified during a 

particular test other than to switch between pressure and 

temperature control. The number of desired TVS cycles was 

generally set to a large number so that the tests would not stop 

during unattended operation. 

Figure 16 shows a flow chart showing the control logic. 

User input parameters follow: 

(1) Lower limit for the tank pressure, Pmin

(2) Upper limit for the tank pressure, Pmax 

(3) Lower limit for the average liquid temperature, Tmin

(4) Upper limit for the average liquid temperature, Tmax 

(5) Time delay tel0 for checking if the maximum pressure or 

temperature has been reached 

(6) Time delay tel1 for checking the rate of tank pressure decay 

while the pump is operating with the vent valve closed 

(7) Time delay tel2 for checking if the minimum pressure or 

temperature has been reached 

(8) The value of tank pressure decay rate (DPDT) at which a 

decision is made whether or not to open the vent valve 

(9) Choice of tank pressure control or average liquid 

temperature control (CntrlFlg) 

Figure 16.—Flow chart of control logic for heating/self-pressurization and thermodynamic vent system (TVS) 

operation; P, pressure; tel, elapsed time for type of control logic; CntrlFlg, flag for user-specified temperatureor pressure-control logic; and T, temperature; DPDT, user-specified tank pressure decay rate in control logic. 
NASA/TP—2014-216633 16 

The control logic was programmed into the test facility 

control software and required information exchange with the 

data system. The data system provided real-time values of the 

tank pressure and the averaged liquid temperature. All of the 

silicon-diode temperature sensors inside the test tank were 

available for use in determining the (arithmetic) average 

temperature. A user interface in the data system allowed convenient selection of the desired sensors. The first attempt to 

define the average temperature using four or five distributed 

sensors (all submerged in the liquid) worked well, and this 

method was kept throughout the test program. For tests at 

90-percent fill, the sensors at the 10-, 30-, 60-, and 80-percent 

fill locations were used; and for tests at 50-percent fill, the 

sensors at the 10-, 20-, 30-, and 40-percent fill locations were 

used. Two recording rates were used by the data system: a 

slower rate when the pump was off (one recording per 90 to 

240 s, typically 180) and a faster rate when the pump was 

operating (one recording per 60 to 180 s, typically 90). The 

control software would turn the pump off and on and share the 

pump status with the data system. Noise in the data could 

cause errors in the decisionmaking steps. Therefore, time 

delays were inserted in the logic. The data system scanned all 

data channels approximately twice a second. 

7.0 Signal Conditioning 

Early in the tests, additional steps were required to improve 

control performance, leading to signal-conditioning improvements for both rate of pressure change and average liquid 

temperature. Calculating the rate of change of pressure using 

consecutive readings could produce a noisy and unstable 

result, especially when one was looking for dP/dt to approach 

a constant value (such as 0). A smoothing function was used: 

  new old d

d 1

d

d

d

d

t

P

t

P

t

P    	  (31) 

where  was typically 0.1 to 0.2. The function worked well to 

smooth the dP/dt signal. The computed value of dP/dt lags the 

true value, but the lag was not a significant issue for TVS 

control because of the much longer time scale of the TVS 

cycles. 

The average liquid temperature, although typically based on 

four sensor readings, was initially rather noisy as in Section 

11.0. The noise caused control problems during temperaturecontrol mode when the minimum-to-maximum temperature 

difference was narrowed (typically 0.17 K). The signal noise 

was reduced by using a running average of 60 consecutive 

scans (about 30 s) of data. This averaging interval is short in 

comparison to the typical TVS cycle time of several hours. 

8.0 Boiloff Data Analysis 

The environmental heat input rate Qe

was determined from 

the measured boiloff rate as follows: 

  t

V P Q m h m h h t o ex g

f g

f

e o fg d

d



 	  















	 

 (32) 

where  is the energy derivative defined in Equation (33) and 

the subscripts e and ex indicate “environmental” and “exit,” 

respectively. 

The first term on the right-hand side assumes that the tank 

contains saturated liquid and vapor and that all heat input 

results in phase change (evaporation). The density ratio 

accounts for the portion of vapor mass that remains in the tank 

and occupying the space of the vaporized liquid. The second 

term on the right-hand side is the heating rate required to heat 

the vapor exiting the tank from the saturation temperature to 

the temperature of the superheated vapor at the tank exit (the 

inlet to the vent line). The tank vent was located at the top of 

the tank, so the uppermost temperature measurement inside 

the tank at the 99-percent fill level was used to obtain he. The 

second term is small at very high liquid fill levels, but it 

becomes significant as the fill level decreases. The last term 

on the right-hand side is ideally zero when the tank pressure is 

kept at a constant value during the boiloff test. In the tests 

reported herein, the pressure drifted slightly and the corresponding change in fluid temperature (resulting in propellant 

mass gain or loss) should be accounted for. The last term was 

obtained from a model assuming saturated liquid and vapor 

(Ref. 16). The energy derivative  is defined as 







 











  P

1 u (33) 

is calculated numerically (using Ref. 12 to obtain fluid 

properties) and is dependent on the fill fraction of the test tank 

as discussed in Reference 16. 

The boiloff rate using the weighing system was calculated as 

fin int

fl int fl fin o t t

m m m 		  , , (34) 

The weighing system calculation was applied to time spans 

where the fluid mass loss was observed to be linear with time. 

The accuracy of the boiloff rate calculated from the weighing 

system depended on the time span used. The estimated 

accuracy was 4 percent for 10 hr of elapsed time. 
NASA/TP—2014-216633 17 

9.0 Boiloff Test Results 

Three boiloff tests were conducted to quantify the total heat 

leak rate into the test tank. True steady-state thermal 

conditions (i.e., steady temperatures, pressures, and boiloff 

rates) did not consistently occur throughout the TVS tests, and 

the actual environmental heat input rate was somewhat variable because of differences in the upper tank wall temperatures 

(e.g., warm unwetted upper wall with axial jet vs. cold-wetted 

upper wall with spray hoops), small differences in the 

uncontrolled vacuum chamber wall temperature (no cryoshroud was used) and vacuum level, and variability in the 

initial thermal conditions prior to the TVS tests. The test 

schedule also did not allow sufficient time to establish true 

thermal equilibrium prior to and during each test. As a result, 

the actual heat input rate during an individual TVS test was 

not necessarily the same as for other tests or for the boiloff 

tests. Nonetheless, the boiloff results provide the best available estimate of heat leak rates for the TVS tests. 

Table II summarizes the boiloff results. All boiloff tests 

were conducted with LO2 and ranged from 34 to 47 hr. Two 

tests were performed at 93- to 94-percent fill, and one at 

39-percent fill. The tank pressure was regulated by automated 

back-pressure control of the tank’s vent system at approximately 0.60 MPa (saturation temperature of 111.5 K). Two 

independent means of measuring the boiloff rate were used: 

(1) the boiloff vent flow was routed through gas flowmeters, 

and (2) the change in fluid mass contained in the test tank was 

measured by the test tank weighing system. 

Unfortunately, the boiloff rate was nearly equal to the fullscale range of the 20-slpm vent line flowmeter, whereas the 

reading based on the much larger 500-slpm flowmeter was 

below the lower accuracy limit of the meter’s turndown ratio. 

The smaller meter was over-ranged in Test A. In general, the 

flowmeter output did not achieve clearly observable steadystate conditions, but it did appear to be approaching 

reasonably steady behavior. The flowmeter readings and 

corresponding tank pressure data used for the boiloff analysis 

are shown in Figure 17. 

The two heat leak measurements from the flowmeter (Tests 

B and D) and the three measurements from the weighing 

system (Tests A, B, and D) were averaged to obtain a 

representative total heat leak rate. The heat leak rate was 

slightly higher for the lower fill-level test, but this trend with 

fill level was not considered to be meaningful. In the 

remainder of this report, a heat input rate of 97 W is used. The 

uncertainty in the heat input rate result was attributed mainly 

to the insufficient time to reach truly steady conditions and 

was conservatively estimated to be no greater than ±10 W 

(10 percent). 

In the two high-fill-level tests, the ullage heating component 

was about 8 to 9 percent of the total. At the lower fill level, the 

ullage heating component was 22 percent of the total. In all 

TABLE II.—SUMMARY OF BOILOFF TEST RESULTS 

Test parameter Test series 

A B D 

Fill level at end of test, percent 94 39 93 

Pressure, MPa 0.61 0.58 0.61 

Total time from start of back pressure 

 control, hr 

34.5 47.0 34.0 

Elapsed time using flowmeter, hr ------ 5.1 3.0 

Elapsed time using weighing system, hr 11.3 8.0 8.0 

Vent flowrate from flowmeter, kg/hr ------ 1.55 1.61 

Vent flowrate from weighing system, kg/hr 1.57 1.42 1.52 

Total heating rate using flowmeter, W ------ 106 95 

Total heating rate using weighing system, W 92 98 92 

Phase change component, percent 91 78 89 

Ullage heating component, percent 8 22 9 

Energy storage/release component, percent 2 0 2 

Figure 17.—Vent flowrate and tank pressure data used for 

boiloff analysis of Test Series B and D. 

tests, the effect of the mass gain or loss in the test tank (due to 

unsteady tank pressure) was very small, from 0 to 2 percent of 

the total. Vacuum chamber pressure levels were similar in 

each of the tests, ranging from 3.0×10–4 to 6.7×10–4 Pa and 

averaged vacuum chamber wall temperatures (24 distributed 

locations) also were similar, ranging from 292 to 294 K. An 

analysis was performed on the vent line to estimate the 

contained vapor mass gain or loss due to temperature change. 

