Jason W. Hartwig

ZIN Technologies, Brook Park, Ohio 

John B. McQuillen and David J. Chato

Glenn Research Center, Cleveland, Ohio

Performance Gains of Propellant Management 

Devices for Liquid Hydrogen Depots

NASA/TM—2013-217834

February 2013

AIAA–2013–0368
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Jason W. Hartwig

ZIN Technologies, Brook Park, Ohio 

John B. McQuillen and David J. Chato

Glenn Research Center, Cleveland, Ohio

Performance Gains of Propellant Management 

Devices for Liquid Hydrogen Depots

NASA/TM—2013-217834

February 2013

AIAA–2013–0368

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

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NASA/TM—2013-217834 1

Performance Gains of Propellant Management 

Devices for Liquid Hydrogen Depots

Jason W. Hartwig

ZIN Technologies

Brook Park, Ohio 44142

John B. McQuillen and David J. Chato

National Aeronautics and Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Abstract

This paper presents background, experimental design, and preliminary experimental results for the 

liquid hydrogen bubble point tests conducted at the Cryogenic Components Cell 7 facility at the NASA 

Glenn Research Center in Cleveland, Ohio. The purpose of the test series was to investigate the 

parameters that affect liquid acquisition device (LAD) performance in a liquid hydrogen (LH2) propellant 

tank, to mitigate risk in the final design of the LAD for the Cryogenic Propellant Storage and Transfer 

Technology Demonstration Mission, and to provide insight into optimal LAD operation for future LH2

depots. Preliminary test results show an increase in performance and screen retention over the low 

reference LH2 bubble point value for a 325×2300 screen in three separate ways, thus improving 

fundamental LH2 LAD performance. By using a finer mesh screen, operating at a colder liquid 

temperature, and pressurizing with a noncondensible pressurant gas, a significant increase in margin is 

achieved in bubble point pressure for LH2 screen channel LADs.

Nomenclature

Dp pore diameter (μm)

DPT differential pressure transducer 

FTS flow through screen

g gravitational acceleration (m/s2

)

GHe gaseous helium 

GH2 gaseous hydrogen 

GN2 gaseous nitrogen 

GRC NASA Glenn Research Center

IPA isopropyl alcohol

LAD liquid acquisition device 

LCH4 liquid methane 

LEO low Earth orbit

LL liquid level above LAD screen (m)

LN2 liquid nitrogen 

LOX liquid oxygen

L/V liquid/vapor

MAWP maximum allowable working pressure

MMH monomethyl hydrazine 

N2O4 nitrous tetraoxide

NBP normal boiling point

OMS Orbital Maneuvering System

PID proportional-integral-derivate
NASA/TM—2013-217834 2

PMD propellant management devices 

RCS Reaction Control System 

RD receiver dewar 

SD silicon diode

STS Shuttle Space Transportation Systems

TRL technology readiness level 

VJ vacuum jacketed 

ΔPBP bubble point pressure (Pa)

ΔPdynamic pressure loss for mass accumulation along channel length (Pa)

ΔPFTS pressure loss for flow through screen (Pa)

ΔPfriction friction pressure loss for flow along channel (Pa)

ΔPhydrostatic hydrostatic pressure (Pa)

ΔPother other sources of pressure fluctuations and differences (Pa)

γ surface tension (mN/m)

ρLH2 liquid hydrogen density (kg/m3

)

θc contact angle (degrees)

1.0 Introduction

In recent years, there has been a decline in the proposed usage of toxic propellants, such as nitrous 

tetraoxide (N2O4) and monomethyl hydrazine (MMH). Safety and environmental concerns over the use of 

these storable propellants (propellants that are liquids at room temperature) have led to the ongoing 

examination of non-toxic cryogenic propellants as an ideal fluid for on-orbit propulsion systems for future 

human and robotic exploration missions. Among the available choices, liquid oxygen/liquid hydrogen 

(LOX/LH2) is the propellant combination of choice owing to an unmatched level of performance relative 

to other systems. In addition, LOX/LH2 has proven flight heritage in launch systems such as Saturn V 

(S4, S4B, S2) and Shuttle Space Transportation Systems (STS), as well as short duration upper stages 

(Centaur, Delta IV). LOX/LH2 is also an attractive candidate for proposed future in-space engines like the

J-2X, and low Earth orbit (LEO) fuel depots, both of which are required to enable long duration space 

missions (Kutter et al. 2008, Kutter 2010, and McLean et al. 2011). However in the optimal design of any 

spacecraft, high performance is not the only consideration. Long term missions will require either a 

reconsideration of propellant choice or a refinement of existing LOX/LH2 technology.

There are challenging aspects when working with cryogenic liquid hydrogen due to its 

thermophysical properties. The low liquid density and normal boiling point (NBP) makes LH2 difficult to 

package, requiring larger storage tanks. Hydrogen’s critical temperature and pressure limits the conditions 

under which hydrogen exists as a liquid. LH2 is particularly susceptible to parasitic heat leak due to its 

low NBP resulting in propellant loss due to boil off. Finally, it is difficult to transfer and maintain single 

phase liquid flow due to its low viscosity, NBP, and surface tension, making vapor ingestion into the 

transfer line highly probable.

