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i

NASA/TP—2012–217469

Solar Thermal Upper Stage Liquid 

Hydrogen Pressure Control Testing 

and Analytical Modeling

A.D. Olsen, E.C. Cady, D.S. Jenkins, F.O. Chandler, G.D. Grayson, and A. Lopez

The Boeing Company, Huntington Beach, California

L.J. Hastings

Alpha Technology, Inc., Huntsville, Alabama

R.H. Flachbart and K.W. Pedersen

NASA Marshall Space Flight Center, Huntsville, Alabama

October 2012

National Aeronautics and

Space Administration

Marshall Space Flight Center • Huntsville, Alabama 35812
ii

Available from:

NASA Center for AeroSpace Information

7115 Standard Drive

Hanover, MD 21076–1320

443–757–5802

This report is also available in electronic form at

<https://www2.sti.nasa.gov/login/wt/>

Acknowledgments

The authors would like to thank the entire Marshall Space Flight Center Test Stand 300 test team for ‘a well done test 

program,’ in spite of significant programmatic and technical challenges. Special thanks to Leslie Curtis (NASA 

Marshall Space Flight Center Program Manager) for her unwavering support, to Joe Gaby and Dave Chato 

(NASA Glenn Research Center), and Jim LeBar, Joe Penitsch, and Mike Stoia (The Boeing Company). The 

timely assistance of Diana LaChance in the final documentation process also deserves special thanks.

An ‘acknowledgment above all others’ is offered to the family of Doug Richards in memory of his technical 

excellence and dedication to this program, to honor who he was as a person. 

TRADEMARKS

Trade names and trademarks are used in this report for identification only. This usage does not constitute an official 

endorsement, either expressed or implied, by the National Aeronautics and Space Administration.
iii

TABLE OF CONTENTS

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

2. PROJECT OBJECTIVES AND TEST ARTICLE DESCRIPTION .................................. 3

2.1 Objectives ...................................................................................................................... 3

2.2 Test Article and Fluid Systems Integration ................................................................... 3

2.3 Test Article/Facility Integration and Implementation ................................................... 8

3. ENGINEERING CHECKOUT TEST ............................................................................... 15

3.1 Approach ...................................................................................................................... 15

3.2 Engineering Test Results ............................................................................................... 16

4. ANALYTICAL CORRELATION OF ENGINEERING CHECKOUT 

TEST DATA ....................................................................................................................... 21

4.1 Standard Modeling ....................................................................................................... 21

4.2 Preliminary Computational Fluid Dynamics Analyses ................................................. 24

4.3 Computational Fluid Dynamics Detailed Modeling ..................................................... 26

5. THIRTY-DAY MISSION SIMULATION ......................................................................... 36

5.1 Mission Simulation Approach ...................................................................................... 36

5.2 Mission Simulation Testing and Evaluation Results ..................................................... 36

6. SUMMARY AND CONCLUSIONS ................................................................................. 44

6.1 Technical Strategy and Test Hardware Performance ..................................................... 44

6.2 Standard Modeling Test Correlations ........................................................................... 44

6.3 Engineering Test Results and Computational Fluid Dynamics Modeling ..................... 45

6.4 Standard Plus Computational Fluid Dynamics Modeling Combination ...................... 45

APPENDIX A—STUSTD INSTRUMENTATION LIST ...................................................... 47

APPENDIX B—AEROSPACE-INDUSTRY TECHNOLOGY PROJECT 

SUPPORTING PERSONNEL .................................................................... 52

APPENDIX C—STUSTD LIQUID HYDROGEN SPACE FLIGHT SYSTEMS

TEST REQUEST SHEET (1/25/99) ............................................................. 54

APPENDIX D—STUSTD 30-DAY MISSION SIMULATION (3/13/99) .............................. 66

REFERENCES ....................................................................................................................... 71
iv

LIST OF FIGURES

 1. Solar thermal orbital transfer vehicle concept ........................................................... 1

 2. STUSTD test article configuration ............................................................................ 4

 3. Top view of STUSTD tank internal components on tank bottom ............................ 4

 4. Test article installation in vacuum chamber ............................................................... 5

 5. Test article instrumentation and integration with facility fluid systems: 

TS300 vacuum chamber ............................................................................................ 6

 6. Test article before MLI installation ........................................................................... 7

 7. Test article during MLI installation ........................................................................... 7

 8. TS300 vacuum chamber ............................................................................................ 8

 9. Facility LH2 fill and drain ......................................................................................... 9

10. Facility LH2 vent tie-in to back-pressure control ....................................................... 10

11. Facility ice bath for back-pressure control ................................................................. 11

12. Control room for TS300 vacuum chamber ................................................................ 13

13. Vacuum chamber pressure during boiloff test ........................................................... 15

14. Tank vented flow rate during stabilization period ...................................................... 16

15. Example of LH2 destratification effects on tank pressure .......................................... 17

16. Example of ullage temperature destratification ......................................................... 17

17. Active TVS ................................................................................................................ 18

18. Passive TVS ............................................................................................................... 19

19. Successful extraction of tank heat by PTVS .............................................................. 19

20. Tank pressure drop due to ullage condensation from PTVS ...................................... 20
v

LIST OF FIGURES (Continued)

21. Final MLI installation ............................................................................................... 22

22. Predicted tank self-pressurization comparison with measured data,

mixer on, and 40-W heater input ............................................................................... 23

23. PTVS energy extraction versus exposed LAD length ................................................. 24

24. Velocity contours illustrating ATVS mixer pump performance at a 95% fill level ...... 25

25. Velocity contours illustrating ATVS mixer pump performance at a 40% fill level ...... 25

26. Fluid iso-surface illustrating lack of ullage penetration at a 30% fill level .................. 25

27. STUSTD model mesh and tank geometry ................................................................. 26

28. CFD modeled versus measured tank pressure during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer) ..................................................... 28

29. CFD modeled versus measured ullage temperature during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer) ..................................................... 28

30. CFD modeled versus measured liquid temperature at 64 cm (25 in) during 

self-pressurization—87% fill, 25-W tank heat leak (case 1, no mixer) ........................ 29

31. CFD modeled velocity and temperature patterns during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer) ..................................................... 29

32. CFD modeled versus measured tank pressure during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer) ..................................................... 30

33. CFD modeled versus measured ullage temperature during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer) ..................................................... 30

34. CFD modeled versus measured liquid temperature during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer (40 cm/15 in)) ................................ 30

35. CFD modeled velocity and temperature patterns during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer) ..................................................... 31

36. CFD modeled velocity and temperature patterns during mixing—90% fill, 

6.7-W tank heat leak (case 3, with mixer) .................................................................. 32
vi

LIST OF FIGURES (Continued)

37. CFD model versus measured tank pressure during self-pressurization 

cycle—90% fill, 6.7-W tank heat leak (case 3, 102 lpm (27 gpm) mixer) ..................... 33

38. CFD model versus measured tank ullage temperature during self-pressurization 

and mixing cycle—90% fill, 6.7-W tank heat leak (case 3, 102 lpm 

(27 gpm) mixer).......................................................................................................... 33

39. CFD model versus measured liquid temperature at (76.3 cm/30 in) during 

self-pressurization and mixing cycle—90% fill, 6.7-W tank heat leak (case 3, 

102 lpm (27 gpm) mixer ............................................................................................. 34

40. Measured tank and vacuum chamber pressures between ground fill 

and on-orbit simulation ............................................................................................. 37

41. Measured LH2 tank and vacuum chamber pressure versus shuttle ascent profile ..... 38

42. Predicted LH2 tank pressure as a function of MET .................................................. 39

43. Measured LH2 tank pressure before and after burn as a function 

of burn number ......................................................................................................... 39

44. Average chamber pressure as a function of burn number .......................................... 40

45. Maximum ullage temperature as a function of burn number ..................................... 41

46. Tank pressure and ullage vent flow rate during ullage vent ........................................ 41

47. Ullage temperature during ullage vent ....................................................................... 42

48. Average feed line flow rate as a function of burn number .......................................... 43
vii

LIST OF TABLES

1. Burn and coast times during the 30-day mission .......................................................... 21

2. STUSTD tank thermal performance ........................................................................... 22

3. STUSTD normal gravity test cases .............................................................................. 27

4. STUSTD CFD model results summary ....................................................................... 35

5. Comparison of burn pressure drop data with predictions ............................................ 40
viii
ix

LIST OF ACRONYMS AND SYMBOLS

AITP aerospace-industry technology project

ATVS active thermodynamic vent system

CFD computational fluid dynamics

CFM cryogenic fluid management

GH2 gaseous hydrogen

GHe gaseous helium

GN2 gaseous nitrogen

H2 hydrogen

LAD liquid acquisition device

LH2 liquid hydrogen

LHSFS liquid hydrogen space flight systems

LN2 liquid nitrogen

MLI multilayer insulation

MSFC Marshall Space Flight Center

PLC programmable logic controller

PTVS passive thermodynamic vent system

RNG renormalization group

ROV remote-operated valve

STUSTD Solar Thermal Upper Stage Technology Demonstrator (program)

TRS Test Request Sheet

TS test stand

TVS thermodynamic vent system

VFD variable-frequency drive 

VJ vacuum-jacketed
x
xi

NOMENCLATURE

Cv valve flow coefficient (dimensionless)

dp delta pressure (kPa)

dt delta time (s)

Tc innermost MLI temperature (coldest layer)

Th external MLI temperature (warmest layer)

a	 ratio of tank self-pressurization rise rate of interest (measured or calculated)

divided by self-pressurization rate for perfectly mixed or homogeneous tank

contents (dimensionless)
xii
1

TECHNICAL PUBLICATION

SOLAR THERMAL UPPER STAGE LIQUID HYDROGEN PRESSURE

CONTROL TESTING AND ANALYTICAL MODELING

1. INTRODUCTION

High-energy cryogenic propellant is an essential element in future manned space exploration programs, particularly those leading to manned missions to Mars. Therefore, NASA and 

its industrial partners are committed to an advanced development/technology program that will 

broaden the experience base for the entire cryogenic fluid management (CFM) community. The 

high cost of microgravity experiments has motivated the Marshall Space Flight Center (MSFC)/

Boeing team to aggressively explore combinations of ground testing and analytical modeling to the 

greatest extent possible. An aerospace-industry technology project (AITP), designed to demonstrate technology readiness for a representative solar orbital transfer vehicle (SOTV) concept 

(fig. 1), was cooperatively performed by a Boeing-led team with representatives from industry, academia, and government.1 A key element of the Solar Thermal Upper Stage Technology Demonstrator 

(STUSTD) program was the development of a liquid hydrogen (LH2) storage and supply system that 

used all of the vented hydrogen (H2) for solar engine thrusting, thereby avoiding boiloff losses. The 

strategy was to balance the LH2 storage tank pressure control requirements with the engine thrusting 

timeline during a 30-day solar thermal propulsion mission. The concept for achieving this strategy 

was to employ an arrangement of thermodynamic venting combined with a capillary screen channel 

liquid acquisition system, multilayer insulation (MLI), and internal tank heaters to reliably provide 

H2, at a controlled rate, to a solar thermal engine in the low-gravity environment of space operation. The STUSTD test data evaluation was initially performed with standard analytical modeling 

tools and later with advanced computational fluid dynamics (CFD) modeling. The test hardware, 

test results, and analytical correlations using both the standard and CFD modeling techniques are 

discussed herein. The analytical modeling correlations are considered a significant element in support of future space exploration efforts. As a matter of convenience, the program and engineering 

checkout test results are summarized in two previous papers; the documentation presented herein is 

much more complete and up to date.2,3
2

Figure 1. Solar thermal orbital transfer vehicle concept.
3

2. PROJECT OBJECTIVES AND TEST ARTICLE DESCRIPTION

2.1 Objectives

The test project objectives were to demonstrate the reliable and efficient storage of cryogenic 

hydrogen both during launch (transport and low Earth orbit) and the low-to-geostationary Earth 

orbit transfer mission simulation during which the solar engine operates. Specific test objectives were 

as follows:

(1) Demonstrate MLI purge evacuation and performance.

(2) Deliver near-saturated H2 vapor at 158–241 kPa (23–35 psia) at 0.91 kg/hr (2 lbm/hr) 

for 140 cycles over 30 days.

(3) Demonstrate venting/lockup for 140 burns over 30 days (without additional overboard 

venting required).

(4) Demonstrate use of MLI and heaters to assist control of tank pressure at 310±27.6 Kpa 

(45±4 psia) while operating.

(5) Demonstrate subcooled liquid acquisition device (LAD)/thermodynamic vent system 

(TVS) operation in 1 g via determination of internal LH2 tank temperature profile.

(6) Demonstrate LH2 loading/unloading and LAD filling for ground operations.

The test article and its integration into facilities at the MSFC Test Stand 300 (TS300)

vacuum chamber are described in the following sections.

2.2 Test Article and Fluid Systems Integration

An overview of the STUSTD test article configuration and a top view picture of tank internal 

components are presented in figures 2 and 3, respectively. A picture of the test article installation in 

the 6.1-m- (20-ft-) diameter vacuum chamber at MSFC TS300 is shown in figure 4.
4

Test Stand

Endcap

With Utility

Feedthroughs

LAD Channels (4)

With PTVS

Spaced 90°

Low-Conduction

Supports

ATVS

Temperature

Diode Rake

Figure 2. STUSTD test article configuration.

LAD

Channels

With PTVS

Temperature

Diode Rake

ATVS

Pump/Mixer

Figure 3. Top view of STUSTD tank internal components on tank bottom.
5

 Figure 4. Test article installation in vacuum chamber.

Figure 5 presents a schematic of the instrumented test article interfaces with the corresponding fluid and vacuum chamber facility elements at TS300. The NASA titanium tank has a volume of 

2 m3 (71 ft3) with 0.707 elliptical heads, a height of 124 cm (49 in), a width of 176 cm (69.3 in), and 

a LH2 capacity of 129 kg (284.1 lbm) with a 98% fill level. Twenty silicon diodes were placed in a vertical array to map the bulk fluid and ullage temperatures, with an additional two diodes on one LAD 

and five diodes placed throughout the MLI. Outflow and ullage pressures were measured along with 

volumetric flowmeter data. A total of 168 data channels were recorded at selectable data rates of 1, 

0.1, and 0.0167 Hz. Further details on both the test article and facility instrumentation are listed in 

appendix A.
6

Figure 5. Test article instrumentation and integration with facility fluid systems: 

TS300 vacuum chamber.

The MLI was double-aluminized Kapton® (DuPont) with alternating layers of Dacron® 

(DuPont) mesh (B4A). Because of problems during the fabrication of the MLI, the north and south 

poles had 100-layer pairs and the remaining tank surfaces had 75-layer pairs of MLI. Test article 

pictures before and during MLI installation are presented in figures 6 and 7, respectively. Further 

details regarding the MLI design, installation, and analytical modeling are presented in section 4.1.1.
7

Figure 6. Test article before MLI installation.

Figure 7. Test article during MLI installation.
8

2.3 Test Article/Facility Integration and Implementation

The MSFC test crew and their areas of responsibility at TS300 are listed in appendix B. The 

integrated hardware utilization and operations approach used by the MSFC/Boeing team are listed 

and defined in this section, beginning with vacuum chamber operations (fig. 8).

Figure 8. TS300 vacuum chamber.