The vent line temperature effect was found to be negligible. 

10.0 Thermodynamic Vent System Data 

Analysis 

In this section, the methodology used to analyze the TVS 

data is explained. 
NASA/TP—2014-216633 18 

10.1 Thermodynamic Vent System Cycle Time 

A TVS cycle was defined to start when the vent valve was 

switched from an open to closed position. To ensure the start 

condition, the previous time step in the data file had to show 

that the vent valve was open. During a cycle, the following 

progression occurred: 

(1) The vent valve closed at or near the cycle minimum 

(pressure or temperature). 

(2) The vent valve remained closed as the liquid was heated 

and the tank pressure rose. 

(3) The peak (temperature or pressure) condition was 

reached, and the vent valve was opened. 

(4) The vent valve remained open as the liquid was cooled 

by operation of the TVS and the pressure dropped. 

(5) The cycle was completed when the cycle minimum was 

reached. The vent valve closed, which coincided with 

the start of the next cycle. 

In the data analysis, the open or closed status of the vent 

was determined from a reading of the TVS vent flowmeter. 

The vent valve was considered to be open whenever the vent 

flowmeter reading exceeded a threshold value. (The value was 

set at 1 percent of the meter’s full-scale range to prevent 

erroneous open or closed classifications due to signal noise or 

to the slight offset observed during zero-flow conditions.) 

Although the vent valve required a finite time to open and 

close and the flowmeter also had a brief startup and shutdown 

response time, flowrate values between the no-flow and 

steady-flow values were very rarely recorded. The TVS cycle 

time was obtained by simply summing the time steps from the 

start to the end of a cycle. The individual cycle times within 

a group of similar successive cycles were arithmetically 

averaged to obtain the average cycle time, cycle t . The 

uncertainty in average cycle time is solely due to the 

uncertainty of the vent valve closing time, discussed shortly, 

and averaging over multiple cycles reduces this uncertainty 

per cycle. All uncertainties provided herein were estimated 

using the methods from Reference 17. 

10.2 Thermodynamic Vent System (or Vent) 

Duty Cycle 

The duty cycle of the TVS TVS is the fraction of time that 

the TVS vent valve is open. This time was obtained by 

numerically integrating the TVS vent status over a complete 

cycle. The exact moment of valve opening was not recorded, 

so the trapezoidal rule was applied to the time step where the 

TVS vent status went from closed to open. The vent could 

open during the same time step as the pump startup or could 

be delayed by 1 to 2 time steps. Reported duty cycle values 

were arithmetically averaged over the available number of 

similar cycles. 

The exact moment of vent closing was also not recorded, so 

the trapezoidal rule was applied here also (to the time step 

where the TVS vent status went from open to closed). 

Application of the rule was complicated by the simultaneous 

shutdown of the pump and the changing of the data-recording 

rate. If the recording rate was slowed down when the pump 

was turned off, then the shutdown had to occur within the first 

portion of the time step (equal to the shorter time step), and 

the shorter time step was used with the trapezoidal rule. The 

slowed-down recording rate was used in most of the tests. 

However, in some tests, the recording rate was increased when 

the pump was shut down and the vent closing could have 

occurred at any point within the time step. In either case, the 

correct time step to use was the shorter one. 

Use of the trapezoidal rule can be viewed as numerically 

equivalent to assuming that the open/close events occur as step 

function changes at the midpoint of the time step; the timing 

uncertainties were thus equal to 1⁄2 of the time step. The 

uncertainty in the total venting time of a TVS cycle was 

dominated by the uncertainties in the opening and closing times. 

10.3 Pump Duty Cycle 

The pump duty cycle p is defined as the fraction of time 

that the pump was on (in operation). The elapsed time that the 

pump was on during a cycle was determined by numerically 

integrating the pump status (on/off) over a complete TVS 

cycle and dividing the result by the complete cycle time. The 

pump status was based on a measurement of the electrical 

power supplied to the pump’s electrical controller. The data 

system’s recording rate was determined by the pump status; 

the data system would switch the recording rate when the 

pump was started and would frequently interrupt (truncate) the 

previous time step. The uncertainty of the pump start time was 

negligible if an interruption was recorded; otherwise, the 

uncertainty was similar to the vent opening uncertainty. Pump 

shutdown always occurred at the same time as the vent valve 

closing discussed in Section 10.2; therefore, the numerical 

integration of the shutdown time step was identical. Because 

the pump was started prior to the delayed opening of the vent 

valve, the pump duty cycle was slightly greater than the TVS 

duty cycle. The uncertainty in the pump duty cycle was similar 

to the TVS duty cycle uncertainty. 

10.4 Turnover Time 

The liquid turnover time  was calculated as the volume of 

liquid contained in the test tank divided by the pump’s 

volumetric flowrate p: 
NASA/TP—2014-216633 19 

p

FVt

   (35) 

Turnover time was used to gauge the time required to 

condition the liquid in the tank by simply flowing all the liquid 

through the heat exchanger. Use of the turnover time is 

admittedly an idealized situation because it assumes that all 

uncooled liquid is first drawn through the pump and heat 

exchanger without any mixing with cooled liquid discharged 

from the heat exchanger via the axial jet and/or spray hoops. 

For the test tank and pump used in the present study,  = 0.65 

hr at 90-percent fill and  = 0.37 hr at 50-percent fill. 

10.5 Vented Mass and Vent Flowrate 

The instantaneous vent flowrate varied during a vent cycle 

and from cycle to cycle. As the tank pressure dropped during 

venting, the flowrate also dropped. Furthermore, if the TVS 

was operated in temperature-control mode, the tank pressure 

minimums and maximums dropped with each additional cycle. 

The amount of mass vented during a cycle was obtained by 

numerical integration (trapezoidal rule) of the flowrate 

readings over a complete cycle. For a number of similar 

cycles, an arithmetic average, , mvent was then calculated. The 

time-averaged vent flowrate per complete TVS cycle mcycle is 

useful for assessing the efficiency of the TVS performance 

and was calculated as 

cycle

vent cycle t

m m (36) 

The average vent flowrate mvent , based only on the time the 

vent valve was open, is 

TVS

cycle

TVS cycle

vent

vent

vent

vent      m

t

m

t

m

m (37) 

10.6 Heat Exchanger Performance 

The instantaneous TVS heat exchanger heat removal rate 

QHX

was calculated as 

  QHX  mvent hout 	 hin (38) 

using the vent flowrate mvent during the vent portion of the 

TVS cycle, the outlet enthalpy hout of the vent flow downstream of the heat exchanger, and the vent flow inlet enthalpy 

hin upstream of the J–T device. The vent flowrate was 

measured by the larger of the two flowmeters and the 

flowmeter data were converted using the methodology 

discussed in Section 5.0. The inlet enthalpy was based on the 

measured test tank pressure and measured temperature 

upstream of the J–T device. In some instances when the tank 

conditions were near saturation (as in Test Series A), the 

pressure and temperature values occasionally returned an 

enthalpy value for vapor. An inferred vapor phase was clearly 

not correct, because the inlet flow was liquid. The error 

resulted from very small measurement errors that were within 

the uncertainty specifications of the sensors. As an errorchecking measure, the data analysis procedure also obtained 

the enthalpy for saturated liquid based on the inlet 

temperature. The procedure then used the lower of the two 

enthalpy values for the inlet. 

The flow downstream of the J–T device was assumed to be 

two-phase and saturated. Thus, the downstream pressure was 

taken to be equal to the saturation pressure corresponding to the 

measured temperature downstream of the J–T device. (Noise in 

the downstream temperature measurement was small, less than 

the uncertainty of the measurement.) Furthermore, the pressure 

drop through the vent side of the heat exchanger was assumed 

to be zero. Modeling showed the vent-side pressure drop to be 

negligible (Ref. 18, personal communication). Therefore, the 

outlet pressure at the heat exchanger on the vent side was the 

same as the saturation pressure downstream of the J–T device. 

The vent flow outlet enthalpy could then be obtained by using 

the measured heat exchanger outlet temperature and pressure. 

The result was always superheated vapor, as expected for a 

well-designed heat exchanger. 