Gravity affects many fluidic processes, such as the separation of liquid and vapor within a propellant 

tank. In 1-g, density controls the location of the liquid/vapor (L/V) interface, as the heavier liquid settles 

to the bottom and lighter vapor rises to the top. In low Bond number (ratio of gravitational to surface 

tension forces) environments, surface tension controls the location of the L/V interface because liquid 

tends towards the walls and vapor tends towards the center of the tank. A robust, flexible system capable 

of delivering single phase liquid flow in varying gravitational and thermal environments is clearly 

required to meet the demands of both future in-space engines as well as the proposed fuel depots.
NASA/TM—2013-217834 3

2.0 Screen Channel LADs

When transferring propellant in microgravity either by pump or pressure fed applications, it is 

necessary to transport only single phase liquid to an engine to ensure efficient thermal conditioning of the 

engine and reliable operation after ignition. Depot tanks will also require high liquid fill fractions due to 

the projected cost of launching and storing propellant in LEO. In Earth’s 1-g field or in milli-g 

environments, propellant transfer is fairly simple. Single phase liquid flow is ensured by an antivortex 

baffle to prevent gas ingestion into the tank outlet. Fluid is transferred by pressurizing the tank and 

opening a valve at the bottom of the tank. In microgravity however, liquid may not sufficiently cover the 

tank outlet making single phase liquid expulsion a challenge.

The need for a special device to separate liquid and gas in cryogenic propellant tanks was required to 

perform engine restarts in a low-g environment (Radcliffe and Transue 1961). Depending on the 

gravitational environment and mission requirements, a variety of propellant management devices (PMD) 

may be required to favorably position the liquid and/or ullage in the tank. One type of PMD, a screen 

channel LAD, uses surface tension forces to acquire and maintain single phase flow to the transfer line.

These LADs are designed and manufactured in a variety of styles, sizes, and geometries. In flight-like 

systems, screen channels, or gallery arms, tend to closely follow the contour of the propellant tank wall 

(shown in Figure 1) and can have different cross section geometries (typically a triangular or rectangular 

shape). The channel side that faces the wall has openings covered with a tightly woven fine mesh screen. 

The screen has micron sized pores which are used to wick liquid into the channel, prevent pores from 

drying out during tank drain, and also act as a barrier to vapor ingestion. As liquid is withdrawn from the 

tank and vapor approaches the screen, surface tension forces at the screen generate a localized area of 

high pressure differential that blocks vapor entrance into the channel, but allow the liquid to flow freely. 

Liquid is wicked along the screen and prevents the pores from drying out if they come into contact with 

vapor. Full communication LADs (i.e., LAD that extends the entire height of the tank wall to maintain 

communication with the liquid at all time) have demonstrated flight heritage in storable systems such as 

the STS Reaction Control System (RCS) and Orbital Maneuvering System (OMS) (Fester et al. 1975 and 

Peterson and Uney 1978) and the design of LADs for storables is well understood (Schweickert 1981 and 

Anglim 1981). Although screen channel LADs have been used in a small scale liquid helium experiment 

in microgravity (Anderson 1989), they have not been used with LH2 in low gravity.

Vane and sponge type PMDs can provide vapor-free liquid under a much simpler design and 

implementation (Jaekle 1991, 1993, 1997) than screen channel LADs. Both types are advantageous in 

systems which experience low-g levels and very low supply demands. However, sponges and vanes may 

not meet the high flow rate requirements of future propellant depots; sponges simply do not scale with the 

projected size of depot tanks, and vanes may not be able to supply large flow rates relative to screen 

channel LADs. Thus, screen channel LADs may be the most extensible and flexible cryogenic propellant 

management device for variable vehicle acceleration and direction, and provide a wide range of flow rate 

regimes to supply single phase liquid for both engine restart and continuous engine operation, as well as 

fluid transfer for fuel cells, life support, and the proposed fuel depots. 

For flight systems, screen channel LAD usage is broken into two categories (Burge and Blackmon 

1973 and Burge et al. 1973). Start baskets, traps, and start tanks are considered small, liquid acquisition 

devices that confine sufficient liquid to start engines until the relatively large accelerations can adequately 

reorient the liquid for the large flow rates required for engine operation. Meanwhile full communication 

channels, distributers, or tank liners are used in systems with small accelerations and flows rates.

The choice of screen for the LAD is dictated by the gravitational environment and desired maximum 

flow rate, which can be estimated through knowledge of the bubble point pressure (Section 3.0). Screens 

are characterized by the screen weave, which refers to the number of wires per inch in each direction, and 

the specific weave pattern used during manufacturing. For example, the 325×2300 Dutch Twill mesh 

screen displayed in Figure 2 has 325 warp wires and 2300 shute wires per square inch of the screen. For a 

Dutch Twill, each shute wire passes under two warp wires before traveling over the next two warp wires. 

Fine mesh Dutch Twill screens create very small pores and provide a tortuous path and resistance against 
NASA/TM—2013-217834 4

Figure 1.—Full Communication Screen

Channel LAD.

Figure 2.—Scanning electron microscope image of a 

325×2300 screen sample.

gas ingestion, which make them popular choices for low surface tension cryogenic liquids. However they 

tend to generate large hydraulic pressure losses during outflow and may become clogged by particulate 

matter. Therefore, no single screen mesh or style exists to optimize the PMD for all propellant tanks and 

missions. 