2.3.1 Ascent to Vacuum

The shuttle ascent pressure profile was simulated as follows. The single stage gaseous nitrogen (GN2) ejector evacuated the chamber per the required profile to a chamber pressure of 40 torr, 

whereupon the chamber’s roughing pumps were engaged and evacuated to <0.1 torr (≈1 hr). Then 

diffusion pumps reduced the chamber pressure to 10×10–5 torr or less within about an hour. The 

vacuum system was arranged to go into a lockup condition in the event of a power failure. The 

vacuum facility instrumentation, listed in appendix B, included the following:

(1) Thermal: Thermocouples were attached to the cold walls and used to map the vacuum 

chamber’s thermal environment. The test engineer and test requestor determined the distribution of 

≈25 thermocouples based on previous test data. The LN2 cold wall system enters a vent condition 

upon power failure.

(2) Dewpoint: A dewpoint meter monitored the chamber gas dewpoint.

(3) Vacuum level: Eight instruments assured coverage from 760 to 10–8 torr. 
9

2.3.2 Facility Liquid Hydrogen Fill and Drain

As shown in figure 9, the LH2 was supplied by trailer and transferred to the test article tank 

via existing vacuum-jacketed (VJ) facility piping. The supply line to the vacuum chamber wall was 

3.8-cm-diameter (1.5-in-diameter) tubing, approximately 18.3 m (60 ft) long. The flow was controlled 

by a throttle valve, overpressure prevented by a relief valve set at 480.6 kPa (69.7 psia), and supply 

temperature provided by a thermocouple located in the flow path near the vacuum chamber wall. 

Test article draining was accommodated through a portion of the fill piping, then out to an existing 

burn stack. 

Chamber Penetration

Fill/Drain, Vent/Press

Fill/Drain Tie-in to

Facility Piping

Fill/Drain Tie-in to

Facility Piping

Figure 9. Facility LH2 fill and drain.

2.3.3 Test Article Piping System

Vacuum-jacketed piping connected the existing facility piping with the test article via facility piping that penetrated the vacuum chamber wall. The inner fill line was wrapped in MLI, and 

the vacuum jacket was connected to a vacuum pump and evacuated to <1 torr before LH2 transfer. 

Foam insulation was applied to the tubing outside the chamber (between the test article and facility 

VJ piping). The test article VJ piping did not extend completely to the test article tank flange. Therefore, there was approximately 0.61 m (2 ft) of bare tubing that required gaseous helium- (GHe-) 

purged MLI. The 1.6-cm (5/8-in) electrically actuated fill/drain valve (supplied by Boeing) resided 

inside the VJ pipe, accessible through a 15.2-cm (6-in) conflat-style port. The outside of the vacuum 

jacket had a bellows, and the internal lines had either a flex hose or expansion/contraction loops to 

account for thermal contraction. Outside the vacuum chamber, the vacuum jacket had a vacuum 

relief valve set at 115.1 kPa (16.7 psig).
10

2.3.4 Facility Liquid Hydrogen Vent Tie-In to Back-Pressure Control 

As shown in figure 10, existing facility piping was used which consisted of 5-cm (2-in) tubing 

that tied off to a high-flow, chill-down leg and to the back-pressure control, flow measurement leg. 

The back-pressure control subsystem was used to maintain the tank ullage pressure at the required 

steady state conditions during the boiloff or heat leak measurement phase of testing. The system was 

composed of four flow control valves with flow coefficients (Cv

s) ranging from 25 to 0.01 that were 

located in the vent line. Each control valve was regulated through a closed-loop control system. This 

control loop manipulated the valve positions based on a comparison of the measured tank ullage 

pressure with the desired set point. An MKS Instruments, Inc., Baratron 0–133 kPa (0–19 psia) absolute pressure transducer (accuracy of ±0.02%) and an MKS delta pressure (dp) transducer (1 torr or 

133 Pa head with an accuracy of ±0.04%) located outside the vacuum chamber were used to measure 

ullage pressure. The 1-torr differential transducer had a tank submersed in an ice bath (fig. 11) as a 

pressure reference source. This tank was charged to 0.5 torr below the desired test article tank ullage 

pressure before the test start. Downstream of the four control valves, the flow converged to supply 

five flowmeters, with flow ranges from zero to 4,250 l/m. These were connected in series with a selector valve between each. The arrangement enabled ullage pressure control to within ±0.00689 kPa 

(±0.001 psi).

Burst Disk Placed in Series

With This Relief Valve

Vent Tie-in to 

Facility Piping

Figure 10. Facility LH2 vent tie-in to back-pressure control.
11

 

Vent Heat

Exchanger

Flowmeter

Enclosure

Back-Pressure

Control Valve

Ullage Pressure

Measurement System

Figure 11. Facility ice bath for back-pressure control.

2.3.5 Test Article Piping

The test article piping resided inside the aforementioned VJ pipe and tied into the facility 

piping, just outside the vacuum chamber via a 11.4-cm- (4.5-in-) diameter conflat-style flange. It was 

1.9 cm (0.75 in) in diameter at the test article tank and expanded to 2.54 cm (1 in) in diameter once 

inside the VJ pipe. 

2.3.6 Overpressure Protection

The Boeing-supplied burst disk was connected to the vent piping, and a relief valve was 

installed in series with this burst disk for two reasons: (1) In the event the burst disk ruptured, the 

relief valve allowed self-pressurization for tank draining and (2) it better ensured that all boiloff 

gaseous hydrogen (GH2) was measured by the flowmeters. All tank ullage and liquid isolations were 

double isolated with a leak-check port between the isolations. This allowed verification that no gas 

was leaking past the first isolation after tanking LH2.
12

2.3.7 Test Article Pressurization

An existing pressurization system that used either a trailer or facility-supplied H2 as its source 

was used. The pressure was regulated and flow controlled by a control valve. This gas tied into the 

vent piping, downstream of the back-pressure control valves.

2.3.8 Test Article Conditioning

Test article tank conditioning was accomplished as follows. A dry nitrogen purge (facility 

nitrogen) was connected to trickle through and maintain a dry test article as soon as it was installed 

in the chamber. Before tanking liquid nitrogen (LN2) or LH2, the vacuum chamber was evacuated 

to <1 torr. Then the test article tank and its associated piping were evacuated to <0.1 torr and backfilled with GN2 (if tanking LN2) or GH2 or GHe (if tanking LH2) to 138 kPa (20±1 psia). This was 

repeated three times. After the last cycle, when the test article tank was again at 138 kPa (20±1 psia), 

the vacuum chamber could be repressurized if required. Just before tanking, the tank and piping 

were subjected to five pressure/vent cycles, pressurizing to 204.8 kPa (29.7 psia) and venting to zero 

pressure using GN2 (if tanking LN2) or GH2 (if tanking LH2).

2.3.9 Simulated Engine Vent Heat Exchanger

The heat exchanger lines were leak tested with a mass spectrometer and 69 kPa (10 psig) GHe 

before being installed into the vacuum chamber. The fluid lines were all welded except for the two 

0.32-cm (1/8-in) Swagelok® fittings for the thermocouples. The fluid lines were proof tested, mass 

spectrometer leak tested, and cold shocked per standard procedures.

2.3.10 Simulated Engine Valve Panel

The Boeing-supplied valve panel was located on the top level of the TS. The panel was delivered with an enclosure purged with GN2. The line that connected this panel with the heat exchanger 

inside the vacuum chamber was required to be 1.27 cm (0.5 in) in diameter. Downstream of the valve 

panel was a water-jacketed heat exchanger designed to warm the vent gas to ground water temperature before reaching the simulated engine orifice.

2.3.11 Test Article Liquid Acquisition Device Vent Line

This line was routed inside the test article VJ pipe. Initially, it was a quarter-inch in diameter 

as it exited the test article tank and expanded to a half-inch in diameter before entering the VJ pipe. 

A cryogenic-rated bellows seal valve served as the isolation for this tube, instead of the supplied 

Marotta solenoid valve. This valve was normally closed with limit switches to indicate valve position.

2.3.12 Multilayer Insulation Purge

This system was supplied by the existing facility pressurization system. Up to 446.1 kPa 

(64.7 psia) GHe was delivered to a metering hand valve, flowmeter, pressure transducer, and vacuumrated isolation hand valve. The requirement for this system was to supply GHe to create 102 kPa 
13

(14.8 psia) of pressure under the MLI. If, during checkouts, the required flow deviated substantially from this, it was recognized that a modification of the system could be required. Installation 

of a relief device was not expected. Tubing with a 1.27-cm (0.5-in) diameter was used between the 

metering valve and test article tie-in. The pressure transducer was positioned on this tubing just outside the vacuum chamber.

2.3.13 Test Article Instrumentation

All test article instrumentation, except for PT4, was provided by Boeing and is listed in appendix A. The MSFC test support personnel mated the test article instrumentation cabling to the facility 

instrumentation junction boxes and verified operation of test article instrumentation. All test article 

instrumentation was National Institute of Standards Technology traceable and included the applicable calibration certifications. All test article instrumentation signals were recorded on MSFC’s lowspeed digital data acquisition system. Test Division supplied all instrumentation signal conditioners. 

Further details on both the test article and facility instrumentation are listed in appendix A.

2.3.14 Automated Controls

The control room for TS300, located in Building 4650 and shown in figure 12, supported two 

main operations.

Figure 12. Control room for TS300 vacuum chamber.

2.3.14.1 Valve, Heater, and Camera Power Control. Facility remote-operated valves (ROVs), 

test article ROVs, facility variable-position valves, test article feed line heat exchanger heaters, test 

article heaters, and the facility vacuum chamber video camera system were controlled by facility 
14

programmable logic controllers (PLCs). The 30-day mission simulation logic was programmed in the 

facility PLC using ladder logic software.

2.3.14.2 Active Thermodynamic Vent System Pump Variable-Frequency Drive Controller.

Depending on the capability of the supplied active thermodynamic vent system (ATVS) pump 

variable-frequency drive (VFD), pump speed was to be set by sending a remote analog signal from 

a PLC or individual analog controller to the VFD. The VFD itself was mounted at the test stand. 

It should be noted, however, that the vendor-supplied controller failed to function with LH2 and 

an MSFC-supplied controller was substituted. The pump vendor, lacking LH2 test facilities, had 

checked out the pump system in LN2. However, the controller start-up characteristics were altered 

considerably by the reduction in electrical resistance that occurred during operation at 20 K, compared with 77 K with nitrogen. The MSFC controller, qualified to operate at LH2 temperatures, was 

substituted and the pump start-up was successful.

2.3.15 Special Test Equipment Fabrication and Integration

2.3.15.1 Test Article Vacuum Jacket Piping. The test article VJ piping was designed, fabricated, and installed by MSFC personnel. It was proof tested, mass spectrometer leak tested, and 

cold shocked per standard MSFC specifications. It consisted of the fill/drain, vent/pressurization, 

and LAD vent tubes residing inside a 15.2-cm (6-in) schedule-10 pipe jacket.

2.3.15.2 Simulated Engine Vent Water Heat Exchange. The simulated engine vent water heat 

exchanger consisted of 1-in-diameter stainless steel tube residing inside a 10.2-cm (4-in) PVC jacket. 

Water flowed through the jacket to warm the gas inside the tube. The length was approximately 

2.4–3.05 m (8–10 ft).

For further details regarding the test article receipt and inspection, test article integration 

into the test facility, systems installation, checkout/verification, and test operations, refer to the Test 

Implementation Plan.4
15

3. ENGINEERING CHECKOUT TEST

3.1 Approach

The engineering tests consisted of four primary phases: (1) Tank LH2 fill and stabilization, 

(2) boiloff measurement to determine the system heat leak, (3) measurement of the tank selfpressurization and heated pressurization rates, and (4) measurements of the passive and active TVS 

thermal extraction capabilities during several simulated ‘burn’ periods (outflow from the tank). 

Details regarding test conditions, control limits, instrumentation versus sampling rates/intervals, 

etc., are defined in the Test Request Sheet (TRS) shown in appendix C.

During the entire engineering test, the MSFC TS300 LN2 cold walls were used to simulate 

the thermal environment of space. After filling the tank with 129 kg (284 lbm) LH2, the tank pressure was held constant at 116 kPa (16.75 psia) in anticipation of a boiloff test. Once the engineering 

tests began, however, a leak was indicated by an increasing chamber pressure (fig. 13). A residual gas 

analyzer verified that H2 was leaking into the vacuum chamber from the test article.

1.5 × 10–5

1.4 × 10–5

1.3 × 10–5

1.2 × 10–5

1.1 × 10–5

1.0 × 10–5

9 × 10–5

8 × 10–5

7 × 10–5

6 × 10–5

0 50,000 100,000 150,000 200,000

Boiloff Test Time (s)

Vacuum Chamber Pressure (torr)

Figure 13. Vacuum chamber pressure during boiloff test.

Subsequently, the LH2 tank was allowed to self-pressurize to an operating pressure of 

311 kPa (45 psia) and the vacuum chamber pressure finally topped out at approximately 5.5×10–5

torr, compared with the initial value of 6.5×10–6 torr. Because the vacuum pressure appeared to be 

stabilized, a decision was made by Boeing and MSFC to continue the test with the higher chamber 

pressure and to account for the increased MLI heat leak in the analytical modeling.
16

3.2 Engineering Test Results

After the vacuum chamber and boiloff achieved a steady state (fig. 14), 10 hr of stable test 

data were obtained. The equilibrium boiloff vent rate was 10.8 l/min (660 in3/min), which corresponded to a system heat leak of 6.7 W.

10,000

9,000

8,000

7,000

6,000

5,000

5,000

3,000

2,000

1,000

0

0 50,000

25

50

100

150

100,000 150,000 200,000

Test Time (s)

Boiloff Flow (scim)

Boiloff Flow (l/m)

Figure 14. Tank vented flow rate during stabilization period.

The system heat leak increased to ≈18 W during the ATVS testing due to the mixer pump 

operation combined with the elevated vacuum chamber pressure. The tank was locked up unmixed 

with 108.9 kg (240 lbm) of LH2 (LH2 fill level of about 94 cm (37 in)) for ≈9 hr without tank heaters being used. The pump was turned on from 12.98 to 13.22 hr and, as expected, mixing the tank 

contents lowered the tank pressure. Destratification in the liquid and ullage is illustrated in figures 15 

and 16, respectively. 
17

140 20

19.5

19

18.5

18

17.5

135

130

125

120

115

2.5 12.7 12.9 13.1 13.3 13.5

Tank Pressure (kPa)

Tank Pressure (psia)

Hours After Midnight 2/6/99

Figure 15. Example of LH2 destratification effects on tank pressure.

44

43

42

41

40

39

38

20 37

22

24

12.5 12.7 12.9 13.1

Hours After Midnight 2/6/99

Temperature (°R)

Temperature (K)

13.3 13.5

Liquid at 30 in (0.76 m)

Liquid at 15 in (0.38 m)

Ullage 41 in (1.04 m)

Ullage 44 in (1.12 m)

Figure 16. Example of ullage temperature destratification.

Once the full tank reached an operating pressure of 310 kPa (45 psia), the vent system flow 

tests began. The system was designed to operate at a mass flow rate of 0.9 Kg/hr (2 lbm/hr) with 

a pressure drop of 75.8 kPa (11 psi) (55.7 kPa (8 psi) from the Joule-Thomson devices, 13.8 kPa 

(2 psi) from the ATVS cold side line, and 6.9 kPa (1 psi) from the flow lines, heat exchanger plate, 

instrumentation, etc.). A mass flow rate of 1.18 kg/hr (2.6 lbm/hr) with a pressure drop of 10.35 kPa 

(1.5 psi) was measured. By reducing restrictions downstream of the valve panel, the system reached 

a mass flow rate of 2.3 kg/hr (5.1 lbm/hr) at a pressure drop of 69 kPa (10 psi). Because this mass 

flow rate was approximately twice the expected value for the same system pressure drop, it indicated 
18

an internal leak inside the tank into the flow path exiting the tank. A real-time pressure drop analysis indicated that the leak was on the downstream side of the ATVS, shown in figure 17, and was 

equivalent to an approximate 1-mm-diameter hole. Because the internal components depended on 

the appropriate pressure and temperature drop to extract energy from the tank contents and thereby 

control tank pressure, the test was continued with the internal leak and higher mass flow rate, because 

the expected pressure drop was now being measured. The ATVS and passive thermodynamic vent 

system (PTVS) required the 55.2 kPa (8 psi) pressure drop through the Joule-Thomson devices to 

control energy extraction and, therefore, tank pressure. 