The heat exchanger performance obtained from 

Equation (38) is an instantaneous value. The flowrate and 

various temperatures and pressures are not constant during a 

vent cycle and furthermore change from cycle to cycle. The 

instantaneous value was calculated for each recorded time 

interval. The total heat removed in each vent cycle was 

numerically evaluated using the trapezoidal rule (with the 

same assumptions as used for the venting time) to obtain the 

total heat removed in the cycle. Dividing this result by the 

total vent time tvent gives the average heat removal rate QHX



for the vent portion of a complete TVS cycle: 

vent

vent

d

t

Q t

Q t HX

HX

 



(39) 

The average heat removal rate over a number of successive 

cycles was found by simple arithmetic averaging of the heat 

removal rates over the total number of available cycles. Note 

that this value was averaged over the venting periods only—

not over the entire TVS cycle. 
NASA/TP—2014-216633 20 

10.7 Quality Downstream of Joule-Thomson 

Device 

The expansion at the J–T device was assumed to be 

isenthalpic (h = constant). The quality x of the two-phase flow 

downstream of the J–T device and also at the heat exchanger 

inlet (assuming that no heat transfer occurs between the J–T 

device and the heat exchanger) was calculated as 

fg

in f

h

h h

x 	  (40) 

where hf and hfg are saturated properties at the downstream 

pressure. All other things being equal, a low value is desirable 

because it indicates greater capacity for heat removal as the 

vent flow is vaporized in the heat exchanger. 

10.8 Energy Balance Applied to a Complete 

Thermodynamic Vent System Cycle 

An overall energy balance on the test tank using a control 

volume approach was applied using data for each temperaturecontrolled test segment. This energy balance provides a check 

on the consistency of the test results. If constant temperature is 

assumed, the change in the propellant internal energy is due to 

the change in fill level and the change in helium internal 

energy is negligible. The vented mass is known, and 

depending on the choice of the control volume boundary, the 

exiting mass at the vent is either liquid if the J–T device and 

heat exchanger are outside the boundary or vapor if they are 

inside the boundary. If the heat exchanger is at the boundary 

of the control volume, then the heat exchange at the heat 

exchanger is included in the energy balance. The energy input 

due to environmental heating is approximately known from 

the boiloff tests, and the TVS and pump duty cycles are 

known. The remaining component in the overall balance is the 

energy input due to pump power dissipation. The energy 

balance can be solved for this and the result compared to the 

specification for the pump provided by the manufacturer. The 

two forms of the energy balance gave nearly identical results 

so only the vapor venting model is discussed. 

If the TVS heat exchanger is within the control volume and 

vapor exits the control volume at the liquid temperature and 

vent pressure, 

 

p

cycle cycle

2 1 ,



	  	

 t

m h T P Q

t

U U

Q

vent e

p



(41) 

where the subscript p refers to the pump. 

10.9 Thermodynamic Vent System Efficiency 

VanOverbeke (Ref. 6) defines a TVS efficiency  as the 

ratio of the boiloff rate moto the time-averaged TVS vent 

flowrate mcycle for the same tank system: 

mcycle

mo



  (42) 

Equation (42) is an attractively simple figure of merit, but 

the comparison can be distorted by variations in ullage 

conditions. If one recalls Equation (32), the boiloff rate for a 

given tank heating rate depends on the fill fraction. That is, the 

boiloff flowrate decreases as the fill fraction is reduced 

because a greater fraction of the heat input is directly absorbed 

by the vapor rather than contributing to liquid vaporization. 

From the present boiloff results (Table II), ullage superheating 

was 8 to 9 percent of the total heating at a fill fraction of 0.93 

to 0.94 and 22 percent of the total heating at a fill fraction of 

0.39. What happens in the ullage also impacts the efficiency of 

the TVS: an axial jet system in normal gravity with increasing 

unwetted wall temperatures will have a lower vent flowrate 

than a system that maintains a cold unwetted tank wall region 

because less energy is removed by the TVS in the former case. 

10.10 Thermodynamic Vent System Test 

Results 

The TVS tests were conducted with differing hardware 

configurations and test conditions as summarized in Table I. A 

more detailed tabulation of the test conditions (test segment 

identification (I.D.), test fluid, control logic mode, number of 

cycles, total test time, median pressure, median liquid 

temperature, median vent pressure, median fill fraction, and 

average vacuum level) is given in Table III. Median values 

(average of observed minimum and maximum values) are 

useful for conditions where the sawtooth TVS profiles trended 

upward or downward over multiple cycles. Key results of the 

data analysis (cycle time, pump and TVS duty cycles, vented 

mass per cycle, vent flowrate (when vent was open), and 

cycle-averaged vent flowrate) are listed in Table IV along with 

estimated uncertainties for various parameters. Other TVS test 

results (heat exchanger heat removal rate, quality downstream 

of the J–T device, turnover time, and TVS efficiency) are 

presented in Table V. These results are all discussed in 

Sections 11.0 and 12.0. 
NASA/TP—2014-216633 21 

TABLE III.—DETAILED SUMMARY OF THERMODYNAMIC VENT SYSTEM (TVS) TEST CONDITIONS 

Test Fluid Logic modea Number 

of 

cycles 

Total 

test time, 

hr 

Median 

pressure, 

MPa 

Median 

temperature, 

K 

Median 

vent pressure, 

MPa 

Median 

fill 

fraction 

Average vacuum 

level, 

Pa 

90A_1 O2 Combined 3 9.7 0.60 111.3 0.24 0.87 6.010–4

50A_1 Combined 2 27.7 .57 110.5 .25 .49 4.8 

90B_1 T control 7 14.6 1.47 111.6 .12 .88 5.5 

50B_1 T control 13 18.3 1.62 111.3 .46 5.1 

50B_2 P control 13 15.8 1.60 111.9 .43 5.1 

90C_1 T control 9 23.4 1.40 111.5 .87 4.7 

90C_2 P control 3 13.0 1.32 112.0 .85 4.1 

50C_1 T control 9 18.3 1.60 111.5 .50 5.1 

50C_2 P control 4 9.3 1.57 111.7 .48 4.7 

50C_3 T control 6 12.2 1.52 111.5 .46 4.1 

90D_1 6 16.7 1.47 111.5 .86 3.5 

90D_2 3 8.1 1.36 111.9 .84 

50D_1 8 20.8 1.63 111.6 .51 2.9 

50D_2 P control 3 11.0 1.61 111.8 .49 3.3 

50D_3 T control 3 7.8 1.60 112.1 .48 3.1 

50E_1 N2 T control 

(with added GHe) 

4 5.0 1.59 100.0 .11 .48 8.5 

50E_2 T control 11 13.8 1.57 100.0 .45 6.4 

50E_3 T control 

(with added GHe) 

5 6.3 1.58 100.0 .42 6.4 

50E_4 P control 13 15.0 1.59 100.3 .40 5.5 

a

T, temperature, or P, pressure. 

TABLE IV.—DETAILED SUMMARY OF THERMODYNAMIC VENT SYSTEM (TVS) TEST RESULTS (PART 1) 

Test Average 

cycle 

time, 

hr 

Uncertainty 

cycle time, 

percent 

Pump 

duty 

cycle 

TVS 

duty 

cycle 

TVS duty 

cycle 

uncertainty 

Average 

vented mass 

per cycle, 

kg 

Vent 

flowrate, 

kg/hr 

Uncertainty 

in average 

vented mass 

per cycle, 

percent 

Cycle-averaged 

vent flowrate, 

kg/hr 

Uncertainty of 

cycle-averaged 

vent flowrate, 

percent 

90A_1 3.2 ±0.2 0.70 0.69 ±0.01 8.8 4.0 ±4 2.7 ±4 

50A_1 13.9 ±0.1 .74 .74 ±0.00 36.1 3.5 ±5 2.6 ±5 

90B_1 2.1 .21 .20 ±0.01 4.6 10.9 ±5 2.2 ±5 

50B_1 1.4 .18 .17 2.8 11.5 ±7 2.0 ±7 

50B_2 1.2 .15 .14 1.9 11.7 ±11 1.6 ±11 

90C_1 2.6 .24 .24 6.4 10.3 ±3 2.5 ±3 

90C_2 4.3 .14 .14 ±0.00 5.7 9.8 ±4 1.3 ±4 

50C_1 2.0 .22 .22 ±0.01 5.1 11.6 ±4 2.5 ±4 

50C_2 2.3 ±0.2 .18 .17 4.6 11.4 ±5 2.0 ±5 

50C_3 2.0 ±0.1 .22 .22 4.9 11.1 ±5 2.4 ±5 

90D_1 2.8 ±0.1 .20 .20 5.9 10.7 ±4 2.1 ±4 

90D_2 2.7 ±0.2 .22 .21 5.8 10.1 2.1 

50D_1 2.6 ±0.1 .18 .18 5.3 11.5 2.0 

50D_2 3.7 ±0.2 .14 .14 ±0.00 5.9 11.4 1.6 

50D_3 2.6 ±0.2 .19 .19 ±0.01 5.5 11.3 2.1 

50E_1 1.2 ±0.4 .43 .42 4.6 8.8 3.7 

50E_2 1.3 ±0.1 .35 .35 3.7 8.6 ±3 3.0 ±3 

50E_3 1.3 ±0.2 .32 .31 3.4 8.7 ±4 2.7 ±4 

50E_4 1.2 ±0.1 .25 .24 2.4 8.6 ±5 2.1 ±5 
NASA/TP—2014-216633 22 

TABLE V.—DETAILED SUMMARY OF THERMODYNAMIC VENT SYSTEM (TVS) TEST RESULTS (PART 2) 