3.0 The Bubble Point

The primary performance parameter characterizing LADs is the bubble point, which is defined as the 

differential pressure required to overcome the liquid surface tension force at the screen pore. Physical 

parameters that affect the bubble point include the geometry and size of the screen pore, the contact angle 

between screen pore and fluid, and the fluid surface tension. Hartwig and Mann (2013a) derived a bubble 

point model based on room temperature bubble point tests:

4 cos γ θ ∆ = C

BP

P

P

D (1)

where γ is the surface tension of the fluid (mN/m), θC is the contact angle between liquid propellant and 

solid screen pore, and DP is the average pore diameter (μm). Based on previously reported cryogenic 

bubble points from Hartwig and McQuillen (2011 and 2012a), the model is only valid for normally 

saturated liquids. Hartwig and Mann (2013b) showed that Equation (1) simplifies to

4γ ∆ = BP

P

P

D (2)

for cryogenic liquids. Effective pore diameters can be determined from room temperature bubble point tests 

conducted in high surface tension liquids (relative to cryogenic liquids) such as isopropyl alcohol (IPA).

To prevent vapor ingestion into the channel during outflow, the total pressure loss in the LAD system 

must be less than the bubble point pressure:

∆ <∆ P P total BP (3)

where the total pressure loss is expressed as:

∆ =∆ +∆ +∆ +∆ +∆ PP PP P P total hydrostatic FTS friction dynamic other (4)
NASA/TM—2013-217834 5

where ΔPhydrostatic is the hydrostatic pressure within the channel, ΔPFTS is the pressure drop across the 

screen, ΔPfriction is the frictional loss down the LAD channel, ΔPdynamic is the dynamic pressure drop due to 

inflow into the channel, and ΔPother is the pressure loss contribution due to vibrations, propellant sloshing, 

and/or transients (Van Dyke 1998). For the orbital conditions of LEO, pressure losses due to flow through 

screen (FTS) and transients become leading order effects. For a given LAD screen, cryogenic liquid, 

operating range of liquid temperature, pressure, and mass flow, the bubble point thus serves as an upper 

limit on the total allowable pressure loss for that system. Since the bubble point is easily measurable and 

readily translated into a deliverable flow rate to an engine or depot receiver tank, it serves as the primary 

performance parameter for characterizing any screen channel LAD.

4.0 Test Objectives

Issues that arise in LH2 cryogenic fluid management can be generically categorized into storage and 

transfer and are being addressed through a suite of ground testing at NASA as part of its technology 

maturation program. To address transfer issues, both a small scale component level bubble point test and 

a full scale screen channel LAD test are being performed at NASA Glenn Research Center (GRC). The 

preliminary bubble point test results are reported here. The purpose of this test was threefold: to give 

future mission designers insight into the influential factors that govern LAD behavior, to mitigate risk in 

flight LAD design/operation, and to demonstrate improved screen retention using three separate methods.

Doing so will raise the technology readiness level (TRL) of LH2 LADs for all future LH2 fueled 

propulsion systems and proposed cryogenic depots, and provide critical data for the mission designers.

5.0 Performance Gains

The current state of LH2 screen channel LADs technology is ascertained through detailed review of 

the literature. As was mentioned earlier, screen channel LADs have proven flight heritage with storable 

propellants, but, only ground testing has been conducted for cryogenic LOX and LH2 using these fine 

mesh screens. Previously reported bubble points for the most popular 325×2300 mesh screen are plotted 

in Figure 3, which includes data for LH2, liquid nitrogen (LN2), LOX, liquid methane (LCH4) and IPA 

from Paynter (1973), Burge and Blackmon (1973), Cady (1973, 1975, 1977), Chato and Kudlac (2002), 

Kudlac and Jurns (2006), and Jurns et al. (2007), respectively. The black line is the predicted values using

Equation (2) and a pore diameter specified from Hartwig and Mann (2013a). Figure 4 plots sparse 

historical data for the finer 450×2750 mesh screen. Chronologically, a considerable amount of LH2 LAD 

ground testing was performed in the early 1970s as part of the STS auxiliary and Centaur upper stage 

propulsion systems. LH2 LAD work progressed into the 1990s as part of technology development for 

propellant transfer vehicles and the Space Station Freedom. In the 2000s, a single bubble point 

measurement in LH2 was conducted at GRC by Chato and Kudlac (2002). 

The resultant NBP LH2 bubble point value at atmospheric pressure for the 325×2300 screen is only 

520 Pa. There was reported difficulty in even resealing the coarser 200×1400 screen after bubble 

breakthrough from Chato and Kudlac (2002). Since the bubble point serves as the upper limit on 

allowable pressure loss in a flight LAD, this low value severely limits the operating flow rates and liquid 

temperature. Over time, the liquid in the tank absorbs heat from the environment, the surface tension and 

thus bubble point will only decrease with increasing liquid temperature. Given this low baseline reference 

level, it was necessary to seek ways to improve the fundamental performance of LH2 LADs.