Probable Leak

Location

Figure 17. Active TVS.

During tank outflow, the bulk fluid exited the LAD channel, passed through the Joule-Thomson device (which expanded and cooled the flow), and then passed through the PTVS tube brazed 

into the LAD apex, shown in figure 18. The PTVS removed energy from the tank by condensation 

and convection. This was an important effect when the LAD surface was exposed. Figure 19 shows 

the effect of the PTVS on the tank ullage temperature with 12.7 cm (5 in) of LAD exposed. Figure 

20 shows that the tank pressure reduction during the burn indicated a larger than expected energy 

extraction rate of 11.3 W (38.5 Btu/hr). Toward the end of the test, the tank bulk fluid level was 

reduced to 40.8 kg (90 lbm) or a fill level of 43.2 cm (17 in). Again, the outflow tests indicated a larger 

than expected PTVS energy extraction rate of 85.6 W (293 Btu/hr). The expected versus analytical 

energy extraction rates are discussed in section 4. 
19

JouleThomson

Device

Figure 18. Passive TVS.

56

55

54

53

52

51

50

49

48

47

46

45

26

27

28

29

30

31

50 150 350

Burn Time (s)

550 750

Temperature (K)

Temperature (°R)

Ullage 44 in (1.12 m)

Ullage 41 in (1.04 m)

LAD (diode submerged and insensitive)

Ullage 43 in (1.09 m)

Figure 19. Successful extraction of tank heat by PTVS.
20

–50

45.1

45

44.9

310

305

44.8

44.7

44.6

44.5

44.4

44.3

150 350

Burn Time (s)

550 750

Tank Pressure (psia)

Tank Pressure (kPa)

Figure 20. Tank pressure drop due to ullage condensation from PTVS.
21

4. ANALYTICAL CORRELATION OF ENGINEERING CHECKOUT TEST DATA

The STUSTD test data evaluation was first performed with standard analytical modeling 

tools and later with advanced CFD modeling. The advantages of standard analytical modeling 

techniques include: (1) a quick turnaround, (2) use on a wide variety of computers, and (3) have 

historically been the industry standard for day-to-day design studies and trades. CFD modeling techniques offer simulation of the actual physical phenomena and, therefore, can generally be applied to 

low-gravity conditions with more confidence than the standard modeling. It is firmly believed that 

with each CFD code improvement increment, the historic dependence of CFM development on 

costly on-orbit experimentation is being significantly reduced. Therefore, although CFD modeling is 

more costly than traditional modeling and requires more experienced personnel, the costs are miniscule compared with on-orbit testing. Hence, the general strategy is to anchor both traditional and 

CFD modeling techniques with available test data, but use the CFD to model selected conditions 

in reduced gravity. Then, the CFD modeling can be used to guide or anchor traditional modeling 

applications to reduced gravity. Application of both standard and CFD techniques to the evaluation 

of the subject data is presented in the following sections.

4.1 Standard Modeling

4.1.1 Multilayer Insulation Modeling and Test Correlation

The solar propulsion concept involves using the LH2 tank as an accumulator, with the tank 

pressure dropping during venting (thrusting) and rising during lockup (coasting). Each pressure 

excursion depends on the vent rate, the insulation design, and the mission duty cycle. Example engine 

thrust durations and coast periods for the 30-day mission are presented in table 1. The burn durations ranged from 0.72 hr to 2.29 hr on day 1 and day 30, respectively. Corresponding coast periods 

between burns ranged from 0.89 hr to 20.26 hr. The long coast periods near the end of the mission require that a high-performance MLI system be used to limit the pressure rise in the tank. 

The MLI blanket arrangement and its integration with propulsion system mission requirements are 

summarized here with further details in appendix D. 

Table 1. Burn and coast times during the 30-day mission.

Day No. Burn No. Burn Duration (hr) Coast Duration (hr) dp/dt, kPa/hr (psi/hr)

1 1 0.23 0.89 6.9 (1)

15 118 0.21 9.20 10 (1.45)

30 140 0.57 20.21 2 (0.29)
22

The MLI system was GHe purged, and the MLI perforated for evacuation efficiency during 

ascent. The MLI blankets were conformed around the tank penetrations and supports to restrict 

the heat leak. The MLI arrangement consisted of four blankets with 25 layers each at the poles, 

and three blankets with 25 layers each on the equator. Each blanket consisted of 1/4-mil aluminized 

Mylar® with Dacron® (DuPont) net spacers. The final MLI installation, shown in figure 21, resulted 

in a calculated thermal performance of 6.6 W, as shown in table 2. Boiloff testing indicated an actual 

heat leak of 6.7 W, which was within 1.5% of the prediction. 

4 Blankets at 

25 Layers Each

SpiralWrapped

Supports

Support

Penetration

Covers

Spiral-Wrapped

Penetrations

Tie-Down

Straps

3 Blankets at

25 Layers Each

Figure 21. Final MLI installation.

Table 2. STUSTD tank thermal performance.

4 blankets; 25 layer-pairs each DAK/B4A; Th=200 K, Tc=25 K

Performance (W)

Budget 4 Blankets Calculated 3 + Blankets Calculated

MLI blankets

MLI seams

MLI supports

MLI margin

Tank supports

1.6

0.5

0,3

–

2.0

1.68

–

–

0.82

2.00*

2.24

–

–

0.82

2.00*

Plumbing 1.0 1.31 1.31

Instrumentation wires 0.6 0.06 0.06

Heat leak margin – 0.23 0.17

Margin 6.0 6.00 6.60

*NASA Glenn Research Center Test Data
23

4.1.2 Tank Pressure Control

The tank self-pressurization rate modeled using the equation for homogeneous conditions 

from Lin et al.5 and a total heat leak of 47.7 W are compared to the measured data “with the mixer 

on” in figure 22. Assuming that α represents the “subject pressure rise rate divided by the theoretical 

rate corresponding to homogeneous or fully mixed conditions.” The predictions were based on α =1, 

whereas the actual α=1.15; i.e., the measured rate was 15% more than the rate with fully mixed conditions. This indicates good mixing of the contents with a small ullage. However, with the mixer off, 

the measured α=3.4 due to stratification within the bulk liquid.

45

43

41

300

250

200

39

37

35

33

31

29

27

0 2 4 6 8 10 12 14

Tank Pressure (psia)

Tank Pressure (kPa)

Hours After Midnight 2/7/99

Actual α = Measured/Homogeneous

 = 1.15 

 Tank Heat Leak = 6.7 W

 Mixer On = 1 W

 Heaters = 40 W

 Total Input = 47.7 W 

Predicted α = Analytical/Homogeneous

 = 1 

Figure 22. Predicted tank self-pressurization comparison with 

measured data, mixer on, and 40-W heater input.

The analytical PTVS energy extraction rate during the burn was 3.1 W (10.6 Btu/hr) compared with a measured rate of 11.3 W (38.5 Btu/hr). With the fluid level at 40.8 kg (90 lbm) (a fill level 

of 17 in), a PTVS energy extraction rate of 7.6 W (26 Btu/hr) was predicted, compared with a measured rate of 85.9 W (293 Btu/hr). The ullage was warmer than expected; a temperature corresponding to saturated conditions had been previously assumed. However, the stratified ullage temperatures 

ranged from 28 K to 31 K (50 °R to 55 °R), that were well above the saturation temperature of 25 K 

(45 °R). The warm ullage was attributed to the following: (1) The pump was unable to disturb the 

liquid-vapor interface sufficiently to achieve ullage cooling, and (2) the heat leak was almost three 

times greater than expected because the vacuum chamber pressure increased and the MLI performance was affected. However, the major reason for the heat extraction rate difference was the large 

temperature difference between the warm, stratified ullage and the LAD channel/PTVS, which was 

much greater than expected. The prediction was based on a perfectly mixed ullage at saturated conditions. Also, underpredictions of condensation heat transfer in the model (using the axial height 

instead of the true exposed LAD length) were left in the model for conservatism. Figure 23 shows 

the large PTVS energy extraction versus exposed LAD length when the LAD diode is exposed to the 

ullage during a burn. Consequently, a rapid drop in LAD temperature occurred whenever the PTVS 

flow was initiated. 
24

500

450

400

350

300

250

200

150

100

50

0

145

131

116

102

87

73

58

44

29

15

0

0 0.508 1.016 1.524

LAD Length Exposed to Ullage (m)

Stratified Ullage

Mixed Ullage

Heat Removal (Btu/hr)

Heat Removal (W)

Figure 23. PTVS energy extraction versus exposed LAD length.

4.2 Preliminary Computational Fluid Dynamics Analyses

FLOW-3D is a general Navier-Stokes equation solver with an extensive history of cryogenic 

tank modeling. FLOW-3D6 offers several options based on what is important to the problem, including application to reduced gravity environments. Consequently, a preliminary study of the TVS mixer 

was performed which indicated that the mixer pump was unable to produce a geyser that penetrated 

or disturbed the liquid-vapor interface sufficiently to achieve adequate mixing of both the ullage and 

bulk liquid. Referring to figures 24 and 25, the CFD analyses indicated minimal disturbances at the 

liquid-vapor interface with fill levels of 90% and 40%, respectively. The fluid iso-surface view, shown 

in figure 26, clearly conveys the very minimal disturbance of the liquid-vapor interface with the 30% 

fill level. It can be argued that the mixer performance will substantially improve in reduced gravity; 

however, that is not always the case. Further, the experimentally measured condensation effects can 

be a significant positive factor in reduced gravity. In conclusion, further testing and analyses would 

be required to resolve the mixer adequacy question for a particular design and mission. 
25

0.725

0.5452

0.3625

0.1827

0

Velocity (m/s)

Figure 24. Velocity contours illustrating ATVS mixer pump performance at a 95% fill level.

0.725

0.5452

0.3625

0.1827

0

Velocity (m/s)

Figure 25. Velocity contours illustrating ATVS mixer pump performance at a 40% fill level.

0.725

0.5452

0.3625

0.1827

0

Velocity (m/s)

Extent of Mixer

Penetration

Figure 26. Fluid iso-surface illustrating lack of ullage penetration at a 30% fill level.

The analytical evaluation of pressure rise rates, mixer performance, and energy extraction 

rates is further pursued in section 4.3. 
26

4.3 Computational Fluid Dynamics Detailed Modeling

As mentioned earlier, the STUSTD data were later evaluated in more detail using a custom version of the commercially available FLOW-3D software to develop the present two-phase 

cryogenic tank model.6,7 The customization enabled the model to treat phase-change effects at the 

liquid-gas interface. First-order approximations for momentum and energy equations, including the 

two-equation k-∈ and renormalization group (RNG) turbulence models, were enabled. The ullage 

region was fully compressible, and liquid density varied with temperature only. Modeling of the heat 

transfer between the liquid, gas, and tank walls was included to capture thermal stratification within 

the fluids. 

The model setup, analysis results, and conclusions are described in the following sections.

4.3.1 Computational Fluid Dynamics Model Description

The model used a 3,366-cell axisymmetric mesh to simulate the 2.01 m3 (71 ft3) ellipsoidal 

STUSTD tank (fig. 27). The mesh was derived from the S-IVB mesh previously shown to be independent of grid size,8 therefore, the STUSTD mesh was expected to have little grid-induced effect. 

The tank had a width of 1.76 m (5.78 ft) and a height of 1.24 m (4.08 ft). LADs were omitted from 

the model because the engineering checkout tests did not include LAD operation. The ATVS was 

located in the center of the tank and was 0.29 m (1 ft) in diameter and 0.19 m (0.63 ft) in height. The 

outlet of the vertical jet was 0.4 cm (0.14 ft or 1.7 in) in diameter.

Tank Wall

ATVS

Heat Exchanger

ATVS

Outlet

ATVS

Inlet

20 W/40 W

Tank Heaters

Figure 27. STUSTD model mesh and tank geometry.

A series of tank wall heaters capable of delivering 20–40 W of heating were imbedded in 

the tank insulation. Because of the axisymmetric mesh, the STUSTD model incorporated these 

variable-power tank heaters as solid obstacles adjacent to the bulk liquid. The actual tank heaters 

were approximately 15 cm (6 in) wide by 66 cm (26 in) long. Four of these tank heater strips were 

evenly distributed on the tank wall. 
27

The ATVS body was considered to be adiabatic, thus contributing a negligible amount of 

heat to the liquid. The self-pressurization models were initially quiescent. The liquid was assumed to 

be saturated at a given tank pressure and allowed to heat up. Ullage and liquid stratification profiles 

were derived from available test data and were used as ‘initial conditions’ on a case-by-case basis.

Table 3 lists the test cases considered in the present analysis. The focus was on normal gravity 

tank self-pressurization and ATVS performance, and the modeling was anchored by previous correlations with MSFC test data.9 Test case 3 included ATVS activation, whereas all the other cases strictly 

treated tank self-pressurization under external heating. The tank heat leak was distributed between 

the tank wall and the imbedded heater obstacles. When heaters were not used, the incoming tank 

heat leak was evenly distributed along the tank walls, including the surface of the inactive heaters. 

Table 3. STUSTD normal gravity test cases.

Test Case Test Fluid

Heat Leak

(W)

Fill Level

(%)

Mixer Type

and Flow Rate

1 LH2 25.7 87 None

2 LH2 25.7 44 None

3 LH2 25.7 90 Axial jet/27 gpm

4.3.2 Computational Fluid Dynamic Results and Discussion

Detailed model results for the normal gravity test cases are presented in figures 28–39. Transient ullage pressure histories, ullage and liquid temperatures, histories, liquid temperature histories, 

temperature contours, and velocity field plots for each case are displayed. Ullage temperatures were 

measured 1.68 m (44 in) from the bottom of the tank, whereas liquid temperatures were recorded in 

the bulk liquid. Because of the different fill levels, recorded liquid temperatures for some test cases 

were measured 0.64 or 0.38 m (25 or 15 in) from the bottom of the tank. The specific temperature 

sensor location is indicated on each plot. 

4.3.2.1 Self-Pressurization. In general, the normal gravity results closely followed the trends 

evident in the test data. Several results from selected cases without mixing are interpreted and discussed in detail in this section. 