Test Logic modea Heat exchanger (HX) 

heat removal rate, 

W 

Uncertainty in HX heat 

removal rate, 

percent 

Average 

quality 

(J–T exit) 

Tank turnover 

time, 

hr 

Pump-on-time to 

turnover-time 

ratio 

TVS 

efficiency 

90A_1 Combined 213 ±4 0.10 0.63 3.61 0.67 

50A_1 Combined 189 ±5 .10 .36 29.05 .69 

90B_1 T control 586 ±5 .16 .64 .70 .83 

50B_1 T control 626 ±8 .33 .77 .90 

50B_2 P control 633 ±11 .31 ------ ---- 

90C_1 T control 561 ±4 .63 1.00 .74 

90C_2 P control 529 ±4 .17 .61 ------ ---- 

50C_1 T control 624 ±5 .16 .36 1.25 .73 

50C_2 P control 615 ±5 .34 ------ ---- 

50C_3 T control 602 ±5 .33 1.35 .76 

90D_1 580 ±4 .62 .90 .85 

90D_2 545 ±4 .17 .61 .97 .85 

50D_1 620 ±5 .16 .37 1.30 .89 

50D_2 P control 619 ±4 .16 .35 ------ ---- 

50D_3 T control 609 .17 .34 1.43 .87 

50E_1 T control 

(with added GHe) 

423 .24 .34 1.55 ---- 

50E_2 T control 414 .33 1.35 ---- 

50E_3 T control 

(with added GHe) 

421 .31 1.31 ---- 

50E_4 P control 415 ±5 .29 ------ ---- 

a

T, temperature, or P, pressure. 

When tests were conducted at more than one fill level, the 

higher fill level was tested first; then the test tank was drained to 

the lower fill level of the second test. For each of the test 

segments, a figure showing the resulting temperature and 

pressure cycles is provided (Figs. 18, 21, 25, 27, 30, 34, 38, 41, 

and 43). The entire histories are shown, with unanalyzed 

regions in lighter grey and the analyzed portions of the curve in 

darker colors. The tank pressure and average liquid 

temperatures used in the detailed analyses are shown. For each 

test, a second figure is provided that shows the measured liquid 

temperatures at a few selected locations from near the tank 

bottom to just below the liquid free surface (Figs. 19, 22, 26, 29, 

31, 35, 39, 42, and 44). In these figures, the plotted liquid 

temperature is a centered 11-point running average (average of 

11 measurements in time—five before and five after the plotted 

time). This was done to remove scatter due to signal noise. In 

addition, the curves are slightly adjusted with small temperature 

offsets to force the curves to align vertically, thus allowing 

easier visual comparison of the results. The temperature offsets 

are very small, much less than the tolerance bands specified by 

the sensor manufacturer, and are the same for each figure. The 

temperature range of these figures also is small; using the 

offsets greatly improves clarity. The data in the figures showing 

liquid temperatures from near the tank bottom to just below the 

liquid free surface (Figs. 19, 22, 26, 29, 31, 35, 39, 42, and 44) 

are for up to six consecutive cycles observed during 

temperature-controlled TVS logic (if pressurized with GHe). 

Finally, some figures (Figs. 20, 23, 32, and 36) show a few 

examples of the temperatures near the liquid-ullage interface, 

and other figures (Figs. 24, 33, 37, and 45) show the ullage, 

tank lid, and unwetted tank wall temperatures observed in some 

of the tests. Only a few representative examples are provided 

because the results were often similar for various tests. 

11.0 Discussion of Results for Each 

Individual Thermodynamic Vent 

System Test 

11.1 Test 90A 

The LO2 was saturated at the total tank pressure of about 

0.6 MPa and saturation temperature of 111 K; no GHe 

pressurant was present. The pump discharge was routed 

through the axial jet. Plots of the tank pressure and average 

liquid temperature are shown in Figure 18. After the initial 

period of about 13 hr in which the TVS heat exchanger was 

operated extensively, three complete cycles were performed 

with a 0.034-MPa pressure-control band. There was a rapid 

pressure drop early in the vent cycle, attributed to breakup of 
NASA/TP—2014-216633 23 

Figure 18.—Tank pressure and average liquid 

temperature for Test 90A. 

the warm thermal layer at the liquid surface by the axial jet 

flow. The initial pressure reduction was far larger than for any 

other tests in the present work and was attributed to the 

augmented thermal stratification at the liquid-vapor interface 

because of the absence of helium (liquid was not subcooled) 

and the high fill level (enhanced free convection due to tall 

wetted sidewall heating). The remaining pressure drop during 

the vent cycle occurred as the heat exchanger cooled the bulk 

liquid, reducing the temperature and the corresponding 

saturation pressure. As expected, the liquid temperature cycles 

were well synchronized with the pressure cycles. The results 

demonstrate the TVS’s ability to control tank pressure and 

liquid temperature. 

The test pressure for Test Series A was far below the TVS 

design point, so the heat exchanger performance was 

substantially reduced for Test Series A because of the low vent 

flowrate of 4 kg/hr. During the vent cycle, the averaged heat 

removal rate of the TVS was 213 W (compared with the design 

point of about 600 W), requiring a high duty cycle of 0.69. 

Also, the average liquid temperature signal was noisy, as seen in 

Figure 18. Once the noise was observed, efforts were begun to 

modify the data system software. Furthermore, the TVS cycles 

were lengthier than desired if the test schedule was to be 

maintained. Because the goal was to demonstrate 6 to 10 cycles 

per test, it was necessary to reduce the control band for pressure 

(or temperature). When control with a narrower band was 

attempted (beginning at 23 hr), variations in the cycle peaks 

were apparent. Some of the variability was attributed to the 

noise in the temperature and pressure signals, as well as to some 

of the delay times in the control logic being too long. 

As seen in Figure 19, there was little difference in liquid 

temperatures at various heights within the liquid. At all sensor 

locations, the liquid temperature responded in a nearly identical manner to the pressure cycle. Although the figure indicates that temperature maximums and minimums increased 

with time, the increase was on the order of only a few 

Figure 19.—Liquid temperatures during Test 90A 

thermodynamic vent system (TVS) cycles. 

Figure 20.—Liquid-ullage interface temperatures during 

Test 90A thermodynamic vent system (TVS) cycles; 

Tsat, temperature of saturated vapor. 

hundredths kelvin. The author does not consider this increase 

to be significant because of the signal noise and logic delay 

time issues mentioned in Sections 6.0, 7.0, and 10.1. The axial 

jet configuration did not provide ullage and unwetted wall 

thermal control capability in the normal-gravity test environment. Although not shown, the tank lid and ullage temperatures increased steadily throughout the test. Figure 20 shows 

the thermal stratification near the liquid-ullage interface. The 

temperature sensor at 93-percent fill was clearly in the ullage 

for all TVS cycles shown. The sensor at 92-percent fill 

appeared to be initially in the liquid, but it was in the ullage 

during the last cycle shown, thus indicating that the liquid 

level had dropped due to TVS venting. The dashed 
NASA/TP—2014-216633 24 

line in the figure is the saturation temperature obtained from 

the tank pressure measurement. The saturation temperature 

should represent the temperature of the liquid-ullage interface 

of a fluid system containing only pure liquid and vapor. The 

maximum-to-minimum saturation temperature difference (or 

maximum liquid thermal stratification) for these TVS cycles 

was slightly less than 1 K. The peak temperature measured by 

the sensor at 90-percent fill increased with successive TVS 

cycles as the liquid level dropped and became closer to the 

sensor’s position. 

11.2 Test 50A 

Test 50A was conducted at about 50-percent fill and without 

GHe pressurant. The pressure and liquid temperature cycles 

are shown in Figure 21. Two complete cycles were obtained 

from 11 to 40 hr. The duty cycle of 0.74 and heat exchanger 

performance of 189 W are similar to those for Test 90A, as 

expected. The slightly larger duty cycle and slightly lower 

heat removal rates are due to the average tank pressure being 

lower than for the previous test (0.57 vs. 0.60 MPa). Although 

the pressure control band is the same as for Test 90A and the 

duty cycle is about the same, the cycle time is much greater 

(13.9 vs. 3.2 hr). Comparing Figures 18 and 21 shows that the 

pressure rise rate was much slower at 50-percent fill than at 

90-percent fill and that the pressure drop due to the mixeroperation-only prior to TVS venting was less. Both results 

from the comparison indicate that less thermal stratification 

developed in each cycle at the lower fill level. The cycle 

control band (pressure or temperature) would have to be 

reduced to obtain the desired number of cycle repeats for the 

given test schedule. When control with a narrower band was 

attempted, beginning at 42 hr, the noisy temperature signal 

was clearly not acceptable and noise in the pressure signal also 

became apparent. 