There are three ways to increase margin in the bubble point over the baseline 325×2300 value. First, 

choosing a finer mesh screen can theoretically decrease the effective pore diameter. While the 325×2300 

has been extensively used in ground tests, data is available in the literature for a finer 450×2750 mesh 

screen and the data is plotted in Figure 4. Burge et al. (1972) and Cady (1973) report LH2 data while 

Castle (1972) reports LN2 and IPA data for this finer mesh screen. The black curve in Figure 4 is based on 

room temperature bubble point tests from Hartwig and Mann (2013a). As shown, for both room 

temperature and cryogenic fluids, higher bubble points can be obtained using the 450 screen.
NASA/TM—2013-217834 6

Figure 3.—Previously reported LH2 bubble point values for the 

325×2300 Dutch Twill screen with helium pressurant.

Figure 4.—Previously reported bubble point values for the 450×2750 

Dutch Twill screen with helium pressurant.

Secondly, one can improve bubble point margin by obtaining a colder liquid temperature, since 

surface tension, and thus bubble point, is an inverse function of temperature. Third, the margin can be 

increased by subcooling the liquid with a noncondensible pressurant such as gaseous helium (GHe). 

Previous tests with LOX and LCH4 demonstrated an increase in margin, but this effect must be confirmed 

with LH2. With such a small baseline reference value, any excess bubble point margin counts when 

working with LH2 LADs.

6.0 Experimental Setup

Testing was performed at the Cryogenic Components Lab 7 (CCL-7) at GRC in Cleveland, Ohio. The 

same samples from Hartwig and Mann (2012a and b) were tested here. Figure 5 shows the LAD screen 

0 5 10 15 20 25

Paynter 73

Burge & Blackmon 73

Cady 73

Cady 75

Cady 77

Chato and Kudlac 02

Kudlac and Jurns 06

Jurns 07

Prediction

0

1000

2000

3000

4000

5000

6000

7000

Bubble Point [Pa]

Surface Tension [mN/m]

LH2

LN2

LOX

LCH

4

IPA

Freon-112

325x2300

0

1000

2000

3000

4000

5000

6000

7000

8000

0 5 10 15 20 25

Castle 72

Burge et. al 72

Cady 73

Prediction

∆

P

BP [Pa]

γ Cos θ

C

 [mN/m]

LH2

LN2

IPA

450x2750
NASA/TM—2013-217834 7

and cup assembly. Each 6.35 cm (2.5 in.) screen sample was welded onto a custom built flange with a 

cover ring and fillet weld to prevent leakage along the screen edges. Testing was always performed in an 

inverted configuration, with liquid above the screen and vapor or gas below the screen. Therefore, the 

purpose of the cup was to create the L/V interface within the screen pores by pressurizing the screen from 

below. Gas was distributed uniformly inside the cup by a custom built cross shaped manifold. Threaded 

rods supported an aluminized mirror to reflect an image of the top of the LAD screen as shown in 

Figure 5. The rods were also used to mount a fiber optic light source.

The LAD screen/cup was mounted inside a receiver dewar (RD) as shown in Figure 6. The RD 

housed the LH2 on top of the LAD screen. The RD was equipped with seven ports for liquid fill and drain, 

vent (x2), power, light source, and instrumentation for both RD and LAD cup, and three side ports for 

pressurant gas fill, drain, and differential pressure transducer (DPT) measurements. The image of the 

LAD screen was reflected off the mirror and recorded through the side portal located halfway up the 

dewar wall. A camera mounted to the viewport viewed and recorded the LAD screen in real time. Images 

were time stamped to synchronize with all other data.

The LH2 (or LN2) supply line was first routed through a supply dewar, as shown in Figure 7. The 

purpose of the supply dewar was to condition the liquid cryogen for the RD. For these tests, the dewar

itself was never filled with liquid; it was simply pumped down to 2 psia to minimize heat leak into the 

liquid feed line. Vacuum jacketed (VJ) lines connected the dewars in series. 

Several facility modifications to the experimental hardware from Jurns et al. (2007) were required 

before flowing LH2 at Cell 7, and to ensure higher accuracy results over previous bubble point tests. All 

pressure control and regulation valves were changed to allow complete remote operation of Cell 7. 

Second, the flow system was modified to remotely flow three different pressurization gases beneath the 

screen: GHe, gaseous nitrogen (GN2), and gaseous hydrogen (GH2). Third, the gas flow control system 

was modified to allow slower ramp rates in pressurization beneath the screen. Fourth, to improve upon 

accuracy of measurements from Hartwig and McQuillen (2012a), dual range high accuracy DPTs were 

installed to measure the differential pressure across the screen as shown in Figure 7. Fifth, smaller 

diameter, 3 mm (1/8 in.) sense lines were used to improve pressure signal response time. Sixth, extra 

silicon diodes were installed to resolve temperature close to the screen. Finally, a new data acquisition 

system was installed to resolve difficulties in synchronizing the pressure, temperature, and flow data with 

the video data.

Figure 5.—LAD cup and diode rake.
NASA/TM—2013-217834 8

Figure 6.—Receiver dewar, temperature instrumentation.