In case 1 (25.7 W, 87% fill, no mixer), shown in figures 28–31, the tank heaters operated at half 

capacity (20 W) and 5.7 W remained evenly distributed along the tank walls. Uniform saturated conditions at 234 kPa, 23.5 K (34 psia, 42.3 °R) and a quiescent velocity field were assumed for the initial 

conditions. The large natural convection vortices generated in the bulk liquid by the tank heaters 

were dispersed and slow, on the order of 0.91 cm/s (0.03 ft/s). The results show that the tank heaters 

impacted the velocity field profiles and magnitudes. Thermal stratification of the ullage was high but 

relatively constant (≈5 K (9 °R)) throughout the self-pressurization phase, with a peak temperature 

of 29.8 K (53.7 °R). Model results modeled an average ullage pressurization rate of 6.9 kPa/hr(1 psi/

hr), which agreed with the measured 6.34 kPa/hr (0.92 psi/hr) rate. Modeled ullage and liquid temperatures at 1.68 and 0.64 m (44 and 25 in), respectively, also agreed well with the measured test data 

and were within 0.28 K (0.5 °R) of the actual sensor readings. 
28

38

37

260

250

240

36

35

34

120,000 125,000 130,000 135,000

Ullage Pressure (psia)

Ullage Pressure (kPa)

Test Data

Model

Time (s)

0.92 psi (6.35 kPa)/hr

1 psi (6.9 kPa)/hr

Figure 28. CFD modeled versus measured tank pressure during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer).

52

50

28.9

27.8

26.7

25.6

24.4

23.3

48

46

44

42

130,000 135,000

Time (s)

120,000 125,000

Ullage Temperature (R)

Temperature (K)

Test Data

Model

Figure 29. CFD modeled versus measured ullage temperature during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer).
29

45

44

25

24.4

23.9

23.3

22.7

22.2

43

42

41

40

120,000 125,000 130,000 135,000

Liquid Temperature (°R)

Liquid Temperature (K)

Test Data

Model

Time (s)

Figure 30. CFD modeled versus measured liquid temperature at 64 cm (25 in) during 

self-pressurization—87% fill, 25-W tank heat leak (case 1, no mixer).

T (°R) T (K) T (°R)

49.0

47.9

46.8

45.7

44.7

43.6

42.5

53.7

51.9

50.1

48.3

46.5

44.7

42.9

(a) t = 123,094 s, Vmax = 0.0311 ft/s (b) t = 130,174 s, Vmax = 0.0293 ft/s

(27.2)

(26.6)

(26)

(25.4)

(24.8)

(24.2)

(23.6)

T (K)

(29,8)

(28.8)

(27.8)

(26.8)

(25.8)

(24.8)

(23.8)

= 0.95 cm/s = 0.89 cm/s

Figure 31. CFD modeled velocity and temperature patterns during self-pressurization—

87% fill, 25.7-W tank heat leak (case 1, no mixer).

For case 2 (25.7 W, 44% fill), shown in figures 32–35, the tank heaters were operating at 20 W. 

The remaining 5.7 W were evenly distributed along the tank walls. The tank was initially quiescent 

and assumed to be saturated at 241 kPa (23.6 K) or 35 psia (42.5 °R). The ullage was stratified by 

4.7 K (8.5 °R) to simulate the temperature distribution in the ullage during the actual test. Both the 

model and test data illustrated a 10-kPa/hr (1.45-psi/hr) tank pressurization rate. Ullage temperature results illustrated a 1.12 K (2.5 °R) rise over the 13.8 kPa (2 psi) pressure band. Modeled liquid 

temperature results of 40 cm (15 in) from the bottom of the tank agreed with test data and showed 

a small rise in temperature. The temperature and velocity field plots illustrated similar bulk liquid 

fluid dynamics for case 2 as for case 1. Ullage thermal stratification was slightly higher at the lower 

fill level.
30

38

37

36

35

34

260

250

240

3,200 4,200

Time (s)

1,200 2,200

Ullage Pressure (psia)

Ullage Pressure (kPa)

Test Data

Model

1.45 psi (10 kPa)/hr

Figure 32. CFD modeled versus measured tank pressure during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer).

56

55

54

53

52

51

50

49

48

31.1

30

28.9

27.8

26.7

3,200 4,200

Time (s)

1,200 2,200

Ullage Temperature (°R)

Ullage Temperature (K)

Test Data

Model

Figure 33. CFD modeled versus measured ullage temperature during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer).

46

45

44

43

42

41

40

25.6

25

24.4

23.9

23.3

22.8

22.2

3,200 4,200

Time (s)

1,200 2,200

Liquid Temperature (°R)

Liquid Temperature (K)

Test Data

Model

Figure 34. CFD modeled versus measured liquid temperature during self-pressurization

—44% fill, 25.7-W tank heat leak (case 2, no mixer (40 cm/15 in).
31

T (°R) T (°R)

54.6

52.6

50.6

48.6

46.7

44.7

42.7

56.3

54.1

51.9

49.6

47.4

45.2

42.9

T (K)

(31.3)

(30.1)

(28.8)

(27.6)

(26.3)

(25.1)

(23.8)

(a) t = 1,837 s, Vmax = 0.0934 ft/s (b) t = 1,837 s, Vmax = 0.0840 ft/s

T (K)

(30.3)

(29.2)

(28.1)

(27)

(25.9)

(24.8)

(23.7)

= 2.85 cm/s = 2.58 cm/s

Figure 35. CFD modeled velocity and temperature patterns during self-pressurization—

44% fill, 25.7-W tank heat leak (case 2, no mixer).

4.3.2.2 Active Thermodynamic Vent System Activation. For test case 3 (6.7 W, 90% fill, 

102 lpm (27 gpm)), in which the ATVS was used, the temperature and velocity field plots for the selfpressurization, ATVS, and repressurization phases of the pressure control cycle are illustrated 

(fig. 36). Figure 36(a) illustrates the temperature and velocity distribution once the tank reached its 

upper pressure limit, hence, the end of the self-pressurization phase. Figure 36(b) displays the temperature and velocity field at the end of the TVS phase, when the ATVS destratified the ullage and 

dropped the tank pressure to its lower limit. Figure 36(c) illustrates the tank conditions at the end of 

the repressurization phase, when the tank reached its upper pressure limit once again.
32

T (°R)

62.8

60.2

47.6

44.8

42.2

38.6

36.8

60.3

48.2

46.1

44.0

41.8

39.8

37.8

(33.5)

(26.8)

(25.6)

(24.4)

(23.3)

(22.1)

(21)

T (°R)

68.1

64.7

61.4

48.0

44.8

41.2

37.8

T (K)

(37.8)

(35.9)

(34.1)

(26.7)

(24.9)

(22.9)

(21)

T (K)

(34.9)

(33.4)

(26.4)

(24.9)

(23.4)

(21.4)

(20.4)

(a) t = 31,658 s, Vmax = 0.0413 ft/s (b) t = 31,908 s, Vmax = 3.82 ft/s

(c) t = 37,898 s, Vmax = 0.0771 ft/s

T (°R) T (K)

= 1.26 cm/s = 1.16 cm/s

= 2.35 cm/s

Figure 36. CFD modeled velocity and temperature patterns during mixing—

90% fill, 6.7-W tank heat leak (case 3, with mixer).

Other aspects of case 3, which simulated the engineering checkout testing performed following boiloff characterization of the STUSTD tank, are summarized graphically in figures 37–39. 

The steady boiloff test determined a total system heat leak of 6.7 W. With the heaters off and 6.7 W 

evenly distributed along the tank walls, the STUSTD tank was locked up and allowed to pressurize to approximately 138 kPa (20 psia), at which point LH2 at 20.4 K (36.8 °R) was jetted axially at 

a rate of 102.2 lpm (27 gpm) into the bulk liquid to mix the tank contents and control tank pressure. 

In case 3, the RNG turbulence model was used instead of the original k-∈ model used in previous 

work.6 RNG is more stable and provides better results for axial jet mixing, and is recommended by 

the FLOW-3D developers. Once the tank pressure dropped to approximately 124 kPa (18 psia), the 

ATVS was shut down and the tank repressurized. During tank self-pressurization, the tank heaters 
33

were powered up to 40 W. The model assumed uniform saturated LH2 conditions at 110 kPa, 20 K 

(16.5 psia, 37.2 °R). The tank was assumed to be initially quiescent. Modeled ullage pressure rise 

rates agreed very well with available test data. During self-pressurization, the STUSTD tank reached 

145 kPa (20 psia) at a rate of 2 kPa/hr (0.29 psi/hr). The tank remained relatively quiescent with maximum velocities on the order of 1.2 cm/s (0.04 ft/s. Natural convection boundary layers were noticed 

as warm liquid was transported toward the free surface by buoyant forces. As expected, the ullage 

became thermally stratified. At the end of self-pressurization 145 kPa (20 psia), peak temperatures 

reached 32.3 K (58.1 °R) and the ullage was thermally stratified by approximately 11.1 K (20 °R). 

21

20

19

18

17

16

144.9

138

131.1

124.2

117.3

110.4

25,000 27,000 29,000 31,000 33,000 35,000

0.29 psi/hr (2 kPa/hr)

Time (s)

Ullage Pressure (psia)

Ullage Pressure (kPa)

Test Data

Model

0.92 psi/hr

(6.35 kPa/hr) –27 psi/hr (–186.3 kPa/hr)

–30.7 psi/hr

(–211.83 kPa/hr)

Figure 37. CFD model versus measured tank pressure during self-pressurization 

cycle—90% fill, 6.7-W tank heat leak (case 3, 102 lpm (27 gpm) mixer).

Test Data

Model

50

48

46

44

42

40

27.8

26.7

25.6

24.4

23.3

22.2

25,000 27,000 29,000 31,000 33,000

Mixer On

35,000

Ullage Temperature (°R)

Ullage Temperature (K)

Time (s)

Figure 38. CFD model versus measured tank ullage temperature during selfpressurization and mixing cycle—90% fill, 6.7-W tank heat leak 

(case 3, 102 lpm (27 gpm) mixer).
34

Test Data

Model

40

39

38

37

36

35

22.2

21.7

21.1

20.6

20

19.4

25,000 27,000 29,000 31,000 33,000 35,000

Mixer On

Liquid Temperature (°R)

Liquid Temperature (K)

Time (s)

Figure 39. CFD model versus measured liquid temperature at (76.3 cm/30 in) 

during self-pressurization and mixing cycle—90% fill, 6.7-W tank 

heat leak (case 3, 102 lpm (27 gpm) mixer.

The test data for case 3 are reported in reference 1. The ullage temperature data 112 and 

104 cm (44 and 41 in) and liquid temperature data 76 and 38 cm (30 and 15 in) from the bottom of 

the tank are available, but for only ≈1,500 s before and after ATVS activation. The model recorded 

ullage pressure and temperature 112 cm (44 in) and liquid temperatures 76 cm (30 in) from the bottom of the tank. Near the tank top, modeled ullage temperatures were higher than the measured test 

data. This may be due to constant heat leak boundary conditions that typically overpredict energy 

exchange between the tank walls and fluids. Both the model and test liquid temperatures showed 

negligible change during tank self-pressurization. 

At 31,658 s (fig. 37), the ATVS in case 3 was turned on and LH2 at 20.4 K (36.8 °R was 

injected into the bulk liquid at a rate of 102 lpm (27 gpm). The modeled average ullage pressure 

drop –186 kPa/hr (–27 psi/hr) during ATVS operation closely followed the test data –212 kPa/hr 

(–30.7 psi/hr). The jet penetrated the bulk liquid short of the free surface but generated enough mixing with the bulk liquid to drop the ullage thermal stratification to approximately 8.9 K (16 °R) and 

the peak temperature to 28.9 K (52 °R).

Following TVS activation, the heaters were turned on and operated at 40 W. The tank quickly 

began to repressurize at a rate of 6.35 kPa/hr (0.92 psi/hr). Approximately 6,000 s after the TVS 

phase, residual motion from the jet was coupled with heater-generated natural convection boundary layers and large-scale vortices continued to circulate within the liquid at 2.1 cm/s (0.07 ft/s). This 

residual motion was most likely responsible for the decrease in ullage temperature observed in the 

model results.

The self-pressurization and destratification model results illustrate steady ullage pressure rise 

rates and temperatures that agreed with the test data. Important cryogenic tank thermodynamic and 

fluid dynamic phenomena, such as natural convection boundary layers and ullage thermal stratifi-
35

cation, were reflected by the model. A summary of the normal gravity, self-pressurization results 

is given in table 4, and provides an overall perspective on how the average self-pressurization rate 

varied with liquid fill level and external heating. With a given heat leak, low fill fraction cases exhibited higher ullage pressure rise rates. For example, at 25.7 W, the 44% fill case (case 2) yielded a selfpressurization rate of approximate 10 kPa/hr (1.45 psi/hr), whereas at an 87% fill case (case 1), the 

ullage pressure rise rate was 6.9 kPa/hr (1 psi/hr). For the two fill levels studied, the incoming heat 

was absorbed by the higher thermal capacity liquid and reduced the tank pressurization rate. However, it is cautioned that experience with other tank geometries and/or propellants has demonstrated 

that the self-pressurization rates do not always vary with fill level in the same way.10

Table 4. STUSTD CFD model results summary.

Test Case

Heat Leak

(W)

Fill Level

%

dp/dt, kPa /hr 

(psi/hr)

1 25.7 87 6.9 (1)

2 25.7 44 10 (1.45)

3* 6.7 90 2 (0.29)

* Mixer on
36

5. THIRTY-DAY MISSION SIMULATION

The top-level purpose of the 30-day mission simulation test was to clearly demonstrate feasibility of a concept wherein LH2 tank pressure is controlled by venting all boiloff through the 

thruster. Concept success depends on managing tank heat load schedule such that the thruster usage 

rate simultaneously satisfies the overboard venting necessitated by tank pressure control for the duration of a mission, in this case for 30 days. More specific objectives are as follows:

(1) Demonstrate MLI purge evacuation and performance. 

(2) Deliver near-saturated hydrogen vapor at 158.7–241.5 kPa (23–35 psia) at 0.91 kg/hr 

(2 lbm/hr) for 140 cycles over 30 days. 

(3) Demonstrate venting/lockup for 140 burns over 30 days (without additional overboard 

venting required).

(4) Demonstrate use of MLI and heaters to assist control of tank pressure at 310±27.6 kPa 

(45±4 psia) while operating.

(5) Demonstrate subcooled LAD/TVS operation in 1 g via determination of internal LH2

tank temperature profile.

(6) Demonstrate LH2 loading/unloading and LAD filling for ground operations.

The mission simulation approach and test results are described in sections 5.1 and 5.2.

5.1 Mission Simulation Approach

The generalized mission sequence of events was as follows: 

(1) Perform MLI purge during tank fill.

(2) Launch preparations.

(3) Ascent flight simulation within the test facility rapid pump-down capabilities.

(4) On-orbit tank self-pressurization with mixing.

(5) Initiate 30-day mission simulation phase.

(6) Monitor system responses to mission transients.

5.2 Mission Simulation Testing and Evaluation Results

The mission simulation testing consisted of the launch and ascent flight phase followed by 

140 simulated engine burns during the 30-day, on-orbit simulation phase. Due to the small tank leak, 
37

the elevated chamber pressure during the on-orbit simulations necessitated adjustments in the engine 

burn flow rates and durations to compensate for the elevated tank pressures. Details regarding various mission phase simulations and the scaling adjustments are presented in sections 5.2.1 and 5.2.2.

5.2.1 Launch Preparations and Ascent Flight

The vacuum chamber was filled with GN2 at 1 atm and the LH2, tank inerted with GHe in 

preparation for the simulated ground fill process. As shown in figure 40, the fill process took the 

first hour; the pressure surge to 172.5 kPa (25 psia) at 0.2 hr occurred when LH2 first appeared in 

the tank, then settled out at about 124.1 kPa (18 psia), as expected. The tank lockup and simulated launch began at ≈1 hr. The initial chamber pressure drop and tank pressure surge is shown in 

figure 40. The chamber pressure follows the shuttle launch profile quite well until 1 hr, at which time it 

begins to deviate. The initial tank pressure jump to approximately 179.4 kPa (26 psia) was predicted, 

based on over 5 kW of ambient heat leak into the tank during the initial phases of launch. The vacuum 

chamber pumps shut down at 1.005 hr, which is why the tank pressure continued to climb. Although 

the chamber vacuum pumps were restarted after a couple of minutes, <30 torr was not achieved until 

after the tank pressure had reached 310.5 kPa (45 psia). When the tank reached 310.5 kPa (45 psia), it 

was vented down to about 241.5 kPa (35 psia), where, due to relatively high vacuum chamber pressure

(10–20 torr) the pressure again increased to nearly 345 kPa (50 psia), and necessitated venting. The 

chamber and tank pressures leveled out at about 1.5 hr, i.e., about 0.5 hr after the simulated launch 

began.