Figure 22 shows the cyclic behavior of the liquid 

temperature at various fill locations. The response was nearly 

identical at each sensor location, and the variations in the 

minimums and maximums are again thought to be insignificant. A key difference for Test 50A was the much larger 

liquid maximum-to-minimum temperature swing per cycle. It 

was about 0.5 K for Test 50A in comparison to 0.15 K for Test 

90A. The reduced wetted sidewall height for Test 50A 

resulted in less warm liquid accumulation at the liquid-vapor 

interface, a longer heating period, and reduced pressure drop 

during the initial portion of the mixing/vent cycle. The total 

cycle time for Test 50A was substantially longer than for any 

other tests reported herein. Figure 23 shows measured 

temperatures near the liquid-ullage interface. An interesting 

observation for Test 50A is the significant drop in liquid level 

from the first to second TVS cycle, as shown. The sensor at 

60-percent fill was clearly in the ullage during both cycles. 

The sensors at 55-, 54-, and 53-percent fill were submerged in 

the liquid during the first cycle and then exposed to the ullage in 

Figure 21.—Tank pressure and average liquid 

temperature for Test 50A. 

Figure 22.—Liquid temperatures during Test 50A 

thermodynamic vent system (TVS) cycles. 

Figure 23.—Liquid-ullage interface temperatures during 

Test 50A thermodynamic vent system (TVS) cycles; Tsat, 

temperature of saturated vapor. 
NASA/TP—2014-216633 25 

Figure 24.—Ullage, lid, and unwetted wall 

temperatures during Test 50A. 

the second cycle. The maximum thermal stratification in the 

liquid, as denoted by the saturation temperature curve, was 

slightly less than 1 K, which is similar to the result at 90-percent 

fill, but for a much longer cycle time (13.9 vs. 3.2 hr). 

Figure 24 shows that the tank lid, ullage, and unwetted wall 

temperatures rose steadily during Test 50A, as one might 

expect with an axial jet mixer in normal gravity. At the 

conclusion of Test 50A, data system software with the 

enhanced signal-conditioning features described in Section 7.0 

was ready for use for all remaining test series. Both tests for 

the Test Series A demonstrated the linked behavior of the 

pressure and temperature cycles for a tank with a single 

component (pure vapor) ullage. The axial jet mixer appeared 

to mix the liquid contents well, but in normal gravity, it did 

not control ullage or unwetted wall temperatures. 

11.3 Test 90B 

Test Series B used the same axial jet hardware configuration 

as Test Series A. The LO2 temperature was again about 111 K, 

but the tank was pressurized with GHe to a total tank pressure 

of about 1.5 to 1.6 MPa. Test 90B conducted at approximately 

90-percent fill resulted in three sets of temperature control 

cycles, with GHe repressurization after the first and second 

sets. Much of the initial pressure drop after pressurization or 

repressurization was attributed to cooling (pressure collapse) 

of the warm GHe pressurant when it was introduced into the 

colder test tank. Figure 25 shows the pressure and temperature 

cycles. Two key results are readily seen. First, the signal 

conditioning worked well, resulting in clean, repeated cycles. 

A larger number of cycles were successfully completed in a 

shorter timeframe, roughly a cycle every 2 hr. Second, the 

pressure slowly dropped with each cycle repeat while the 

liquid temperature was controlled within the lower and upper 

set points, as expected based on the thermodynamic model 

presented in Section 1.2. 

The third set of TVS cycles was analyzed, with the results 

summarized in Tables III to V. The test conditions were near 

Figure 25.—Tank pressure and average liquid 

temperature for Test 90B; GHe, gaseous helium. 

Figure 26.—Liquid temperatures during Test 90B 

thermodynamic vent system (TVS) cycles. 

the TVS heat exchanger design point as reflected by a duty 

cycle of 0.20 and a heat exchanger performance of 586 W. 

Because the tank pressure was much higher than for Test 

Series A, the vent flowrate was much higher (10.9 vs. 3.5 to 

4.0 kg/hr). The temperature control band was 0.17 K with a 

corresponding 0.08 MPa swing in pressure. The recorded 

temperature swing was slightly greater because of a small 

upward spike in temperature at mixer startup. The cause was 

not investigated, but the spike may have been due to rapid 

vaporization of cold liquid being pumped through external 

valves and lines that had warmed up while the pump was off. 

As shown in Figure 26, the temperatures were again very 

uniform at the various sensor locations within the ullage. 

Although not shown, the tank lid and ullage temperatures for 

Test 90B slowly increased throughout the test. The lid and 

ullage temperatures were significantly higher than the liquid 

temperature because the axial jet (in a normal-gravity 

environment) was unable to cool these locations. 
NASA/TP—2014-216633 26 

11.4 Test 50B 

Test 50B followed Test 90B, after which the tank was 

drained to ~50-percent fill, exposing the cold upper tank 

sidewall to the ullage. The tank was then repressurized with 

GHe. In Test 50B, the upper tank wall, tank lid, and ullage 

temperatures increased throughout the test. As the unwetted 

tank sidewall warmed up, the ullage temperature rose as well, 

which in turn increased the partial pressure of the GHe. The 

ullage warming effect was opposed by the operation of the 

TVS, which decreased the GHe partial pressure by expanding 

the ullage. Initially in Test 50B, the ullage warming effect 

appeared to dominate. Eventually a tipping point was reached 

(as shown in Fig. 27) after 30-plus hours of operation, and the 

tank pressure decreased while operating the TVS in temperature control mode. The observed normal-gravity tank 

pressure behavior in Test 50B could be different from what 

would occur in low gravity. 

A portion of the data from Figure 27 is magnified in 

Figure 28, where seven complete TVS cycles under temperaturecontrolled logic are shown. The temperature maximums and 

minimums were stable, whereas the pressure cycles exhibited a 

repeatable cyclic behavior with a declining trend of pressure. The 

liquid temperature distribution is shown in Figure 29 and is 

similar to the axial jet test results presented in Section 11.3. 

After demonstrating that the TVS could reduce tank pressure 

using temperature-control logic, a second test with pressurecontrol logic was performed. Test 50B was the first test to 

implement pressure-control logic with GHe in the tank. As seen 

in Figure 27, beginning at 65 hr, the tank pressure was kept 

within the pressure-control set points and the average liquid 

temperature increased as expected, based on model predictions. 

The duty cycles and the cycle-averaged vent flowrate were 

lower under pressure control than under temperature control 

because less energy removal is required (via TVS venting) 

when the tank and liquid are permitted to warm up. 

11.5 Test 90C 

In Test Series C, the pump discharge was split, with an 

unknown portion routed through the axial jet and the remainder 

through the two spray hoops near the top of the tank. For the 

approximate 90-percent fill test, about four temperaturecontrolled cycles were initially completed. Then the test tank 

was repressurized, and the TVS was operated uninterrupted in 

temperature-control mode from about 15 to 40 hr, as shown in 

Figure 30. After that, the control logic was switched to 

pressure control for three cycles. The same trend was 

observed, as seen before in Test 50B under pressure control: 

the liquid temperature rose while the pressure cycled within 

the control band. 

Figure 27.—Tank pressure and average liquid 

temperature for Test 50B. 

Figure 28.—Closeup of seven temperature-controlled 

pressure and temperature cycles for Test 50B. 

Figure 29.—Liquid temperatures during Test 50B 

thermodynamic vent system (TVS) cycles. 
NASA/TP—2014-216633 27 

Figure 30.—Tank pressure and average liquid 

temperature for Test 90C. 

Figure 31.—Liquid temperatures during Test 90C 

thermodynamic vent system (TVS) cycles. 

Figure 32.—Liquid-ullage interface temperatures during 

Test 90C thermodynamic vent system (TVS) cycles. 

Figure 33.—Ullage and lid temperatures during 

Test 90C. 

Figure 31 shows a sample of the liquid temperature 

behavior. The plot looks very similar to those seen in the 

previous axial jet tests, but with an indication of some shortterm randomness at the individual locations. The randomness 

could be caused by the diminished strength of the axial jet 

flow because some of the pump discharge was going through 

the spray hoops. Figure 32 shows the measured temperatures 

near the liquid-ullage interface. The sensor at 91-percent fill 

appears to have been in the ullage for all cycles. The sensor at 

90-percent fill was initially in the liquid (or very near the 

liquid) but was in the ullage at the end of the cycles. The 

sensor at 89-percent fill may have become exposed to the 

ullage during the last cycle. The temperature scale was much 

smaller than in the similar figures for Test Series A (pure 

liquid and vapor). Also, the temperatures on the liquid side of 

the interface were essentially the same as for the bulk liquid 

(the 85-percent sensor reading is also used in Fig. 30). 

Therefore, there was minimal thermal stratification in the 

liquid. The reduction in stratification may have been partially 

due to the shorter cycle time in comparison to Test Series A, 

but it was more likely due to the presence of helium in the 

ullage and the resulting subcooled liquid. 