Figure 7.—Cell 7 Liquid Hydrogen Flow System.
NASA/TM—2013-217834 9

LH2 was filled offsite in portable 250 gal dewars and was connected to the flow system through a 

flexible VJ line. LN2 was connected through the same VJ line via a roadable 500 gal dewar, and was used 

to perform all pretest check outs and cold shocking of the hardware and to collect bubble point data. GHe 

was available from a portable tuber trailer while GH2 and GN2 were available from a high pressure Kbottle. Gas pressure and flow rate into the LAD cup were controlled by a set of three low flow control 

valves.

Liquid pressure inside the RD was primarily controlled using a proportional-integral-derivate (PID) 

loop. To increase pressure, gas was added to the ullage space in the RD. The pressure would also increase

due to heat leak into the RD. The vent valve was opened to restore pressure back to ambient. To reduce 

pressures below atmospheric, Cell 7 ejectors were used to pump down on the liquid. Temperature of the 

LH2 was primarily controlled by initial conditioning of the liquid inside the roadable dewar. Cell 7 air 

ejectors were used to reduce the liquid temperature by reducing pressure in the ullage space. Parasitic heat 

leak into the RD provided more than sufficient thermal energy to increase liquid temperature. The 

pressure of the incoming pressurant gas was controlled by the set of flow control valves. The pressurant 

gas temperature was measured using a silicon diode (SD) located inside the cup (SD3) and controlled by 

heaters inside of the LAD cup, in a PID control loop.

Figure 5 and Figure 6 show the locations of temperature measurements. Diodes were mounted 

directly to the liquid (SD1) and gas (SD2) sides of the screen. Two diodes were mounted inside the LAD 

cup to measure the temperature of the incoming gas (SD3 and SD4). SD6-9 were mounted in a staircase 

as depicted in Figure 5 to resolve the temperature gradient in the bulk liquid above the LAD screen. 

SD10-14 provided temperature and liquid level inside the RD. Pressure was measured inside the LAD cup 

and in the RD. The differential pressure across the LAD screen was measured using a 0 to 1250 Pa (0 to 5 

in H2O) and 0 to 2500 Pa (0 to 10 in H2O) high accuracy DPT while a 0 to 7500 Pa (0 to 30 in H2O) DPT 

was used for LN2 bubble point tests. Gas flow rates into the cup were also measured.

Diodes measured temperature to within ± 0.1K. The pressure at the LAD screen was measured to 

within 2300 Pa (0.333 psia). The raw DPT reading across the screen was accurate to within 0.32 Pa 

(0.0013 in H2O), and the total uncertainty in reported bubble point pressures is estimated to be 2.3 Pa 

(0.093 in H2O), which is less than 3.6 percent at the lowest measured pressures.

7.0 Experimental Methodology

The methodology for conducting the LH2 bubble point test was as follows. Prior to bubble point tests, 

all supply lines and both dewars were purged with GHe. A GHe flow was established across the LAD 

screen to prevent flooding of the cup during LH2 fill. Next, LH2 was transferred to condition the facility 

plumbing, RD, and eventually fill the RD to the desired liquid level. Once the RD was filled, the liquid 

was conditioned to the desired temperature and pressure at the LAD screen using methods discussed 

earlier. 

To conduct a bubble point test, the screen was resealed by slowly decreasing pressurant gas flow into 

the cup. The desired pressurant gas type and pressurant gas temperature was then selected. When 

switching over to autogenous pressurization, the vapor was allowed to flow through the lines, LAD cup, 

and screen for up to 10 minutes to purge other gases from inside the network prior to controlled bubble 

point tests. Then, the pressure underneath the LAD screen was slowly increased using the flow control 

valves. Eventually a bubble would break through the wetted screen, breaking the screen down, as 

indicated both by visual observation and an increase in the DPT signal. The bubble point was taken as the 

moment when a visible bubble penetrated the liquid laden screen. The time stamp at breakthrough was 

noted and synchronized to the data file to extract the corrected bubble point pressure and any other 

relevant parameters. Bubble point measurements were then repeated at similar pressures and/or 

temperatures for repeatability and consistency before moving on to the next data point. Note that bubbles 

broke through various locations on the screen, indicating that there was no preferential breakdown spot 

due to screen defects.
NASA/TM—2013-217834 10

Data reduction for the LH2 bubble point tests was straightforward. Details are presented in Hartwig 

and McQuillen (2012a). To compare with theoretical predictions, all experimental bubble point values 

were corrected for head pressure:

( ) exp LH2 ∆P T P DPT gLL BP , = −ρ (5)

where LL is the liquid level above the screen in the RD as determined by the vertical silicon diode rake.

Thermodynamic properties such as liquid density and surface tension were computed using REFPROP. 

8.0 Test Matrix

The overall test plan is outlined on the left hand side in Figure 8. Three fine mesh screen channel 

LAD samples (325×2300, 450×2750, 510×3600) were tested in two cryogenic liquids (LH2 and LN2), 

using two pressurization schemes (autogenous GH2/LH2 and GN2/LN2 and noncondensible GHe/LH2 and 

GHe/LN2) at several different pressurant gas temperatures. Bubble point data was collected across the 

widest possible range of liquid temperatures and pressures within facility limitations. The lower limit of 

data collection was limited by the Cell 7 ejector’s ability to pull vacuum on the RD to chill the liquid to 

the saturation temperature at 31 kPa (4.5 psia). The upper limit of 155 kPa (22.5 psia) was dictated by the 

maximum allowable working pressure (MAWP) of the RD. This range allowed independent examination 

of both liquid temperature and pressure effects on bubble point using two different pressurant gases, as 

both saturated and subcooled liquid states were achievable. This is indicated in the blue shaded region in 

Figure 8. The pressure range corresponds to a range in LH2 temperatures of 16.7 < T < 21.1K, which 

represents typical thermal conditions within a low pressure LH2 propellant tank. More bubble point data 

was collected during this test series than the sum total of all previous LH2 LADs tests to date.