60

50

40

30

20

10

0

800

700

600

500

400

300

200

100

0

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

Time (hr)

Chamber Pressure (psia)

Chamber Pressure (torr)

Tank Pressure

Chamber Pressure

Figure 40. Measured tank and vacuum chamber pressures between ground fill 

and on-orbit simulation.

As indicated in figure 41, the chamber pressure followed the shuttle launch profile quite well 

until 1 hr, at which time it began to deviate because of the chamber pump shut-down. Also, the diffusion pumps cannot maintain the shuttle ascent profile after the free molecular flow regime begins 

in the 10–2 to 10–3 torr range.
38

Because of the inability of the vacuum chamber pumping train to follow the Shuttle Launch 

profile (reaching hard vacuum in 10–15 min), the capability of the MLI blankets to evacuate GHe 

rapidly enough to maintain the LH2 tank pressure below 206.8 kPa (30 psia) could not be verified. 

However, it was demonstrated that the MLI purge system worked well in allowing LH2 loading in 

GN2 at 1 atm.

Because of the launch simulation problems, the vacuum chamber was pumped down to the 

lowest pressure the leak would allow (–130 μtorr at 310.5 kPa (45 psia) and the LH2 tank was vented 

down to 172.5 kPa (25 psia) and refilled with LH2 prior to starting the 30-day mission.

5.2.2 On-Orbit 30-Day Mission

The 30-day mission simulation testing and analytical modeling using the standard model 

described in section 4.1 are presented in this section. The analytically modeled pressure excursions 

for each of the 140 burns are shown in figure 42; further details are in appendix D. The pressures are 

arranged as 311 kPa + 0 –Y (45 + 0 –Y psia), rather than as 311 kPa ±Y kPa. The maximum predicted pressure excursion at the penultimate apogee burn (139) was 54.4 or ±27.2 kPa (7.88 or ±3.94 

psia). The transitory effects of the ullage conditioning vents are shown. These vents were designed to 

mix the ullage and reduce the ullage temperature stratification to simulate conditions more representative of low-g operation.

A variety of factors are responsible for the measured LH2 tank pressure before and after 

each of the 140 engine burn simulations presented in figure 42. The sharp rise in tank pressure for 

burns l–3, 13, and 14 was because the ATVS was off during these burns. The more moderate rise 

for burns 4–12 and 15–17 showed that, in spite of the excessive heat leak, the ATVS would have 

removed sufficient thermal energy to assure pressure control had the coast periods and burn durations been adjusted to compensate for the high chamber pressures, i.e., burns were not long enough 

for the ATVS and PTVS to remove the extra heat. Note that, for burns 18 and 19, the tank pressure 

dropped; this was due to burn durations l5 min longer (1,700 s versus 800 s). Even with the ATVS off 

Chamber Pressure

60 800

50

40

30

20

10

0

700

600

500

400

300

200

100

0

0.9 0.95 1

Time (hr)

1.05 1.1

Chamber Pressure (torr)

Chamber Pressure (psia)

Tank Pressure

Shuttle

Launch Profile

Figure 41. Measured LH2 tank and vacuum chamber 

pressure versus shuttle ascent profile.
39

during these burns, the PTVS was able to remove more than enough heat, thanks to the longer burn 

times. In figure 43, the tank pressure drop during the burns is the difference between the measured 

tank pressure before and after each burn.

46

45

44

43

42

41

40

39

38

37

317.4

310.5

303.6

296.7

289.8

282.9

276

269.1

269.2

255.3

0 20 40 60 80 100 120 140

Tank Pressure Excursions (psia)

Tank Pressure Excursions (kPa)

Burn Number

Figure 42. Predicted LH2 tank pressure as a function of MET.

Burn Number

LH2 Tank Pressure (psia)

LH2 Tank Pressure (kPa)

55

53

51

49

47

45

43

41

39

37

35

379.5

365.7

351.9

338.1

324.3

310.5

296.7

282.9

269.1

255.3

241.5

0 20 40 60 80 100 120 140

Tank Pressure Before Burn

Tank Pressure After Burn

Figure 43. Measured LH2 tank pressure before and after burn as a 

function of burn number.

As shown in table 5, the tank pressure drops during the first 20 burns were as expected; 

however, due to stratification effects on PTVS performance, the actual pressure drops for burn numbers 40–139 were 46%–55% less than the predicted values.
40

Table 5. Comparison of burn pressure drop data with predictions.

Tank Pressure Change During Burn

Burn Number Modeled, kPa (psid) Actual, kPa (psid)

20 4.7 (0.69) 4.8 (0.70)

40 9.8 (1.42) 5.3 (0.77)

60 14.3 (2.08) 6.48 (0.94) 

80 18.42 (2.67) 8.14 (1.18)

100 22.29 (3.23) 11.32 (1.64)

120 26.4 (3.83) 16.15 (2.34)

122 40.23 (5.83) 31.05 (4.50)

136 44.85 (6.50) 35.54 (5.15)

139 54.37 (7.88) 38.85 (5.63)

After burns 20, 21, 33, 48, 109, and 117, ullage conditioning vents were performed. The conditioning effects on LH2 tank and chamber pressures are obvious in figure 44 and 45, respectively. 

Conditioning effects on ullage temperature are shown in figure 45. It can be seen that, as the tank 

emptied, the average chamber pressure decreased whereas the maximum ullage temperature steadily 

increased, as expected.

0

140

130

120

110

100

90

80

70

60

50

40

20 40 60 80 100 120 140

Burn Number

Average Chamber Pressure (

μtorr)

Figure 44. Average chamber pressure as a function of burn number.
41

0

40

45

50

55

60

65

20 40 60 80 100 120 140

22.2

25

27.8

30.6

33.3

36.1

Maximum Ullage Temperature (°R)

Maximum Ullage Temperature (K)

Burn Number

Figure 45. Maximum ullage temperature as a function of burn number.

A detailed plot of the ullage conditioning vent following burn 48 is shown in figure 46. When 

the valve was actuated for the burn, the ATVS pump was switched from low speed during coast to 

high speed during burn. The ATVS pump was returned to low speed during the ullage conditioning vent, which was indicated by the high vent flow, and the sharp drop in tank pressure, to 297 kPa 

(43 psia). Note that the tank pressure rebounded to 300 kPa (43.5 psia) when the ullage vent was 

closed. This was typical behavior whenever a vent or burn was stopped, and was caused by transient

338.1

331.2

324.3

317.4

310.5

303.6

296.7

289.8

0 0.5 1

Burn Number

1.5 2

30

25

20

15

10

5

0

Average Tank Pressure (kPa)

Vent Flow (kg/hr)

Tank Pressure

Vent Flow

Figure 46. Tank pressure and ullage vent flow rate during ullage vent.
42

LH2 boiling as a result of the tank pressure reduction. This characteristic persists until the selfpressurization is sufficient to suppress the boiling.

During a tank conditioning vent or depressurization, the boiling is very effective in destratifying and cooling the ullage, as shown in figure 47. The top four ullage temperatures (T1–T4) converge 

at about 46 °R (25.6 K), which is essentially the measured LH2 temperature. 

58

56

54

52

50

48

46

44

32.2

31.1

30

28.9

27.8

26.7

25.6

24.4

0 0.5 1

Time (hr)

1.5 2

TT1

Temperature (K)

Temperature (°R)

TT4

TT3

TT2

Figure 47. Ullage temperature during ullage vent.

Burns 22–43 were performed with the ATVS off: the steady increase in pressure reflected the 

inability of the PTVS to extract sufficient heat to keep up with the increased heat flow into the tank 

and the long coast times. On burns 44–140 (the rest of the mission), the ATVS was on. This was done 

in recognition of the fact that the extra heat flux from high vacuum chamber pressure made control 

of the tank pressure intractable without using all of the vent capability available (ATVS plus PTVS).

The variations in tank pressure also affected the overall vent flow rate, as shown in figure 48. 

The pressure affected both the flow through the VISCO JETs® and the flow through the leak. Note 

that the flow rate was generally dropping as the tank emptied. This implied that there was some head 

effect on the leakage flow. (Head should have had essentially no effect on VISCO JET flow.)
43

8.0

7.5

7.0

6.5

6.0

5.5

5.0

4.5

4.0

3.6

3.4

3.2

3

2.7

2.5

2.3

2

1.8

0 20 40 60 80

Burn Number

Average Flow Rate (kg/hr)

Average Flow Rate (lbm/hr)

100 120 140

Figure 48. Average feed line flow rate as a function of burn number.

Returning to figure 43 at burn 56, there was a significant change in pressure control capability. This was where the coasts were reduced (from approximately 7,200 s to 4,800 s) to correct for 

the increased MLI heat flux. The ATVS, together with the PTVS, was able to extract enough heat to 

more than keep up with heat input from the shortened coast periods. On burns 60 and 64, the ATVS 

was turned off; note the immediate jump in tank pressure for these burns. This series of burns, from 

56 to 80, showed that the ATVS and PTVS could effectively maintain constant tank pressure, as 

designed.

At burn 80, the tank pressure again trended upward, as the shortened coasts had also reached 

7,500 s. This continued until burn 90, when the coast time was again reduced (from approximately 

9,100 s to 7,100 s). However, by burn 100, coast times had again reached 9,100 s, and ullage conditioning 

vents were required at burns 104, 109, and 117 to control the tank pressure below approximately 331 kPa 

(48 psia) until burn 121. Burn 121 was increased from 749 s to 2,187 s to simulate the STUS apogee burns 

(121–140). Even though the coast times for burns 121 to 128 increased from approximately 15,000 s to 

20,000 s, the longer burn times allowed the ATVS and PTVS to control tank pressure. After run 128, 

the combination of long coast times (approximately 20,000 to 27,000 s) and excessive MLI heat leak 

from high vacuum chamber pressure overcame the ability of the ATVS and PTVS to maintain 

constant tank pressure. Fortunately, run 139, with the final long coast, was reached before the tank 

pressure reached 349 kPa (49 psia). 
44

6. SUMMARY AND CONCLUSIONS

The programmatic strategy of the AITP was to assemble a technology team (comprised of 

individuals and small groups representing industry government, academia, and small business) that 

would enable the sharing of technology skills and facilities in a cost effective and productive manner. 

The STUSTD program was a major element of the cryogenic solar propulsion AITP cooperatively 

performed by a Boeing-led team. The technical strategy, hardware performance, test results, and 

analytical modeling accomplishments are summarized below.

6.1 Technical Strategy and Test Hardware Performance

The STUSTD technical strategy was to balance the LH2 storage tank venting requirements 

with the engine thrusting timeline during a 30-day solar thermal propulsion mission, thereby assuring that there were no vent losses. Proof-of-concept testing was performed at MSFC’s TS300 vacuum 

chamber using a 2 m3 (71 ft3) LH2 tank with an arrangement of thermodynamic venting combined 

with a capillary screen channel LAD, MLI, and internal tank heaters. An engineering test series was 

first conducted to check out the test equipment and to calibrate or anchor analytical modeling. The 

checkout testing was followed by a 30-day mission simulation test designed to demonstrate the concept feasibility.

The checkout testing began with an ascent flight simulation that was compromised by vacuum 

chamber pump-down limitations and a chamber control system malfunction. The on-orbit checkout 

phase began with a LH2 boiloff test that indicated a system heat leak of 6.7 W versus a predicted 

heat leak of 6.6 W, i.e., the actual heat leak was within 1.5% of the prediction. However, upon activation of the ATVS, a small H2 leak caused the chamber pressure to elevate from 6.75×10–6 to 5×10–5

torr and increased the tank heat leak to 18 W. The testing was continued, recognizing that analytical 

modeling would have to reflect the heat leak increase. 

Analytical correlations of the engineering checkout test data were initially performed primarily with standard analytical modeling tools and limited CFD modeling. The initial modeling was 

later bolstered by a comprehensive CFD correlation effort involving the checkout data. However, 

due to budgetary constraints, only the standard modeling was applied to the 30-day mission.

6.2 Standard Modeling Test Correlations

The original PTVS thermal extraction rate prediction was based on a perfectly mixed ullage 

at saturated conditions plus a LAD length based on axial height instead of true exposed LAD 

length. Consequently, in every case, the measured thermal extraction rates exceeded the predictions 

by a factor of 10 with the PTVS. The ‘higher than expected’ extraction rates were attributed to the 

‘warmer than expected’ ullage due the previously mentioned elevated MLI heat leak in combination 

with inadequate TVS pump mixing. (See CFD modeling results in sec. 6.3.) Modifications to the 

standard analytical algorithms and simulations to account for the relatively warm ullage and actual 
45

LAD length (and, therefore, increased thermal extraction of the PTVS) were made before the 30-day 

mission simulation. The standard analytical modeling, anchored by the engineering checkout test 

data and the CFD mixer simulations, enabled pretest predictions that correlated quite well with the 

30-day mission simulation test results.

6.3 Engineering Test Results and Computational Fluid Dynamics Modeling

Based on previous experience with reduced gravity pressure control within the S-IVB LH2

tank, a 3,366-cell axisymmetric mesh was selected to simulate the STUSTD LH2 tank.6 Application 

of this CFD model to the engineering checkout testing yielded the following conclusions: 

• The TVS mixer was unable to produce a geyser that penetrated or disturbed the liquid-vapor interface sufficiently to achieve adequate mixing with fill levels of 90% and 40%.

• With a 90% fill level (10% ullage), the analytical and measured self-pressurization rates correlated 

quite well prior to the mixing cycle; however, the modeled rate was three times the measured rate 

following the mixing cycle. 

• Stabilized self-pressurization conditions prior to the first mixing cycle are more easily simulated 

than the transient thermodynamic conditions immediately following a mixing cycle. 

• Modeling ullage and liquid temperatures were more successful. Compared with measured data, 

modeled ullage and liquid temperatures were generally within 10% and 2%, respectively.

Therefore, even though additional ullage pressure control modeling improvements are needed, 

to a significant degree, tank heater-liquid dynamics, natural convection boundary layers, ullage thermal stratification, and ATVS operation were captured by the numerical simulations. Prescribed 

boundary conditions at the tank walls combined with temporary boiling heat transfer and residual 

stratification effects represent substantial limitations in self-pressurization CFD modeling between 

mixing cycles. However, this limitation can be addressed by setting up a computational interface 

between a thermal code, such as SINDA and the CFD code, thereby enabling the tank calculated 

structural temperatures to interact with calculated interior thermodynamics and vice versa.