Figure 33 confirms flow through the spray hoops. The lid 

temperature, measured by a thermocouple on the outside 

surface of the lid cyclically varied as the pump was turned on 

and off. The silicon-diode sensor located at the 99-percent fill 

location warmed when the pump was off and cooled to the 

liquid temperature (111 K) when the pump was on. The 

minimums and maximums of both sensors were steady over 

time, indicating the ability of the spray hoops to control ullage 

and lid temperatures and maintain steady-state thermal 

conditions. The lid temperature remained under 200 K for Test 

90C; whereas in the axial-jet-only tests, the lid temperature 

was higher (e.g., ~220 K in Fig. 24, Test 50 A). 
NASA/TP—2014-216633 28 

11.6 Test 50C 

Smooth operation of the TVS was demonstrated in Test 50C, 

as shown in Figure 34. Nine temperature-controlled cycles were 

conducted first, followed by four pressure-controlled cycles. 

Then the control logic was switched back to the initial 

temperature-control set points. Because the liquid temperatures 

had increased during the pressure-controlled operation, the TVS 

had to vent steadily for a longer time to cool the liquid and 

return to the minimum-temperature set point. Once the 

minimum set point was reached, the TVS operated through six 

more temperature-controlled cycles. The cycle times, duty 

cycles, and vented mass for the first and third segments (both 

temperature-controlled) are very similar as seen in Table IV. In 

Figure 35, occasional short-term randomness in the liquid 

temperature at various locations again occurred. Otherwise, the 

behavior is very similar to that seen when the entire pump 

discharge was routed through the axial jet. 

Figure 36 shows the liquid-ullage interface temperatures. 

Results are similar to the Test 90C results discussed in Section 

11.5. The drop in liquid level is apparent because the 

temperature at 52 percent rose when the sensor was exposed to 

the ullage. Again, there was very little thermal stratification on 

the liquid side, which the author attributes to the shorter cycle 

times and liquid subcooling due to helium pressurization. 

The effect of the spray hoops is evident in Figure 37. The 

lid temperature cycled as the pump was turned on and off, as 

did the ullage temperature at the 99-percent fill location. Two 

additional temperatures are shown, both closer to the liquid 

surface. The sidewall temperature shows a very slight cyclic 

oscillation as does the ullage temperature at the 80-percent fill 

location. Both ullage temperatures dropped to the liquid 

temperature (111 K) when the pump was running. The longterm trend is one of steady cyclic oscillations, thus indicating 

the ability of the spray hoops to control and maintain the 

thermal conditions in the upper tank region. 

Figure 34.—Tank pressure and average liquid 

temperature for Test 50C. 

Figure 35.—Liquid temperatures during Test 50C 

thermodynamic vent system (TVS) cycles. 

Figure 36.—Liquid-ullage interface temperatures during 

Test 50C thermodynamic vent system (TVS) cycles. 

Figure 37.—Ullage, lid, and unwetted wall 

temperatures during Test 50C. 
NASA/TP—2014-216633 29 

11.7 Test 90D 

In Test Series D, all pump discharge flow was routed 

through the spray hoops in the ullage. The test at ~90-percent 

fill started with temperature-controlled TVS cycles, which 

were executed successfully. When control was switched to 

pressure-controlled logic, a few pressure cycles were obtained, 

but the average liquid temperature climbed steadily, as seen in 

Figure 38. The attempt at pressure control in Test 90D was the 

only occurrence in the entire test program where the TVS did 

not operate successfully. The unsuccessful result was not 

totally unexpected because active mixing of the liquid was due 

solely to liquid intake at the pump inlet and droplets falling on 

the liquid surface. Results may have been different if a wider 

pressure control band were tried. However, no further testing 

was attempted with pressure control at 90 percent. A second 

segment of temperature-controlled cycles was then performed. 

Figure 39 shows the liquid temperatures for six representtative temperature-controlled cycles. The results are clearly a 

bit different from any of the previous tests. The temperatures 

are still cyclic and repeat themselves quite well, but the upper 

locations show upward spikes when the pump is started (at the 

cycle peaks) and the bottom location shows an unusual lower 

minimum. The temperature at 5-percent fill seems to have 

twice the amplitude of the temperature at 45 percent, but there 

is no overall long-term divergence or convergence of the 

measured liquid temperatures. The lid and ullage temperatures for Test 90D were very similar to those for Test 90C. 

Performance specifications are given in Tables IV and V. 

Duty cycles and vented mass are slightly lower than for the 

combined axial jet and spray hoops test (Test 90C). 

Figure 40 provides more detail about the pressure-controlled 

portion of Test 90D. During the first self-pressurization period, 

the liquid temperatures became equal and increased at the same 

rate. Just before the first cycle peak, some thermal stratification 

began to develop. The first and second vent cycles did not 

destratify the liquid because the thermal gradient from 5- to 

85-percent fill persisted. The third vent cycle with a greater 

pressure drop (the set points were changed and the control logic 

switched to temperature control as the vent cycle was in 

progress) shows both a temperature reduction and some 

destratification. The incomplete destratification was not seen 

in other tests and is an indication that spray-hoops-only 

configuration was not totally effective in mixing the liquid 

sufficiently to achieve isothermal conditions. 

11.8 Test 50D 

Next, the tank was drained to 50 percent, and the sprayhoops-only configuration was tested first with temperature 

control logic, then with pressure control, and finally with 

temperature control again. The TVS appeared to 

Figure 38.—Tank pressure and average liquid 

temperature for Test 90D. 

Figure 39.—Liquid temperatures during Test 90D 

thermodynamic vent system (TVS) cycles. 

Figure 40.—Pressure and liquid temperatures for 

pressure-controlled operation in Test 90D. 
NASA/TP—2014-216633 30 

perform satisfactorily in each segment. Figure 41 shows the 

pressure and liquid temperature cycles, and Figure 42 shows 

the individual temperature histories for six representative 

liquid locations. The trends are similar to those seen with the 

same hardware configuration at 90-percent fill (Test 90D). 

The upper sensors show a temperature spike where the pump 

turned on. The spike might have been caused by warmer liquid 

falling on the liquid surface as a result of spray droplet contact 

with the tank lid and upper tank wall. The lowest sensor shows 

a wide peak-to-peak swing in temperature. There does not 

seem to be any long-term divergence or convergence of the 

temperature cycles; the cyclic results look repeatable and 

consistent. Table III shows similar performance results for the 

two temperature-controlled test segments. The duty cycles and 

vented mass are slightly lower than for Test 90D. 

Figure 41.—Tank pressure and average liquid 

temperature for Test 50D. 

Figure 42.—Liquid temperatures during Test 50D 

thermodynamic vent system (TVS) cycles. 

11.9 Test 50E 

At the conclusion of the LO2 test program, time was granted to 

conduct a single LN2 test. The LN2 test was performed at the 

nominal 50-percent fill level and a liquid temperature of 100 K. 

GHe pressurization was used to obtain tank pressures near 

1.6 MPa, and the hardware was configured for tank mixing by 

the axial jet only. Figure 43 shows the pressure and liquid 

temperature cycles. Both temperature- and pressure-controlled 

tests were performed. A unique result for Test Series E was the 

simultaneous control of both pressure and temperature with an 

ullage containing GHe. Simultaneous control was demonstrated 

with four TVS cycles at the start followed later by another five 

cycles beginning at about 22 hr. During simultaneous pressure 

and temperature control, the TVS was operated using temperature control logic. When the TVS was venting, the test 

facility operator would carefully monitor the decreasing 

pressure. When the pressure reached a specified lower limit, the 

operator would manually control additional GHe flow to 

maintain the tank pressure at the lower limit while the TVS 

continued to vent and cool the liquid. When the liquid 

temperature reached the lower set point, the TVS venting would 

stop automatically and the operator would simultaneously stop 

adding GHe to the tank. As seen in Figure 43, both temperature 

and pressure remained within the control bands during these test 

segments. The separate temperature- and pressure-controlled 

segments of the test behaved very much the same as the 

previous LO2 tests with GHe. When the pressure-controlled 

test segment ended at a lowered tank pressure (at 21 hr), GHe 

was added to restore the tank pressure to the initial value at the 

start of Test Series E. 

The liquid temperatures plotted in Figure 44 look very similar 

to those for the other tests using the axial jet. The LN2 test series 

(Test Series E) was the only one where the lid, unwetted wall, 

Figure 43.—Tank pressure and average liquid 

temperature for Test 50E; GHe, gaseous helium. 
NASA/TP—2014-216633 31 

and ullage temperatures decreased as the tests progressed, as 

shown in Figure 45. The upper tank temperatures decreased 

because the tank was warm prior to filling and the tank was 

only filled to the 50-percent level—unlike the other 50-percent 

fill tests, where the tank was drained from the 90-percent 

level. Therefore, the upper tank continued to cool as the tests 

were conducted. As shown in Table III, the performance 

specifications for the first and third segments (the two 

segments with simultaneous temperature and pressure control) 

show some differences, such as duty cycles and vented mass. 

The differences are attributed to the initial unsteady thermal 

conditions associated with the cooling of the upper tank early 

in the test. 

Figure 44.—Liquid temperatures during Test 50E 

thermodynamic vent system (TVS) cycles. 

Figure 45.—Ullage, lid, and unwetted wall 

temperatures during Test 50E. 