Figure 8.—LAD Bubble Point Test Plan.
NASA/TM—2013-217834 11

The test matrix for the 450×2750 screen in LH2 is outlined in Table I. Stable liquid temperature 

during testing permitted an isothermal marching scheme in data collection, as shown. For temperature 

variation, the blue shaded region in Figure 8 was divided into six equally spaced isothermal slices across 

the temperature range. For pressure variation, controlled breakthroughs were spaced in 6.9 kPa (1.0 psi)

intervals from the saturation pressure up to the MAWP. Data was collected across the entire range of 

liquid temperatures in Table I using both GHe and GH2 as pressurant gases. The pressurant gas feed line 

was immersed within the liquid inside the RD and the residence time of the gas inside the LAD cup was 

long enough such that the baseline gas temperature was approximately the liquid temperature. To collect 

data at elevated gas temperatures, the heaters inside the LAD cup were activated and the entire grid in 

Table I was repeated for two more pressurant gas temperatures. The entire process was repeated for the 

other two meshes. Parametric bubble point testing for both LH2 and LN2 required approximately 3 weeks 

of continuous testing with a total test time of approximately four months for cold shocks, check outs, 

testing, and screen change out.

TABLE I.—TEST MATRIX: 450×2750

LH2

PSAT,

kPa (psia)

T,

K

PSCREEN,

psia

29.3 (4.25) 16.7 4.25 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

41.4 (6) 17.5 ----- -- 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

55.2 (8) 18.4 ----- -- -- -- 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

75.8 (11) 19.3 ----- -- -- -- -- -- -- -- 12 13 14 15 16 17 18 19 20 21 22

101.3 (14.5) 20.3 ----- -- -- -- -- -- -- -- -- -- 14.5 15 16 17 18 19 20 21 22

131+ (19+) 21.1+ ----- -- -- -- -- -- -- -- -- -- ----- -- -- -- -- 19 20 21 22

9.0 Preliminary Results and Discussion

Figure 9 is a screen shot of an ideal bubble point test. The raw DPT signal across the screen, the 

liquid screen side temperature (SD1), bulk liquid temperature (SD11), and RD ullage pressure during a

run of six successive controlled breakthroughs for a 450×2750 using GHe in LH2 are shown. 

Breakthrough pressures occurred at roughly 90 to 95 percent of the maximum DPT signal. As shown, 

screen side and bulk liquid temperature gradually rose as ullage pressure was increased. Good signal to 

noise (SNR) ratio was also achieved, even in warmer liquid temperatures as the DPT signals decreased

due to decreased surface tension of the liquid.

The low baseline bubble point value for LH2 using a 325×2300 screen is only 520 Pa. The easiest way 

to improve upon this low reference value is to use a finer mesh. Figure 10 plots bubble point pressures taken 

at saturated or near saturated liquid states using GHe for the 325×2300 and 450×2750 screens. The solid 

lines are predictions based on Equation (2). “Near saturation” is defined as (PSCREEN – PSAT) ≤ 2.0 psi.

Higher bubble points are obtained using the finer 450 mesh screen, with an increase in 27 percent margin 

over the baseline 325 screen. While the model qualitatively agrees with the data, it under predicts the data 

(model predicts lower bubble point pressure) for both meshes, even for saturated liquid states where the 

theory is in best agreement. Since the data is taken at 20K and the curves are based on room temperature 

bubble point tests, the pore diameter is clearly temperature dependent. Comparing the difference between 

data and model, the 325×2300 pore diameter achieves a 7 percent gain in bubble point and a 9 percent gain 

for the 450×2750 screen at 20K over the model curve using the room temperature pore diameter. Figure 10

also implies that the 450×2750 screen may deliver a higher flow rate from a flight propellant tank over the 

325×2300 screen.
NASA/TM—2013-217834 12

Figure 9.—Screenshot of an ideal test run during 450×2750 GHe/LH2 testing.

The second way to improve upon the low reference bubble point value is to directly increase the 

surface tension of the liquid. Figure 11 plots LH2 bubble point data as a function of the liquid side screen 

temperature, SD1. Again, only bubble points taken in saturated or near saturated liquid states are plotted 

to permit direct comparison between screen meshes. As shown for both meshes, higher bubble points are 

always achieved in colder liquid temperatures, since surface tension is an inverse function of temperature. 

Model and data are again in qualitative agreement when surface tension is based on SD1, but not the bulk 

liquid temperature SD11. Screen pores could be shrinking at colder liquid temperature due to the thermal 

coefficient of contraction of the warp and shute wires. Because the gain is different for each screen, 

geometrical considerations of the actual pore structure must also be taken into account. Nonetheless, extra 

margin is obtained for both meshes at cryogenic hydrogen temperatures.