6.4 Standard Plus Computational Fluid Dynamics Modeling Combination

CFD modeling has steadily improved each time it has been challenged with another set of 

cryogenic test data such as the subject STUSTD data. Therefore, these CFD and standard analytical 

correlations of an ellipsoidal LH2 tank with an axial jet mixer have, without a doubt, substantially 

improved the analytical code validation database for two-phase, on-orbit cryogen storage. In 

fact, existing CFD codes in combination with traditional models can already be used to at least 

bracket reduced gravity effects. An overwhelming justification for continued development of CFD 

modeling lies in its innate capability to simulate reduced gravity effects on basic thermodynamics, 

fluid dynamics, and heat transfer. It is this reduced gravity capability that can provide the ‘bridge’ 

or scaling between normal and reduced gravity CFM. Eventually it will become standard practice 

to use the CFD codes to anchor traditional codes for reduced gravity applications, thereby enabling 

rapid turnaround modeling with the less costly and user-friendly traditional codes.
46

Summation statement: Substantial strides have occurred in recent years as CFD codes continue 

to be challenged with various sets of cryogenic test data and computer speeds increase. With each CFD 

code improvement increment, the historic dependence of in-space CFM development on prohibitively 

expensive on-orbit testing is being significantly reduced.
47

APPENDIX A—STUSTD INSTRUMENTATION LIST

Channel Measurement Description MFG Serial No. Range Units Cal Card Group

101 Chamber dew point E&H 2009pw-4 -112 to 68 F se 2

102 Chamber dew point E&H 2012pw-4 -112 to 68 F se 2

103 Ullage differential pressure MKS 536864 1 TORR se 6

calc Active TVS in/out delta pressure PSI 960094-A 0.5 PSID n/a

38 Raw millivolts DPT4 PSI 960094-A -50 TO 50 MV se

41 External tank heater 1 voltage FLEXCORE 8050052 50 VDC Jul-99 se

42 External tank heater 2 voltage FLEXCORE 8050053 50 VDC Jul-99 se

43 External tank heater 3 voltage FLEXCORE 8050054 50 VDC Jul-99 se

44 External tank heater 4 voltage FLEXCORE 8050055 50 VDC Jul-99 se

104 LH2 Vent flow rov 20-671 FTI 2402000 165 ACFM May-00 se 8

105 LH2 Vent flow rov 20-672 FTI 1606628 55 ACFM May-00 se 8

106 LH2 Vent flow rov 20-673 Hastings 10879 20 SCFM May-00 se 8

107 LH2 Vent flow rov 20-674 MKS 0558A01738232 2918 SCIM May-00 se 8

139 Engine feedline flow rate HASTINGS 18018 10 SCFM se

140 MLI He purge rate FTI 807219 1 ACFM se

100 ISUS Vent flow rate Flow Tech., Inc. 16010078 1760 ACFM Oct-00

50 TVS pump frequency 0-400 HZ n/a

70 Tank heater On / Off 1000 OFN n/a

45 External tank heater 1 current FLEXCORE 8050107 1 AMP Jul-99 se

46 External tank heater 2 current FLEXCORE 8050108 1 AMP Jul-99 se

47 External tank heater 3 current AAC 3498 5 AMP Jul-99 se

48 External tank heater 4 current AAC 3499 5 AMP Jul-99 se

51 TVS current leg A 0-5 AMP n/a

52 TVS current leg B 0-5 AMP n/a

53 TVS current leg C 0-5 AMP n/a

108 Ullage/Vent line pressure MKS 1000 TORR se 6

66 Ullage/Vent line pressure TELEDYNE TABER 781298 50 PSIG Jul-99 fb 6

110 Reference pressure vessel MKS 1000 TORR se 6

111 Ullage pressure MKS 96 PSIA se

112 GN2 supply on roof STELLAR TECH 940849 5000 PSIG Jul-99 fb 14

113 GN2 to ejector STELLAR TECH 951773 1000 PSIG Jul-99 fb 14

114 GN2 supply to digicell DYNISCO 24651 3000 PSIG Jul-99 fb 14

115 Digicell loader operator TELEDYNE TABER 880264 5000 PSIG Jul-99 fb 14

116 2nd stage firex STATHAM 15 1000 PSIG Jul-99 fb 24

118 LN2 Tank ullage pressure TELEDYNE TABER 781319 100 PSIG Jul-99 fb 11

119 LN2 Pump line pressure TELEDYNE TABER 860548 300 PSIG Jul-99 fb 11

120 GN2 Supply pressure MB ELECTRONICS 41029 100 PSIG Jul-99 fb 11

121 Air manifold TELEDYNE TABER 888102 5000 PSIG Jul-99 fb 14

122 Repress stage two DYNISCO 27686 1500 PSIG Jul-99 fb 14

123 Repress stage one TELEDYNE TABER 930869 1000 PSIG Jul-99 fb 8,10,14
48

Channel Measurement Description MFG Serial No. Range Units Cal Card Group

124 LH2 F/D trailer fill TELEDYNE TABER 880705 100 PSIG Jul-99 fb 13

125 LH2 Vent sytem TELEDYNE TABER 781316 100 PSIG Jul-99 fb 13

126 LH2 Vent sytem @ pressure control valves TELEDYNE TABER 931250 100 PSIG Jul-99 fb 8,13

127 LN2 Supply system rov 20-1700 TELEDYNE TABER 891412 100 PSIG Jul-99 fb 10

128 GH2 Upstream DLR - 803 TELEDYNE TABER 932076 6000 PSIG Jul-99 fb 10

129 GH2 Downstream DLR - 803 STELLAR TECH 941210 1000 PSIG Jul-99 fb 10

130 GH2 Feed line TELEDYNE TABER 950379 500 PSIG Jul-99 fb 10

131 GN2 Downstream DLR-2082 TELEDYNE TABER 890938 2000 PSIG Jul-99 fb 10

132 GHE Upstream DLR-2304 DYNISCO 32121 1000 PSIG Jul-99 fb 17

133 GHE Downstream DLR-2304 TELEDYNE TABER 751800 200 PSIG Jul-99 fb 17

134 Tank pressurization TELEDYNE TABER 890934 2000 PSIG Jul-99 fb 11

141 Helium purge of MLI STATHAM 89 2 PSIG Aug-99 fb

165 ISUS Vent flow pressure TELEDYNE TABER 891351 50 PSIA May-00

166 Downstream burts disk TELEDYNE TABER 860736 100 PSIG Oct-00 fb

69 LAD pump on / off 100 PER Jul-99

136 Feedline pressure downstrm micrometer valve CEC 2359 50 PSIA Jul-99 se

calc Feedline pressure downstrm heat xchanger SENSOTRON 1517-3 50 PSIA

167 Feedline pressure downstrm heat xchanger TELEDYNE TABER 850786 100 PSIA Nov-00 fb

39 Raw millivolts PT4 SENSOTRON 1517-3 139 TO 

540

R n/a se

204 LN2 pump line temperature TYPE E 27 TO 558 R se 11

205 LN2 Return line temp. TYPE E 27 TO 558 R se 11

216 Diff pump #1 water outlet TYPE K -100 TO 

500

K se 11

217 Diff pump #2 water outlet TYPE K -100 TO 

500

K se 11

218 Diff pump #1oil TYPE K -100 TO 

500

K se 11

219 Diff pump #2 oil TYPE K -100 TO 

500

K se 11

220 LH2 F/D rov 20-601 TYPE E 27 TO 558 R se 13

222 LH2 F/D rov 20-676 TYPE E 27 TO 558 R se 13

224 LH2 Vent line @ chamber penetration TYPE E 27 TO 558 R se 13

225 LH2 Vent line flow box TYPE E 27 TO 558 R se 8,13

226 LN2 Heat Exchanger #1 TYPE K -100 TO 

500

K se 14

227 LN2 Heat Exchanger #2 TYPE K -100 TO 

500

K se 14

228 LN2 Heat Exchanger #3 TYPE K -100 TO 

500

K se 14

231 LN2 heat exchanger rov 20-1704 TYPE E 27 TO 558 R se 17

232 LN2 supply rov 20-1700 TYPE E 27 TO 558 R se 10

233 LH2 vent heat exchanger water discharge temp TYPE E 27 TO 558 R se 13

234 Flowbox temperature - feedback TYPE E 27 TO 558 R se 8

235 Ice bath reference pressure vessel temp TYPE E 27 TO 558 R se 6

236 DP1 hoffman box temperature TYPE E 27 TO 558 R se 6

237 Engine feedline temp before ROV's TYPE E 27 TO 558 R se

238 MLI He purge temperature TYPE E 28 TO 558 R se
49

Channel Measurement Description MFG Serial No. Range Units Cal Card Group

239 ISUS Vent flow temperature TYPE E 27 to 558 R se

240 Cabinet 39 air temperature TYPE E 28 to 558 R se

241 Chamber air temperature TYPE E 28 TO 558 R se

142 Ring 1 #1 TYPE E 27 TO 558 R se

143 Ring 1 #4 TYPE E 27 TO 558 R se

144 Ring 1 #6 TYPE E 27 TO 558 R se

145 Ring 1 #8 TYPE E 27 TO 558 R se

146 Ring 2 # 11 TYPE E 27 TO 558 R se

147 Ring 2 # 15 TYPE E 27 TO 558 R se

148 Ring 3 # 21 TYPE E 27 TO 558 R se

149 Ring 3 # 25 TYPE E 27 TO 558 R se

150 Ring 4 #27 TYPE E 27 TO 558 R se

151 Ring 4 #31 TYPE E 27 TO 558 R se

152 Ring 5 # 37 TYPE E 27 TO 558 R se

153 Ring 5 # 41 TYPE E 27 TO 558 R se

154 Ring 6 # 43 TYPE E 27 TO 558 R se

155 Ring 6 # 47 TYPE E 27 TO 558 R se

156 Ring 7 # 53 TYPE E 27 TO 558 R se

157 Ring 7 # 57 TYPE E 27 TO 558 R se

158 Ring 8 # 59 TYPE E 27 TO 558 R se

159 Ring 8 # 63 TYPE E 27 TO 558 R se

160 Ring 9 # 67 TYPE E 27 TO 558 R se

161 Ring 9 # 69 TYPE E 27 TO 558 R se

162 Ring 9 # 71 TYPE E 27 TO 558 R se

163 Ring 9 # 73 TYPE E 27 TO 558 R se

164 Ring 9 # 74 TYPE E 27 TO 558 R se

37 DPT4 diaphragm temperature (active TVS region) LAKESHORE 27 to 855 R ce

40 PT4 diaphragm temperature SENSOTRON 140 / 540 R se

1 Tank fluid 44 inch LAKESHORE 27 to 855 R ce

2 Tank fluid 43 inch LAKESHORE 27 to 855 R ce

3 Tank fluid 41 inch LAKESHORE 27 to 855 R ce

4 Tank fluid 40 inch LAKESHORE 27 to 855 R ce

5 Tank fluid 37.5 inch LAKESHORE 27 to 855 R ce

6 Tank fluid 35 inch LAKESHORE 27 to 855 R ce

7 Tank fluid 32.5 inch LAKESHORE 27 to 855 R ce

8 Tank fluid 30 inch LAKESHORE 27 to 855 R ce

9 Tank fluid 27.5 inch LAKESHORE 27 to 855 R ce

10 Tank fluid 25 inch LAKESHORE 27 to 855 R ce

11 Tank fluid 22.5 inch LAKESHORE 27 to 855 R ce

12 Tank fluid 20 inch LAKESHORE 27 to 855 R ce

13 Tank fluid 17.5 inch LAKESHORE 27 to 855 R ce

14 Tank fluid 15 inch LAKESHORE 27 to 855 R ce

15 Tank fluid 12.5 inch LAKESHORE 27 to 855 R ce

16 Tank fluid 10 inch LAKESHORE 27 to 855 R ce

17 Tank fluid 7.5 inch LAKESHORE 27 to 855 R ce

18 Tank fluid 6 inch LAKESHORE 27 to 855 R ce
50

Channel Measurement Description MFG Serial No. Range Units Cal Card Group

67 Tank fluid 5 inch LAKESHORE 27 to 855 R ce

20 Tank fluid 4 inch LAKESHORE 27 to 855 R ce

21 TVS 10 inch from LAD bottom LAKESHORE 27 to 855 R ce

22 TVS 10 inch from LAD top LAKESHORE 27 to 855 R ce

23 Inlet line of active TVS LAKESHORE 27 to 855 R ce

24 Outlet line of active TVS LAKESHORE 27 to 855 R ce

25 MLI tank surface, 5/8" below equator at 90 deg LAKESHORE 27 to 855 R ce

26 MLI tank surface, 17.5" below equator at 90 deg LAKESHORE 27 to 855 R ce

27 MLI layer 25, 90 deg at equator LAKESHORE 27 to 855 R ce

28 MLI layer 50, 90 deg at equator LAKESHORE 27 to 855 R ce

29 MLI layer 75, 90 deg at equator LAKESHORE 27 to 855 R ce

49 Feedline heat xchanger surface center LAKESHORE 27 to 855 R ce

68 Feedline fluid temp LAKESHORE 27 to 855 R ce

32 LAD vent line fluid temp LAKESHORE 27 to 855 R ce

33 Feedline heat xchanger surface outlet LAKESHORE 27 to 855 R ce

34 Feedline heat xchanger surface inlet LAKESHORE 27 to 855 R ce

35 Feedline heat xchanger fluid outlet LAKESHORE 27 to 855 R ce

36 Feedline heat xchanger fluid inlet LAKESHORE 27 to 855 R ce

71 ROV 20-1854 position 1000 OFN n/a se

72 ROV 20-1856 position 1000 OFN n/a se

73 VPV 20-661 position 100 PER n/a se

74 VPV 20-662 position 100 PER n/a se

75 VPV 20-663 position 100 PER n/a se

76 ROV 20-1859 open indication 1000 OFN n/a se

77 ROV 20-1859 close indication 1000 OFN n/a se

178 Diffusion pump primer GRAN.-PHILLIP 10-3 TO 

760

TORR se 23

179 LN2 tank vacuum jacket GRAN.-PHILLIP 10-3 TO 

760

TORR se 23

180 LH2 F/D vac jacket @ rov 1502 GRAN.-PHILLIP 10-3 TO 

760

TORR se 13

181 LN2 tank vacuum jacket pump (ctrl GE room) GRAN.-PHILLIP 10-3 TO 

760

TORR se 23

182 Chamber wall box MKS 1,000.00 se 23

183 Chamber wall box MKS 1,000.00 TORR se 23

184 Pump one foreline (outside pump house) GRAN.-PHILLIP 10-3 TO 

760

TORR se 24

calc Internal chamber pressure MKS 0.01 TO 

1000 

UTORR 6

calc Chamber wall box GRANVIL-PHILLIP 0.01 TO 

1000 

UTORR 6

calc Internal chamber pressure GRANVIL-PHILLIP 0.01 TO 

1000 

UTORR 6

188 RGA Pressure MKS 1,000.00 se 24

calc RGA Pressure GRANVIL-PHILLIP 0.01 TO 

1000 

UTORR 24

192 GRAN.-PHILLIP 10-3 TO 

760

TORR se 24
51

Channel Measurement Description MFG Serial No. Range Units Cal Card Group

193 test article v.j. piping - top level GRAN.-PHILLIP 10-3 TO 

760

TORR se 13

194 GRAN.-PHILLIP 10-3 TO 

760

TORR se 24

185 Raw millivolts for VP3010 MKS 5300 MV se

186 Raw millivolts for VP3011 GRANVILL-PHILLI -120 MV se

187 Raw millivolts for VP3012 GRANVIL-PHILLIP 4000 MV se

189 Raw millivolts for VP3014 GRANVIL-PHILLIP -120 MV se
52

APPENDIX B—AEROSPACE-INDUSTRY TECHNOLOGY PROJECT 

SUPPORTING PERSONNEL

EP91

EP92 Kevin Pedersen Project Engineer/Mechanical Systems Engineer

EP93 Robert Lake Instrumentation Engineer

EP93 Jim Crawford Data Systems Engineer

EP94 Jacob Yarbrough Control Systems Engineer

LB&B Technicians:

Mech David Cole Lead

Mech Bo Jones

Mech Culley Cantrell

Instr Ralph Keller Lead

Instr Doug McBride

Instr Tabitha Neely

Cont Joel Wheeler Lead

Cont Greg Wirt

Cont Cleveland Green

Support Personnel:

CR20 Rosalyn Patrick Safety

LB&B John Webster Safety

HEI Alvin Eidson Safety

HEI Bill Horn Quality

LB&B Dennis Crawford Quality

LB&B Donald Lott Welder/Fitter

LB&B Van Gassoway Welder/Fitter

CSC Fausto Latini Data Systems Support

CSC Wade Farris Data Systems Support

LB&B Paul Readus Drafting

LB&B Tony Hightower Drafting

LB&B Paula Nave Documentation

LB&B Tammie Rockhill Configuration Control

LB&B Jon Keel Area Coordinator

LB&B Ron Beasley Area Coordinator
53

During Test Only:

EP92 Dennis Strickland Test Conductor

EP92 Melanie Ramsey Test Conductor

EP92 Alan Murphy Test Conductor

EP92 Collie Kellet Test Conductor

EP93 Keary Smith Instrumentation Engineer

EP93 John Wiley Instrumentation Engineer

EP93 Curtis Thompson Instrumentation Engineer

EP93 Jason Elmore Data Systems Engineer

EP94 David Jones Control Systems Engineer

EP94 Ron Smith Control Systems Engineer

EP94 Judith Gregory Control Systems Engineer

CSC Jay Carter Data Systems Support

CSC David Jarrell Data Systems Support

CSC Frankie Hinkle Data Systems Support

CSC Ken Dodd Data Systems Support

LB&B Larry Billions Control Systems Technician
54

APPENDIX C—STUSTD LIQUID HYDROGEN SPACE FLIGHT SYSTEMS 

TEST REQUEST SHEET (1/25/99) 

STUSTD Boil-off & Engineering Checkout

Test Name: STUSTD LHSFS Boil-off and Engineering Checkout 

(Sections 5.1, 5.2, 5.5, 5.6, 5.10, and 5.11 of the STUSTD Test Plan). Sections 5.3 [Simulated Launch] and 5.4 [Boiloff Without Cold Walls] are no longer part of the boil-off or Engineering Checkout plans. The Simulated Launch will 

be part of the 30-day Mission Simulation.

Test Project ID: Test Requester: Ed Cady, Al Olsen, [Boeing-HB], Dan Vonderwell, [Leon 

Hastings, Robin Flachbart]

Test Date: Feb. 3, 1999 Test Conductor: Kevin Pedersen

Test Objectives: 

• Develop a data base for the performance of the Solar Thermal Upper Stage Technology Demonstrator (STUSTD) Liquid 

Hydrogen Storage and Feed System (LHSFS) before the 30-day mission simulation. Data obtained from these tests will 

characterize and verify the expected performance of the passive and active TVS, the tank heaters, and the valve system.

• Fill the STUSTD tank with LH2 and monitor thermodynamic characteristics (Sections 5.1 and 5.2 of the Test Plan).

• Measure boil-off with LN2-cold-walls to determine the heat leak into the tank system (Section 5.5 of the Test Plan)

• The tank will be slowly pressurized from the pressure maintained during the boil-off tests (Section 5.5 of the STUSTD Test Plan, 

16.5 to 17 psia) to the operating pressure of 45 psia. The flow metering valve will be calibrated to 2.0 lbm/hr. The Active 

Thermodynamic Venting System (ATVS) and Passive Thermodynamic Venting System (PTVS) thermal extraction capability 

will be characterized as a function of burn (vent) time and tank fill level.

• The tank pressure will be controlled to 45 ± 4 psia throughout the entire Engineering Checkout process.

• After the Engineering Checkout process, the tank will be drained and safed per sections 5.10 and 5.11. 

Test Fluid: LH2

Approximate Test Duration: ~312 hrs (13 days) [refer to attached Engineering Test Process description and timeline and to the 

Test Plan]

DSU Sampling Rate: Per STUSTD LHSFS Test Plan

High: 1 Hz [1 sample per second] (used when test data is changing in a 

transitory manner, such as during a simulated engine burn or initialization of the ATVS)

Mid: 0.1 Hz [1 sample per 10 seconds] (used when the high data rate is not indicated) 

Low: 0.0167 Hz [1 sample per 60 seconds] (may be used during periods of relative inactivity as 

determined by the Boeing representative) 

Sampling Start/End Time: Data collection will continue from Section 5.1 (Tank safing), through Section 5.6 (Engineering 

Checkout), and terminate after the tank is safed again (Section 5.11)

Test data should be dumped to the location accessible by Boeing (i.e. the Jetson server) every two 

hours so that Boeing-HB can track the progress of the tests. This is a modification to the “every 

6 hours” called out in the Test Plan. The increased frequency of the data dumps is due to the fact 

that the frequently updated Excel file is no longer available. 

Vacuum Chamber/Pump: Evaculate the vacuum chamber to 1X10-6 Torr or less and maintain that value for at least 8 

hours prior to tanking and during the tank safing and inerting operation. . Hard vacuum 

conditions are to be maintained throughout the remainder of testing. The MLI GHe purge will 

not be used.

AT NO TIME IS THE PRESSURE EXTERNAL TO THE TANK (IN THE VACUUM 

CHAMBER) TO EXCEED THE PRESSURE INSIDE THE TANK. 

Vacuum Chamber Purge: Dry GN2 (dewpoint 395°R or lower), one atmosphere. 

Vacuum Chamber Coldwalls: Cold walls are used at steady state (~140°R) for Section 5.5 of the Test Plan; the cold walls will be 

maintained throughout the boil-off test and the Engineering Checkout (Sections 5.5 and 5.6). Cold 

walls will have to reach a steady-state value prior to the cold wall boil-off test. 
55

Back Pressure Control System: Setup ice bath and constant pressure volume/ice bath two days prior to testing to obtain steady state 

environment. Check ice bath at 1 to 2 hour intervals. Fill or pack as needed to prevent temperature 

drift. Document all operations in the test log

Flow Meter Box: Setup all flow meter heating purges to 39 o

C a minimum of 24 hours prior to test to obtain steadystate conditions on all hardware. Document all operations in the test log

Other Vacuum Systems: All vacuum jacketed systems are pumped down prior for three days prior to testing (Engineering 

Checkout) 

RGA Operation: Periodic as needed (write DSU times and file # on plots) 

Camera Operation: If possible (i.e. if the tapes are external to the vacuum chamber and can be replaced) it would be 

preferable to periodically operate at least one camera to capture any events internal to the chamber.

Test Article Internal Purge: Tank is purged with 7 ± 1 psig (~22 psia) GHe in Section 5.1 of the Test Plan. AT NO TIME IS 

THE PRESSURE EXTERNAL TO THE TANK (IN THE VACUUM CHAMBER) TO 

EXCEED THE PRESSURE INSIDE THE TANK.

Test Article Fill Level: 284 lbm LH2 prior to initiation of Section 5.6 of Test Plan. Filled until TT-04 reads liquid H2. 

 284 – 279 lbm LH2 during metering valve calibration

 279 – 270 lbm LH2 during high-fill-level ATVS / PTVS characterization

 270 – 165 lbm LH2 during “high power” vent

 165 – 159 lbm LH2 during medium-fill-level ATVS / PTVS characterization

 159 – 85 lbm LH2 during “high power” vent

 85 – 79 lbm LH2 during low-fill-level ATVS / PTVS characterization

Test Article Heat Guards: N/A

Test Article Shroud Purge: N/A

Test Article Shroud: N/A

Ground Hold Duration: Not Tested

Test Article Ullage Applicable only during boil-off tests (Sections 5.4 and 5.5 of Test Plan)

Pressure Control: 

Tank Replenishment: N/A. Tank is drained, safed, and inerted after the Engineering Checkout.

Chamber Pressurization: Hard vacuum (1X10-6 Torr) is maintained throughout the boil-off testing and the Engineering 

Checkout (Section 5.6).

Specific Test Hardware Parameters: 

Real time adjustments of some of the parameters listed below may be required based on actual measured STUSTD heat leak. Document all 

operations in the test log.

ATVS Pump: Absolute minimum setting: 22Hz (~400 RPM pump rate)

 “Low” setting: 30.5 Hz (550 RPM)

 “High” setting: 39 Hz (700 RPM)

 Absolute maximum setting: 50 Hz (900 RPM)

Tank Heaters: (4) heaters, set at 20, 30, or 40 Watts total as indicated

Flow Rate: (ROV 20-1854 opened) 2.0 lbm/hr

Tank Pressures: Use the backpressure control subsystem to maintain the selected ullage pressure (16.5 – 17 psia) 

within a tolerance of ± 0.001 psia.

The engineering checkout testing will begin with a pressure of 16.5 – 17 psia. The tank will be 

allowed to pressurize to the nominal operating pressure of 45± 4 psia.

The 45 ± 4 psia nominal operating pressure will be maintained during Engineering tests.
56

Additional General Instructions and Comments: 

• External HEX plate heaters and / or the valve panel hot-water heat exchanger should be on at all times before flowing to 

the valve panel. The H2 flowing through the valve panel should never drop below 360°R.

• Generally, whenever the tank heaters are on the ATVS pump should be on at the “low” setting. Note that this does not 

constitute “activation” of the ATVS, since the flow control valve has not been opened and there is no thermal extraction. 

The pump is turned on when the heaters are on to avoid non-homogenous stratification issues and to avoid “hot-spots” in 

the tank. The ATVS pump will NOT be turned on during the two Engineering Test Process high power vents (see 

attached timeline) even through the tank heaters are activated. Since ROV 20-1854 is opened and LH2 is being 

removed from the tank, this would cause the ATVS to extract heat from the tank and would lower the tank pressure to 

unacceptable levels. 

• There is no simulated launch during the Engineering Tests. The simulated launch from Section 5.3 of the Test Plan is 

now part of the 30-day mission simulation.

• There is no “without cold walls” boil-off test during the Engineering Tests. Data with hot walls and LN2 in the chamber 

will be taken as part of the 30-day mission simulation.

• Tank is to be drained, safed, and inerted after the Engineering Tests (prior to the 30-day mission simulation).

• Tank safing and inerting, boil-off stability requirements, etc. from Sections 5.1 – 5.11 can be found in the STUSTD 

LHSFS Test Plan. 

• Specific comments on the Engineering Test Process, Section 5.6 of the Test Plan, can be found on the pages following.

•

Engineering Checkout (Section 5.6) Comments and Instructions: 

A printout of the revised Engineering Test Process, Section 5.6 of the Test Plan, is found at the end of this document.

As a general rule, the following things are important, and the associated actions in the Engineering Test Process should be followed as 

closely as possible:

• Vent (“burn”, ROV 20-1854 open) times

• Heater power

• End of test-section tank pressure (45 psia). In other words, we always want to be at 45 psia before going on to the next test section

• Data rate switches

On the other hand, other things are not important and represent analytic approximations. They should not be followed, since they are 

essentially educated guesses at this time (in fact, the purpose of the Engineering Test is to refine these analytic models). Any controlling 

code generally should not key on the following parameters:

 

• Self-pressurization hold times

• Tank pressures after each item

• Tank fluid levels

STUSTD Engineering Test Process Description and Goals

The Engineering Test occurs the baseline cold wall boil-off test (Section 5.5 of the Test Plan).

Before the Engineering Tests, the tank has stabilized with LH2 at ~17 psia to the level of TT-04, which corresponds to a liquid level of 

about 40 inches.

For each of the following sections, the important parameter is to get the tank pressure back to 45 psia before going on to the 

next section. This will take a certain amount of time, depending upon, for example, whether or not the tank heaters are on. Generally, the 

amount of time it takes to get back to 45 psia and the fluid level in the tank are not important -- it is important to get back to 45 psia before 

going on to the next section. The times and tank levels are estimates only. The point of the entire Engineering Test Process is to refine 

the computational models to reflect the experimental results, not to try to make the experimental results match the models.

One exception is the burn (vent by opening ROV 20-1854) times called out in the Engineering Test Process. If the Process calls for a 

30 minute burn (vent) at a certain data acquisition rate, Boeing would like to keep these times as written. However, “hold” times can (and 

will) be flexible to allow the tank pressure enough time to reach 45 psia before the next section or segment of the current section.
57

SECTION A (Items 0 – 19 of the Engineering Test Process)

Brief Description

Essentially, we are trying to pressurize the tank to 45 psia (and learn a few things along the way). The tank heaters begin at 40 Watts, 

and then are reduced to 20 Watts in an attempt to slowly raise the tank pressure. The goal of this section is to have a final pressure of 45 

psia in the tank after a certain amount of time has elapsed. The important parameter is getting to the tank pressure of 45 psia, not the 

amount of time that it takes to get there (although, of course, we are interested in knowing how long it takes). The heaters will be operated 

in 6-hour blocks, until 45 psia is reached. Generally, the ATVS pump is activated to de-stratify the tank contents. When the tank reaches 

45 psia, stop Section A and go on to Section B.

The important thing is to get the tank to 45 psia. Period. The pressures in the P(psi) column of the Engineering Test Process 

spreadsheet are analytical estimates, and are by no means meant to be used for a controlling logic. We want to heat the tank until the 

pressure reaches 45 psia under saturated, mixed conditions, and the pressures after each heater cycle will be whatever it is.

 Item #17 says that we should run the tank heaters for 106 minutes – then the tank will be at 45 psia. Again, this is an estimate. It may 

reach 45 psia before or after this time. If we reach item #18 before the tank is at 45 psia, continue to heat the tank / run the pump until 45 

psia is reached. When the tank reaches 45 psia, stop Section A and go on to Section B.

The operation of the pump, tank heaters, pressure transducers, temperature transducers, data acquisition system, controller code, etc. 

will be verified in this section. We will also get some additional information about the tank heat leak and fluid stratification from the selfpressurization portion of the test.

Goals

We want to slowly bring the tank from a pressure of ~18 psia to 45 psia. The end pressure is important, not the time it takes to get 

there. We will make sure that the tank heaters, pump, temperature and pressure transducers, data acquisition system, etc. are all operating 

properly. Some tank heat-leak and stratification data will be taken by monitoring the self-pressurization rate.

Tank Initial State

284.1 lbm (less boiloff) LH2 at 16.5 – 17 psia. The fluid level is near TT-04.

What is Important?

Getting the tank pressure to 45 psia. Making sure that this is the mixed fluid pressure. The time it takes to get the tank to 45 psia is 

not important, but we’d like to run the heaters in 6-hour blocks.

Tank Final State

The 284.1 lbm (less boiloff) LH2 will be well mixed at 45 psia at a fluid level slightly above TT-01.

SECTIONS B & C (Items 20 – 69 of the Engineering Test Process)

Brief Description

 

This purpose of these sections is to verify the remote operation of ROV 20-1854, to adjust the metering valve in the GH2 vent line 

(upstream of TT-31 and PT-01), and to provide some preliminary results for the LHSFS PTVS thermal extraction rate and thermodynamic 

performance. The metering valve needs to be adjusted to ensure that we’re flowing 2.0 lbm/hr GH2 when ROV 20-1854 is opened. That 

way, we can determine key STUSTD parameters later on at the appropriate flow rate. Also, the 2.0 lbm/hr flow rate is used for the 30-day 

mission simulation.