12.0 Observations of All Thermodynamic 

Vent System Test Results 

The results in Tables IV and V clearly indicate better TVS 

performance at higher tank pressure. The heat exchanger 

performance is about 3 times better for Test Series B, C, and 

D than for Test Series A. Thus, the duty cycles are much 

lower for Test Series B, C, and D than for Test Series A. The 

improvement is due primarily to the higher vent flowrate at 

the higher tank pressure: the greater the vented mass, the 

greater the capability of the vent flow to remove energy from 

the tank and the lower the required heat exchanger duty cycle. 

Figure 46 shows the effect of tank-to-vent pressure drop on 

vent flowrate. All of the LO2 tests exhibit a common trend of 

increasing flowrate with pressure drop. For the same pressure 

drop, the vent mass flowrate of the LN2 tests was about 

25-percent lower. The volumetric flowrate in the downstream 

vent line where the vent gas had warmed to room temperature 

was approximately equal for the LO2 tests pressurized with 

GHe and the LN2 tests. The higher mass flowrate for O2 was 

due primarily to its greater density. The flowrates plotted in 

Figure 46 are based on the flowmeter reading in the vent line. 

In general, the agreement between vent flowrates determined by flowmeter readings were in good agreement with 

calculated values based on the tank weighing system, as 

shown in Figure 47. Two notable exceptions are visible in 

Figure 47: Tests 50E_1 and 50E_2—the first two tests performed with LN2. The cause of the discrepancy was not 

determined clearly, and it is not known which of the two vent 

flowrate measurement methods was better. The most probable 

cause of error was considered to be startup transient behavior 

because the first and second tests had discrepancies (with the 

first being larger), whereas the third and fourth LN2 tests 

showed good agreement. Because of the uncertainty in the 

LN2 tests, the LN2 flow vent flowrate data should be considered with caution. 

The heat removal rate of the heat exchanger was 

proportional to the vent mass flowrate and the enthalpy change 

Figure 46.—Effect of pressure drop on vent flowrate. 
NASA/TP—2014-216633 32 

Figure 47.—Comparison of flowrates measured by vent 

flowmeter and tank weighing system. 

Figure 48.—Heat exchanger heat removal 

rate versus vent mass flowrate. 

of the vented fluid, as shown in Figure 48. The data points 

with a heat removal rate above 500 W are for Test Series B, C, 

and D (LO2 tests with GHe pressurization). The data pair near 

200 W are for Test Series A (LO2 tests without GHe 

pressurization). The points slightly above 400 W are for Test 

Series E (LN2). The N2 data fall below the line fitted through 

the O2 data because of the lower specific enthalpy change of 

the vent flow across the heat exchanger for N2 compared with 

O2 (about a 10-percent difference). The quality downstream of 

the J–T device was 0.10 for Test Series A; 0.16 to 0.17 for 

Test Series B, C, and D; and 0.24 for Test Series E. The lower 

quality for Test Series A was a result of operating the vent 

system at a slightly higher back pressure than for the 

remainder of the test program. 

Figure 49.—Relationship between heat exchanger heat 

removal rate and thermodynamic vent system (TVS) duty 

cycle for temperature-controlled TVS operation. 

Figure 49 shows the relationship between TVS heat removal 

rate and TVS duty cycle. Only data points for the temperaturecontrolled tests are shown. Close inspection of Table IV 

reveals that the duty cycles were always less for pressure 

control than for temperature control for otherwise similar test 

conditions (e.g., compare Tests 50B_1 and 50B_2, Tests 

90C_1 and 90C_2, Tests 50C_1 and 50C_2, or Tests 50D_1 

and 50D_2). Because the liquid temperature rose under 

pressure control, less heat was removed, thus allowing shorter 

duty cycles. When two sets of temperature-controlled operation were conducted, similar results were obtained for duty 

cycles and vented mass per cycle (e.g., Tests 50C and 50D). 

The sole exception was in Test 50E, which may be related to 

unsteady test conditions during the early part of Test Series E, 

as discussed in Section 11.9. 

Further examination of Tables IV and V shows that the TVS 

performance (duty cycles, cycle-averaged vent flowrate, and 

heat removal rate) for a given hardware configuration was 

essentially the same for the high and low fill levels when 

temperature-controlled operation was used (e.g., Test Series 

A, B, C, and D). No significant effects of fill level were found. 

There was a subtle difference in performance for the various 

hardware configurations. The combined axial jet and spray 

hoops configuration had the largest TVS duty cycles and timeaveraged vent flowrates. All the data were carefully examined, 

and no attributable differences in vacuum level, vacuum 

chamber wall temperature, or other environmental variables 

were found. With the combined mode used as the basis for 

comparison, for the axial jet, the TVS duty cycle was about 

16-percent lower and the time-averaged vented mass was 

about 14-percent lower. The time-averaged vented mass was 

expected to be less for the axial jet configuration because the 

temperature of the unwetted tank wall, lid, and ullage were not 

controlled with an axial jet in the 1g environment. In all the 

LO2 axial jet tests, these temperatures were increasing with 
NASA/TP—2014-216633 33 

time; thus not all the tank’s energy input was removed and the 

vented mass was less than if complete temperature control had 

occurred. 

Similarly, when the performance of the spray-hoops-only 

configuration was compared with that of the combined mode, 

the TVS duty cycle was about 13-percent lower and the timeaveraged vented mass was about 14-percent lower. However, 

the difference is not due to thermal control of the upper tank 

region because temperatures here exhibited essentially identical behavior for these two configurations. The most probable 

explanation is incomplete mixing and thermal control of the 

liquid during operation of the spray-hoops-only configuration. 

The apparently successful operation of the TVS with the 

spray-hoops-only configuration was somewhat surprising. The 

high heat input rate for the tests may have promoted rather 

vigorous free convection within the test tank. Perhaps the free 

convection provided partial mixing while not achieving isothermal liquid temperatures. 

When the overall energy balance was applied to the results, 

the average pump power dissipation was 75 W for the axial jet 

(Test Series B), 160 W for the combined axial jet and spray 

hoops (Test Series C), and 85 W for the spray-hoops-only 

configuration (Test Series D). There is no cause for the actual 

power dissipation to vary for these test series. The value for 

the combined mode is probably the closest to the true value 

because the combined mode represents the most likely of the 

three test series to be at uniform thermal conditions, as 

assumed in the model. For Test Series B, the missing thermal 

energy increase was probably storage in the upper tank region; 

whereas for Test Series D, the missing amount could have 

been the thermal energy increase in the liquid (due to 

incomplete mixing/cooling) that was not observed directly. 

For the LO2 tests using temperature control, Table V 

provides additional performance metrics. The turnover time, 

computed from Equation (35) is listed, as well as the pump-on 

time (equal to p tcycle). The ratio of pump-on time to turnover 

time is also provided. One observation is that the ratio was 

much greater than unity for Test Series A; this is due to (1) the 

reduced performance of the heat exchanger resulting from the 

lower vent flowrate and (2) the larger control band for 

temperature. Note that, for a fair comparison, the ratio of 

pump-on time to turnover time should only be compared for 

tests using the same control bands because using a wider 

control band increases the pump-on time. For Test Series B to 

E (all with temperature-control bandwidths of 0.17 K), the 

ratio of pump-on time to turnover time was roughly equal to 1, 

ranging from 0.7 to 1.4. Therefore, the combined mixing and 

cooling process may be viewed as largely due to the action of 

pumping the entire liquid volume through the heat exchanger. 

The ratio of pump-on time to turnover time was lower at the 

high fill-level tests than at the lower fill-level tests. The reason 

for fill-level dependence is not known. One thought is there 

may have been a significant time delay in the startup 

performance of the heat exchanger. The delay would have a 

greater impact on the low fill-level tests where the turnover 

time was shorter. 

The axial jet configuration was most efficient, with ratios 

significantly less than 1, an indicator that the bulk liquid was 

well mixed by the jet flow. The ratios were larger for Test 

Series C and D (combined mode and spray hoops) than for 

Test Series B (axial jet) because some or all of the mixing 

flow was routed through the spray hoops. The liquid droplets 

impinging on the unwetted tank wall and lid absorbed heat 

before falling to the liquid surface. The thermal energy gain of 

the falling droplets was added to the liquid; thus, longer TVS 

and pump-on time were required. 

The uncertainty estimates provided in Tables IV and V 

show typical values of 0.1 to 0.4 percent for the cycle time, 

0.01 for the duty cycle, 4 to 5 percent for the vent mass and 

vent flowrate, and 4 to 5 percent for the heat removal rate. The 

exception is Test Series 50B, where larger uncertainties 

resulted from a slower than desired data-recording rate. The 

recording rate was similar to those for the other LO2 test series 

(Test Series A, C, and D), but the significantly shorter cycle 

times for Test Series 50B led to larger uncertainties in the 

computed performance metrics. 