Finally, the third method to improve upon the low reference bubble point value is to indirectly raise 

the surface tension of the liquid. Figure 12 plots all of the 325×2300 bubble point data obtained with both 

GHe and GH2 as pressurant gases. The solid line is the model prediction using a room temperature pore 

diameter. Several trends are obvious. First, GHe outperforms GH2 across the entire range of LH2

temperatures, and the disparity between gases is far outside the uncertainty in the data. Compared to 

trends with LOX and LCH4 from Hartwig et al. (2012) and Hartwig and McQuillen (2012b) this disparity 

in performance between pressurization gases is greatest at the coldest LH2 temperatures. Second, while 

data and theory are in qualitative agreement, the room temperature pore diameter model underpredicts 

GHe data and overpredicts the GH2 data. Therefore, noncondensible pressurization increases margin for 

the bubble point, but autogenous pressurization decreases margin for the bubble point, relative to the 

model predicted curves. Difference in performance between the two pressurization schemes were 

explained previously in Hartwig et al. (2012) and Hartwig and McQuillen (2012b). In summary, during 

autogenous pressurization, vapor condenses across the screen into the liquid, warming the liquid at the 

screen, reducing the surface tension and thus the bubble point. Meanwhile pressurization with the 

noncondensible GHe causes liquid to evaporate away from the screen cooling the liquid temperature at 

the screen, increasing surface tension and thus the bubble point.
NASA/TM—2013-217834 13

Figure 10.—Liquid hydrogen bubble point as a function of screen mesh. 

Solid lines are model predictions based on Equation (1) using room 

temperature pore diameters.

Figure 11.—Liquid hydrogen bubble point as a function of liquid screen 

side temperature (SD1). Solid lines are model predictions based on 

Equation (1) evaluated using room temperature pore diameters.

0

200

400

600

800

1000

1200

0 0.5 1 1.5 2 2.5 3 3.5

450x2750 GHe

325x2300 GHe

∆

P

BP [Pa]

γ Cos θ

C

 [mN/m]

325x2300

450x2750

LH2

400

500

600

700

800

900

1000

14 15 16 17 18 19 20 21 22

450x2750 GHe

325x2300 GHe

∆

P

BP [Pa]

SD1 [K]

325x2300

450x2750
NASA/TM—2013-217834 14

Figure 12.—Liquid hydrogen bubble point as a function of pressurization 

gas for the 325×2300 screen. The solid line is the model prediction based 

on Equation (1) evaluated using a room temperature pore diameter.

Trends here are in excellent agreement with previous bubble point tests in LOX and LCH4 where 

similar disparity in performance existed between the autogenous and noncondensible pressurization 

methods (Hartwig et al. 2012 and Hartwig and McQuillen 2012a). Unlike LOX and LCH4 test results 

however, disparity between GH2 and GHe in LH2 is far outside experimental uncertainty. Analysis of the 

test data for LOX, LCH4, and LH2 shows that this disparity increases as the saturation temperature 

decreases, with the most pronounced difference at LH2 temperatures. On average, GHe data is 13 percent

higher than the GH2 data for the 325×2300 screen. 

In light of the apparent under prediction of the GHe data, Equation (1) may still be used as a lower 

bound to predict screen channel LAD performance for a flight LH2 system using helium pressurization, 

since the actual breakdown pressure is higher for all liquid temperatures tested here. As stated earlier, the 

bubble point represents the maximum allowable pressure drop before gas or vapor is ingested into the 

channel and transfer line to an engine or receiver depot tank. Therefore, Figure 12 implies that using GHe 

to pressurize and/or subcool a flight LH2 propellant tank during expulsion results in higher margin in the 

total allowable pressure loss for the LAD. Higher bubble point pressures can translate into higher flow 

rates to the engine or receiver depot tank.

10.0 Future Work

Future plans include parametric analysis of 325×2300, 450×2750, and 510×3600 bubble point data in 

both LH2 and LN2 to thoroughly examine the effect of liquid temperature and pressure and type of 

pressurization gas on the bubble point pressure. In addition, heated pressurant gas data will also be 

analyzed to determine how warm gas affects LAD performance. Full scale LAD outflow tests are also 

planned, and these tests will examine the various pressure loss terms from Equation (3). 

11.0 Conclusion

A fully robust system has been designed and characterized to examine the effect of screen mesh, 

liquid temperature and pressure, and pressurization type and temperature on the primary performance 

parameter for screen channel liquid acquisition devices, the bubble point. Data has been collected for 

400

450

500

550

600

650

700

750

800

16 17 18 19 20 21 22

GHe

GH2

∆

P

BP [Pa]

SD1 [K]

325x2300

LH2
NASA/TM—2013-217834 15

three different fine mesh screen channel liquid acquisition device samples in liquid hydrogen over thermal 

conditions representative of a low pressure LH2 propellant tank. Preliminary results obtained here in LH2

are in qualitative agreement with previous test results using liquid oxygen and methane.