TRANSIENT BEHAVIOR:

The initial flow rate may be < 2 lbm/hr. As such, we should wait several (5) minutes before adjusting the metering valve to a value of 

2 lbm/hr. Boeing-HB would like to measure and record the transient mass flow rate as a function of time. 

Burn (vent by opening ROV 20-1854) times are given in the Test Process. It’s important that we try to stick with these times. What is

flexible, on the other hand, are the hold times dictated by items 29, 39, 49, 59, and 69. Again, the important thing is to get the ullage 

pressure back up to 45 psia after items 29, 39, 49, 59, and 69, not the hold time it takes to get there. Run the pump on low for 5 

minutes during the unvented hold times (probably towards the end, when the pressure is approaching 45 psia) to mix the fluid.

Goals

Primarily, to adjust the metering valve so that we are flowing 2.0 lbm/hr GH2 when ROV 20-1854 is opened. We’ll also collect some 

very good data on the PTVS system performance.
58

Tank Initial State

284.1 lbm (less boiloff) LH2 at 45 psia. Fluid level is slightly above TT-01.

What is Important?

It is important that the metering valve is adjusted to deliver 2.0 lbm/hr of GH2 when ROV 20-1854 is opened and transients have died 

down. It is also important that we try to stick to the burn times of these sections, although these can be flexible if the need arises (i.e., lack 

of test time, etc.). The given hold times are not important (in fact, probably are not entirely correct), but the tank pressure at the end of the 

hold time is important. Any “burn” controlling logic should use the burn times indicated in these sections. Any “hold” controlling logic 

should not be hold time-oriented, but should be pressure-oriented.

Tank Final State

~279 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-01 and TT-02.

SECTIONS D & E (Items 70 – 137 of the Engineering Test Process)

Brief Description

These sections are used to determine the heat extraction capabilities of the PTVS and ATVS for different burn times and different 

pump speeds with a near-full tank. Recall that the ATVS is “activated” by opening ROV 20-1854 with the pump on. After each burn that 

the PTVS or ATVS has been used (30 minutes in Section D, and one hour in Section E) the tank will be allowed to self-pressurize (no tank 

heaters) back to 45 psia. The self-pressurization hold time indicated in items 81, 93, 103, 115, 127, and 137 is an estimate. The important 

thing is to get the pressure back up to 45 psia after the actions preceding the hold. Run the pump on low for at least 5 minutes during 

the unvented hold times (probably towards the end, when the pressure is approaching 45 psia) to mix the fluid.

 

Goals

We are trying to determine the heat extraction capabilities of the passive and active thermal venting systems for different burn times 

with a near-full tank. 

Tank Initial State

~279 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-01 and TT-02.

What is Important?

It is important that we try to stick to the burn times of these sections, although these can be flexible if the need arises (i.e., lack of test 

time, etc.). The given hold times are not important (in fact, probably are not entirely correct), but the tank pressure at the end of the hold 

time is important -- that is, we want to be back at 45 psia after each hold, before the next burn. Any “burn” controlling logic should use the 

burn times indicated in these sections. Any “hold” controlling logic should not be hold time-oriented, but should be pressure-oriented.

Tank Final State

~270 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-03 and TT-04.

SECTION F (Items 138 – 150 of the Engineering Test Process)

Brief Description

 

This is a “high power vent”. DO NOT RUN THE ATVS PUMP WHEN THE TANK HEATERS ARE TURNED ON. Essentially, it is 

intended to vent (“burn”) a great deal of fluid out of the tank at a near-constant pressure. Note that the tank heaters are on at 30 Watts.

High Power Vent:

 A reasonable amount of operator (i.e. human) observation and perhaps intervention will be required for this section. If the tank 

ullage pressure gets below 41 psia, this means that we are extracting too much energy, and the power to the tank heaters will have to be 

increased to 40 Watts. On the other hand, if the tank pressure gets above 49 psia, we are putting too much energy into the tank and the 

tank heater power will have to be decreased to 20 Watts. If the heater power has to be changed, this is fine -- as long as the time and the 

heater power is recorded.

We would like to vent/burn H2 for around 50 hours. It is somewhat important that we stick to this time (we are trying to have ~ 170 

lbm LH2 left in the tank after the burn/vent), but we can be flexible if necessary. 
59

After High Power Vent:

According to items 144 – 150, we are turning off the heater, continuing venting for approximately 2.5 hours, closing ROV 20-1854, 

and then allowing the tank to self-pressurize. This is all in an attempt to get the tank pressure back to 45 psia after the high power vent. 

The operator will need to exert some control at this point. That is: if we are above 45 psia after the 50 hour vent, turn off the heaters, 

vent until the tank is at ~44.5 psia, close ROV 20-1854, and then allow the tank to self-pressurize to 45 psia. If we are below 45 psia 

after the 50 hour vent, turn off the heaters, close ROV 20-1854, and allow the tank to self-pressurize to 45 psia. If we are well below 

45 psia (say, 42 psia) after the 50 hour vent, the tank heaters can be turned on for a little bit to raise the pressure to 45 psia more quickly. 

Run the pump on low for at least 5 minutes during the unvented hold times (probably towards the end, when the pressure is approaching 45 

psia) to mix the fluid.

The end result of this section is that there is much less fluid in the tank after the high power vent. This will allow us to test the 

performance of the passive and active thermal venting systems with the LADs only partially submerged. Since the LADs are no longer 

completely submerged, heat will begin to be extracted from the tank fluid by condensation effects in addition to convection effects. We are 

very interested in this data. 

Goals

Reduce the amount of fluid in the tank so that we can test the P/ATVS performance with the LADs only partially submerged.

Tank Initial State

~270 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-03 and TT-04.

What is Important?

The vent/burn time is important, since we are trying to end up with ~170 lbm of fluid in the tank at the end of this section. After the 

50 hour vent/burn, the operator should return the tank pressure to 45 psia using the methods described in the “Brief Description”. A return 

to 45 psia is important.

 

Tank Final State

~170 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-09 and TT-10.

SECTION G (Items 151 – 192 of the Engineering Test Process)

Brief Description

This section is used to determine the heat extraction capabilities of the PTVS and ATVS for different pump speeds with a partiallyfilled tank (LAD surface is exposed).

After each hour-long burn that the PTVS or ATVS has been used, the tank will be allowed to self-pressurize (no tank heaters) back to 

45 psia. The hour-long burn times are important. The self-pressurization hold times indicated in items 170, 182, and 192 are estimates. 

The important thing is to get the pressure back up to 45 psia after the actions preceding the hold. Run the pump on low for at least 5 

minutes during the self-pressurization hold times (probably towards the end, when the pressure is approaching 45 psia) to mix the fluid.

 

Goals

We are trying to determine the heat extraction capabilities of the passive and active thermal venting systems for different burn times 

with a partially-full tank. 

Tank Initial State

~170 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-09 and TT-10.

What is Important?

It is important that we try to stick to the burn times of these sections, although these can be flexible if the need arises (i.e., lack of test 

time, etc.). The hold times of items 162, 164, 166, and 168 are important. We want this data for these hold times. However, the selfpressurization hold times are not important (in fact, probably are not entirely correct), but the tank pressure at the end of the hold time is 

important -- that is, we want to be back at 45 psia after each hold, before the next burn.

Tank Final State

~159 lbm (less boiloff) LH2 at 45 psia. Fluid level is around TT-10.
60

SECTION H (Items 193 - 205 of the Engineering Test Process)

Brief Description

 

This is another “high power vent”. DO NOT RUN THE ATVS PUMP WHEN THE TANK HEATERS ARE TURNED ON. 

Essentially, it is intended to vent (“burn”) a great deal of fluid out of the tank at a near-constant pressure. Note that the tank heaters are on 

at 30 Watts.

The goal of this Section is to end up with ~90 lbm of LH2 left in the tank. 

High Power Vent:

 Much like Section F, a reasonable amount of operator (i.e. human) observation and perhaps intervention will be required for this 

section. If the tank ullage pressure gets below 41psia, this means that we are extracting too much energy, and the power to the tank heaters 

will have to be increased to 40 Watts. On the other hand, if the tank pressure gets above 49 psia, we are putting too much energy into the 

tank and the tank heater power will have to be decreased to 20 Watts (or even turned off). If the heater power has to be changed, this is 

fine -- as long as the time and the heater power is recorded.

We would like to vent/burn H2 for around 35 hours. It is of minor importance that we stick to this time (we are trying to have ~90 

lbm LH2 left in the tank after the burn/vent). 

After High Power Vent:

According to items 199 – 205, we are turning off the heater, continuing venting, closing ROV 20-1854, and then allowing the tank to 

self-pressurize. This is all in an attempt to get the tank pressure back to 45 psia after the high power vent. The operator will need to exert 

some control at this point. That is: if we are above 45 psia after the 35 hour vent, turn off the heaters, vent until the tank is at ~44.5 

psia, close ROV 20-1854, and then allow the tank to self-pressurize to 45 psia. If we are below 45 psia after the 35 hour vent, turn 

off the heaters, close ROV 20-1854, and allow the tank to self-pressurize to 45 psia. If we are well below 45 psia (say, 42 psia) after 

the 35 hour vent, the tank heaters can be turned on for a little bit to raise the pressure to 45 psia more quickly. Run the pump on low for at 

least 5 minutes during the unvented hold times (probably towards the end, when the pressure is approaching 45 psia) to mix the fluid.

The end result of this section is that there is much less fluid in the tank after the high power vent. This will allow us to test the 

performance of the passive and active thermal venting 

systems with the LADs mostly uncovered. Condensation effects will dominate the heat extraction capabilities of the TVS. We are very 

interested in this data. 

Goals

Reduce the amount of fluid in the tank so that we can test the P/ATVS performance with the LADs mostly uncovered.

Tank Initial State

~159 lbm (less boiloff) LH2 at 45 psia. Fluid level is around TT-10.

 

What is Important?

The vent/burn time is relatively important, since we are trying to end up with ~90 lbm of fluid in the tank at the end of this section. 

After the 35 hour vent/burn, the operator should return the tank pressure to 45 psia using the methods described in the “Brief Description”. 

A return to 45 psia is important.

 

Tank Final State

~90 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-13 and TT-14.

SECTION I (Items 206 – 246 of the Engineering Test Process)

Brief Description

This section is used to determine the heat extraction capabilities of the PTVS and ATVS for different pump speeds with a partiallyfilled tank (LAD surface is mostly exposed). 

After each hour-long burn that the PTVS or ATVS has been used, the tank will be allowed to self-pressurize (no tank heaters) back to 

45 psia. The hour-long burn times are important. The self-pressurization hold time indicated in items 225 and 237 is an estimate. The 

important thing is to get the pressure back up to 45 psia after the actions preceding the hold. Run the pump on low for at least 5 minutes 

during the self-pressurization hold times (probably towards the end, when the pressure is approaching 45 psia) to mix the fluid.

 
61

Goals

We are trying to determine the heat extraction capabilities of the passive and active thermal venting systems for different burn times 

with a partially-full tank (LADs mostly exposed). 

Tank Initial State

 

~90 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-13 and TT-14.

What is Important?

It is important that we try to stick to the burn times of these sections, although these can be flexible if the need arises (i.e., lack of test 

time, etc. However, the self-pressurization hold times are not important (in fact, probably are not entirely correct), but the tank pressure at 

the end of the hold time is important -- that is, we want to be back at 45 psia after each hold, before the next burn.

Tank Final State

~79 lbm (less boiloff) LH2 at 45 psia. Fluid level is between TT-13 and TT-14.
62
63
64
65
66

APPENDIX D—STUSTD 30-DAY MISSION SIMULATION (3/13/99)

Ullage conditioning vents on burns 29, 30, 54, 55, 82, 83, 121, 129, and 137.
67
68
69
70
55

REFERENCES

 1. “Solar Thermal Upper Stage Technology Demonstrator Program,” Final Report, NASA 

Marshall Space Flight Center Cooperative Agreement NCC-8-61, The Boeing Company, 

Huntington Beach, CA, October 1999.

 2. Cady, E.; and Olsen, A.: “Solar Thermal Upper Stage Technology Demonstrator Program,” 

AIAA Paper 96-3011, July 1996. 

 3. Olsen, A.D.; Cady, E.C.; Jenkins, D.S.; and Hastings, L.J.: “Solar Thermal Upper Stage Cryogen 

System Engineering Checkout Test,” AIAA Paper 99-2604, June 1999.

 4. “Test Implementation Plan (TIP),” Marshall Space Flight Center Test Position 300 Plan No. 

302-TIP-001, July 23, 1998. 

 5. Lin, C.S.; Van Dresar, N.T.; and Hasan, M.M.: “Pressure Control Analysis of Cryogenic 

Storage Systems,” AIAA Paper 91-2405, June 1991.

 6. FLOW-3D Computational Modeling Power for Scientists and Engineers, Technical Manual, 

Version 7.1, Flow Science, Inc., 1992.

 7. FLOW-3D User’s Manual, Version 9.0, Flow Science, Inc., Santa Fe, NM, 2005.

 8. Grayson, G.D.; and Lopez, A.: “Cryogenic Tank Modeling for the Saturn AS-203 Experiment,” 

AIAA Paper 2006-5258, March 9, 2006. 

 9. Hastings, L.J.; Flachbart, R.H.; Martin, J.J.; et al.: “Spray Bar Zero Gravity Vent System 

for On-Orbit Liquid Hydrogen Storage,” NASA/TM—2003–212926, Marshall Space Flight 

Center, Huntsville, AL, October 2003. 

10. Flachbart, R.H.; Hastings, L.J.; Hedayat, A.; et al.: “Testing of a Spray-Bar Thermodynamic 

Vent System in Liquid Nitrogen,” Paper C1-G-05, Cryogenic Engineering Conference and 

International Cryogenic Materials Conference, Keystone, CO, August 29–September 2, 2005.

71
56

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Solar Thermal Upper Stage Liquid Hydrogen Pressure Control 

Testing and Analytical Modeling

A.D. Olsen,* E.C. Cady,* D.S. Jenkins,* F.O. Chandler,* 

G.D. Grayson,* A. Lopez,* L.J. Hastings,** R.H. Flachbart, 

and K.W. Pedersen

George C. Marshall Space Flight Center

Huntsville, AL 35812

National Aeronautics and Space Administration

Washington, DC 20546–0001

Unclassified-Unlimited

Subject Category 20

Availability: NASA CASI (443–757–5802)

Prepared by the Propulsion Systems Department, Engineering Directorate

*The Boeing Company, Huntington Beach, CA **Alpha Technology, Inc., Huntsville, AL

M–1345

Technical Publication

NASA/TP—2012–217469

low-gravity cryogenic fluid management, liquid hydrogen thermodynamic vent system/liquid acquisition device 

thermal integration, on-orbit liquid hydrogen pressure control, use of liquid hydrogen boiloff for solar propulsion

01–10–2012

UU 88

NASA

U U U

The demonstration of a unique liquid hydrogen (LH2) storage and feed system concept for solar thermal 

upper stage was cooperatively accomplished by a Boeing/NASA Marshall Space Flight Center team. The 

strategy was to balance thermodynamic venting with the engine thrusting timeline during a representative 

30-day mission, thereby, assuring no vent losses. Using a 2 m3 (71 ft3) LH2 tank, proof-of-concept 

testing consisted of an engineering checkout followed by a 30-day mission simulation. The data were 

used to anchor a combination of standard analyses and computational fluid dynamics (CFD) modeling. 

Dependence on orbital testing has been incrementally reduced as CFD codes, combined with standard 

modeling, continue to be challenged with test data such as this.

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