The dissipated pump power calculated from the energy 

balance (Eq. (41)) for the temperature-controlled test segments 

was reasonably consistent for axial jet (Test Series A and B), 

spray-hoops-only (Test Series D), and combined axial jet and 

spray hoops operation (Test Series C). The combined mode 

had the highest calculated value (average of all pertinent test 

segments) of 159 W, which the author considered to be the 

most correct value for the pump power because the tank 

conditions are thought to be mostly isothermal as assumed in 

the energy balance analysis. The value for the axial jet tests 

was 75 W; a value that is probably reduced because of the 

energy stored in the tank lid and unwetted wall. (Lid and 

unwetted wall temperatures increased during the axial jet 

tests.) The value for the spray-hoops-only tests was 84 W and 

is an indication of energy storage somewhere in the system. A 

likely explanation is that the storage was in the liquid because 

the liquid may not have been well mixed with spray hoops 

circulation only. The energy balance method was not applied 

to the LN2 tests (Test Series E) because a boiloff test was not 

conducted with LN2 and the tank heating rate (when filled 

with LN2) was unknown. 

The last column in Table V gives the TVS efficiency based 

on Equation (42). To make the comparison more “fair,” an 

adjusted boiloff flowrate was calculated assuming that the 

second and third terms on the right-hand side of Equation (32) 

were equal to zero for the same tank heating rate. The 

resulting boiloff flowrates for the LO2 tests were 1.85 to 

1.87 kg/hr. TVS efficiencies were computed only for 

temperature-controlled tests. The efficiency was poorest for 

Test Series A. The author attributes this to the inefficient heat 
NASA/TP—2014-216633 34 

exchanger performance at the off-design operating condition 

(low tank pressure) and the resulting high pump/TVS duty 

cycles, where an increased amount of pump energy was 

dissipated into the tank. Much improved efficiencies for the 

axial jet configuration were obtained in Test Series B. The 

higher efficiency at the lower fill level was probably due to 

greater unwetted wall heating (less energy removed by the 

TVS). The efficiencies for Test Series C (combined mode) 

were not as good as for Test Series B (axial jet). The lower 

efficiency for Test Series C was due to the compete cooling of 

the entire tank and contents and the greater amount of energy 

removal. Test Series D (combined mode) had better efficiency 

than Test Series C. The Test Series C to Test Series D 

efficiency comparison may be misleading if the liquid was not 

well mixed by the spray-hoops-only configuration as 

speculated. 

13.0 Comparison of Test Results With 

Thermodynamic Vent System Model 

Results from several test sets were selected for comparison 

with the model presented in Section 1.2. Test Series C was 

selected because the combined use of the axial jet and the spray 

hoops provided the best assurance that the tank contents were 

isothermal as the model assumes. The pressure data for the first 

temperature-controlled set from Test 90C (~90-percent fill) is 

shown in Figure 50 along with the predicted minimum pressures 

for each cycle. The model input parameters were the initial 

pressure, initial fill level, temperature (constant), pump duty 

cycle, environmental heat input rate, and the pump power 

dissipation rate. The agreement between the model and experimental results is quite good, with the model predicting a slightly 

faster rate of pressure drop. In Figure 51, the next portion from 

Test 90C is shown. Here, the pressure was kept constant, 

causing the temperature to rise. The model predicted a slightly 

lower temperature rise rate, but again the agreement with the 

data was quite good. 

Similar comparisons of the model predictions with the 

experimental results from Test 50C (~50-percent fill) are 

shown in Figure 52 for a temperature-controlled segment and 

in Figure 53 for a pressure-controlled segment. The agreement 

is very good for both cases. Again, all values input to the 

model were matched to the test values. 

The last comparison was for the nitrogen Test 50E (axial jet, 

~50-percent fill) using combined pressure and temperature 

control. Figure 54 shows the experimental fill levels for the 

third test segment plotted against the predicted values. The 

agreement is good. Also shown is the predicted amount of 

added helium as a percentage of the initial helium mass. The 

quantity of added helium was not measured in the tests. The 

amount added was small, about 0.7-percent of the initial mass 

was required for each TVS cycle. 

Figure 50.—Comparison of model predictions with 

data at 90-percent fill with temperature control. 

Figure 51.—Comparison of model predictions with 

data at 90-percent fill with pressure control. 

Figure 52.—Comparison of model predictions with 

data at 50-percent fill with temperature control. 
NASA/TP—2014-216633 35 

Figure 53.—Comparison of model predictions with 

data at 50-percent fill with pressure control. 

Figure 54.—Comparison of model predictions with data

at 50-percent fill with systematic helium addition. 

14.0 Conclusions 

An active thermodynamic vent system (TVS) was 

successfully demonstrated in a liquid oxygen tank pressurized 

with helium. Tests were conducted at nominal fill levels of 

90 and 50 percent. The liquid was heated above its normal 

boiling point to represent thermodynamic conditions in a

candidate lunar lander ascent stage. Three different hardware 

configurations for mixing the cooled liquid with the bulk fluid 

were tested: a submerged axial jet, a pair of spray hoops in the 

ullage, and combined operation of both. The combined mode 

maintained temperature throughout the entire tank, whereas 

the axial jet maintained liquid temperatures but did not control 

ullage temperatures. The spray-hoops-only mode was generally able to maintain liquid temperatures, although one test set 

was not successful. The liquid mixing may have been assisted 

by free convection driven by the high heat leak into the test 

tank. An initial test without helium pressurization verified the 

expected TVS behavior for the pure liquid and vapor condition 

in which the liquid temperature was controlled when the TVS 

was used to maintain tank pressure within a specified band. 

The presence of helium in the ullage decoupled the tank 

pressure from the liquid temperature. The experimental results 

showed that if the tank pressure was kept within a control 

band, the liquid temperature rose with each successive TVS 

cycle. Similarly, if the temperature was kept within a control 

band, the tank pressure would drop with each successive TVS 

cycle. A simple thermodynamic model was developed that 

predicts these trends. Correlation of modeling results shows 

that the amount of temperature rise or pressure drop per cycle 

was proportional to both the helium mole fraction and the heat 

input into the tank. The tank fill level was also important 

because the amount of temperature rise or pressure drop per 

TVS cycle increased as the ullage volume decreased. 

An additional test with liquid nitrogen pressurized with 

helium was conducted to demonstrate that simultaneous control 

of pressure and liquid temperature was possible by systematic 

introduction of a small amount of helium during each TVS 

cycle. The added helium occupied the volume of liquid 

removed from the tank during TVS venting cycles and existed 

as part of a binary ullage mixture along with propellant vapor. 

Glenn Research Center 

National Aeronautics and Space Administration 

Cleveland, Ohio, April 24, 2014 

NASA/TP—2014-216633 37 

Appendix—Symbols 

C flowmeter coefficient 

CntrlFlg flag for user-specified choice of temperature- or 

pressure-control logic 

DPDT user-specified tank pressure decay rate in control 

logic 

F liquid fill fraction 

GHe gaseous helium 

h enthalpy 

id propellant identification (e.g., “oxygen” or 

“methane”) 

LCH4 liquid methane 

LH2 liquid hydrogen 

LN2 liquid nitrogen 

LO2 liquid oxygen 

m mass 

m averaged mass 

mmass flowrate 

maveraged mass flowrate 

N2 nitrogen 

O2 oxygen 

P pressure 

Q heat/energy transfer 

Qheat input rate 

Qaveraged heat input rate 

R individual load cell (weighing system) reading 

T temperature 

t time 

t averaged time 

U total internal energy or uncertainty 

u specific internal energy 

V volume

x fluid quality (mass fraction of vapor) 

 first coefficient in weighing system calibration 

equation 

 second coefficient in weighing system calibration 

equation 

 third coefficient in weighing system calibration 

equation 

 constant in smoothing function for rate of change 

of pressure 

 efficiency 

 duty cycle 

density 

 tank turnover time 

 energy derivative defined in Equation (33) 

 volumetric flowrate 

Subscripts 

cal flowmeter calibration 

cp critical point 

cycle single thermodynamic vent system cycle 

e environmental 

el0 elapsed time control logic for checking maximum 

temperature or pressure

el1 elapsed time control logic for checking rate of 

pressure decay

el2 elapsed time control logic for checking minimum 

temperature or pressure 

ex exit 

f saturated liquid

fin final 

fl fluid 

fluid flowmeter fluid conversion 

fm flowmeter 

g saturated vapor 

He helium 

HX heat exchanger 

in inlet of heat exchanger or Joule-Thomson device 

int initial 
NASA/TP—2014-216633 38 

j load cell identification number (1 to 3) 

l liquid 

lc load cell 

min minimum 

max maximum 

new value at current time step during data acquisition 

o out (via venting or boiloff) 

old value from previous time step during data 

acquisition 

out outlet of heat exchanger 

p pump 

ref reference 

sat saturated 

TVS thermodynamic vent system 

tp triple point 

u ullage 

v vapor 

vent thermodynamic vent system vent 

1 initial state 

2 final state 

Superscripts 

* dimensionless quantity in correlation 

† dimensionless quantity in correlation 

‡ derived from mass balance 

 derived from energy balance 

Acronyms 

J–T Joule-Thomson 

MLI multilayer insulation 

TVS thermodynamic vent system 
NASA/TP—2014-216633 39 

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