The low bubble point of the 325×2300 screen hinders the operating range over which single phase 

liquid may be extracted from an LH2 propellant tank. By using the finer 450×2750 screen, operating at a 

colder liquid temperature, and pressurizing the system with a noncondensible pressurant, such as gaseous 

helium, a higher margin in the bubble point pressure is achieved over the 325×2300, thus potentially 

increasing the total allowable flow rate from a liquid hydrogen fuel depot. Bubble points in excess of 900 Pa

are achievable with the 450×2750 mesh, which represents a 170 percent increase over the baseline NBP 

value for a 325×2300 screen. Results indicate that Equation (1) using a room temperature pore diameter can 

be used as a lower bound to predict the GHe bubble point pressure across the full range of temperatures 

representative of a flight LH2 propellant tank. For all screens and temperatures tested, pressurization with

GHe outperformed pressurization with GH2. Compared to test results in other propellants, the disparity in 

performance between the noncondensible and autogenous pressurization schemes is greatest at LH2

temperatures. Results here have direct impact on future LAD and pressurization system design for low 

surface tension liquids, especially future cryogenic hydrogen fueled depots.

References

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Systems—Volume II—Supporting Experimental Program,” NAS8–27685 MDAC Report MDC G5038.

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Advanced Cryogenic Space Propulsion Systems” AIAA/SAE Propulsion Conference Las Vegas, NV, 

Nov. 5–7, 1973.

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System,” NASA CR–134810.

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Systems,” NASA CR–135218.

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Chato, D.J., and Kudlac, M.T. (2002) “Screen Channel Liquid Acquisition Devices for Cryogenic 

Propellants,” AIAA–75–1196, 38th Joint Propulsion Conference and Exhibit Indianapolis, IN, 

July 7–10, 2002.

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Technology for Space Shuttle Reaction Control Tanks,” AIAA–75–1196, 11th AIAA/SAE 

Propulsion Conference, Anaheim, CA, Sept. 29–Oct. 1, 1975.

Hartwig, J.W., and Mann, J.A. (2013a) “A Simplified Model for Predicting the Breakdown Point of 

Screen Channel Liquid Acquisition Devices,” AiCh Journal (submitted for publication).

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AiCh Journal (submitted for publication).

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Oxygen at Elevated Temperature,” 41nd AIAA Thermophysics Conference, Jun. 27–30, 2011, 

Honolulu, HI.

Hartwig, J.W., and McQuillen, J. (2012a) “Analysis of Screen Channel LAD Bubble Point Tests in Liquid 

Methane at Elevated Temperature” 50th Aerospace Sciences Meeting, Jan. 9–12, 2012, Nashville, TN.
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Hartwig, J.W., and McQuillen, J. (2012b) “Screen Channel LAD Bubble Point Tests in Liquid Methane,” 

Journal of Thermophysics and Heat Transfer (submitted for publication).

Hartwig, J.W., McQuillen, J., and Jurns, J. (2012) “Screen Channel LAD Bubble Point Tests in Liquid 

Oxygen,” Journal of Thermophysics and Heat Transfer (submitted for publication).

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AIAA–91–2172.

Jaekle, D.E., (1993) “Propellant Management Device Conceptual Design and Analysis: Sponges,” 

AIAA–93–1970.

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AIAA–97–2811.

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Methane of a Screen Channel Capillary Liquid Acquisition Device,” 54th JANNAF Conference, 

Denver, CO, May 14–17, 2007.

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AIAA–2006–5054, 42nd Joint Propulsion Conference, Sacramento, CA, Jul. 9–12, 2005.

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Kutter, B., Zegler, F., O’Neil, G.O., and Pitchford, B. (2008) “A Practical, Affordable Cryogenic 

Propellant Depot Based on ULA’s Flight Experience,” AIAA Space Conference and Exposition, 

Sep. 9, 2009, San Diego, CA.

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Cryogenic Propellant Depot for Near Term Applications,” 2011 IEEE Aerospace Conference, Big 

Sky, MT.

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Cryogenic Test,” Martin Marietta Corp., Denver, CO, MCR–73–97.

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Performance Gains of Propellant Management Devices for Liquid Hydrogen Depots 

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14. ABSTRACT

This paper presents background, experimental design, and preliminary experimental results for the liquid hydrogen bubble point tests 

conducted at the Cryogenic Components Cell 7 facility at the NASA Glenn Research Center in Cleveland, Ohio. The purpose of the test 

series was to investigate the parameters that affect liquid acquisition device (LAD) performance in a liquid hydrogen (LH2) propellant tank, 

to mitigate risk in the final design of the LAD for the Cryogenic Propellant Storage and Transfer Technology Demonstration Mission, and to 

provide insight into optimal LAD operation for future LH2 depots. Preliminary test results show an increase in performance and screen 

retention over the low reference LH2 bubble point value for a 325×2300 screen in three separate ways, thus improving fundamental LH2

LAD performance. By using a finer mesh screen, operating at a colder liquid temperature, and pressurizing with a noncondensible

pressurant gas, a significant increase in margin is achieved in bubble point pressure for LH2 screen channel LADs. 

15. SUBJECT TERMS

Cryogenic fluids; Cryogenics; Bubbles; Helium; Interfacial tension; Cryogenic fluid management; Liquid acquisition devices; Liquid 

hydrogen; Pore diameter; Contact angle 

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