..
J
TEST DATA ANALYSIS OFA SPRAY BAR ZERO-GRAVITY
LIQUID HYDROGEN
VENT
SYSTEM
FOR
UPPER STAGES
A. Hedayata,
J.
W.
Baileyb,
L.
J.
Hastingsc, and
R.
H. Flachbarta
aNASA Marshall Space Flight Center
Marshall Space Flight Center, Alabama 35812, USA
bSverdrup Technology, Inc.
Marshall Space Flight Center Group
Huntsville, Alabama 35806, USA
CAlpha Technology, Inc.
Huntsville, Alabama 35806, USA
ABSTRACT
To support development of a zero-gravity pressure control capability for liquid
hydrogen (LH2), a series of thermodynamic venting system (TVS) tests was conducted in
1996
and 1998 using the Marshall Space Flight Center (MSFC) multipurpose hydrogen test
bed
(MHTB).
These
tests
were
performed
with ambient heat leaks
=20
and
50
W
for tank
fill levels of
90%,
50%,
and
25%.
TVS performance testing revealed that the spray bar was
highly effective in providing tank pressure control within
a
7-kPa band (131-138 Wa), and
complete destratification of the liquid and the ullage was achieved with all test conditions.
Seven of the MHTB tests were correlated with the TVS performance analytical model. The
tests were selected to encompass the range
of
tank fill levels, ambient heat leaks, operational
modes, and ullage pressurants. The TVS model predicted ullage pressure and temperature
and bulk liquid saturation pressure and temperature obtained from the TVS model were
compared with the test data. During extended self-pressurization periods, following tank
lockup, the model predicted faster pressure rise rates than were measured. However, once the
system entered the cyclic mixinghenting operational mode, the modeled and measured data
were quite similar.
INTRODUCTION
Maintaining propellant tank pressure control while minimizing propellant boiloff loss
is
a significant challenge associated with the storage
of
cryogens in the near-zero-gravity
environment of space. Traditionally, auxiliary thrusters are used to settle the propellants in
order to accomplish tank venting. However, such systems incur weight penalties (associated
with the propellant and hardware required to perform the settling burns) that increase with the
number of settling sequences required during the mission. In addition, tank ventingkesettling
may become necessary at inopportune moments in a mission timeline, thereby increasing
mission complexity. The
TVS
concept, which enables tank pressure control through venting
without resettling, is presented in the following section.
i
THERMODYNAMIC VENTING SYSTEM CONCEPT
Major components
of
the TVS design, shown in
FIG
1,
consist
of
a recirculation
pump, Joule-Thomson (J-T) expansion/shutoff valve, and a parallel flow concentric tube
heat exchangedspray bar apparatus. The pump extracts propellant from the tank and flows
it through the heat exchangedspray bar apparatus. The fluid reenters the tank through
orifices in the spray bar that expel the fluid radially into the tank, resulting in propellant
destratification and ullage condensation. When pressure control within the tank cannot be
maintained through mixing alone (bulk liquid is saturated at the ullage pressure), a small
amount
of
fluid extracted from the recirculation
flow
is passed through the J-T valve where it
is expanded to a lower pressure and temperature. The subcooled liquid is then passed through
the heat exchanger, which extracts heat from the recirculation flow, and subsequently vented
to the environment.
If
ullage instead of liquid enters the recirculation line, vapor is vented
through the J-T valve and ullage depressurization occurs much as it would in a normal
gravity environment. Details
of
the TVS hardware development effort are provided by Lak
and Wood
[I].
THERMODYNAMIC VENTING SYSTEM TEST SETUP
The MHTB aluminum tank is cylindrical with both a height and diameter of
3.05
m, and
2:
1
elliptical domes. The
MHTB
insulation concept consists of a foadmultilayer insulation
(MLI)
combination. The foam element enables the use of a payload bay-type gaseous
nitrogen purge during ground-hold periods. The variable density 45-layer, double-aluminized
DuPont@ Mylar@
MLI
provides thermal radiation protection during the simulation
of
External Vent Back-Pressu re
Component
Box
Zero-Gravity
TVS
Hardware
Zero-Gravity
TVS
Installed in Tank
FIGURE
1.
TVS
hardware
configuration.
on-orbit vacuum conditions. Detailed descriptions of the MHTB and the insulation system are
provided by Martin and Hastings
[2].
The MHTB tank is enclosed within an environmental
shroud, which simulates a ground-hold conditioning purge similar to that in a payload bay,
and enables the imposition of uniform temperatures (ranging from
80
to
320
K)
on the MLI
external surfaces.
Testing was performed at the MSFC east test area thermal vacuum facility, test stand
300. The facility systems, in combination with the MHTB’s shroud, enabled simulation of
orbital environmental conditions. During the TVS testing, the chamber maintained vacuum
levels in a low
lo4
torr level with the cold walls operating, and torr without the cold
walls.
4
TEST
PROCEDURES
Boiloff testing was conducted to determine the ambient heat leak into the MHTB tank
and to set up consistent baseline conditions prior to each of the TVS tests. The first TVS test
series was conducted with the vacuum chamber liquid nitrogen (LN,) cold walls operating
to produce a minimum heat leak condition of
19-20
W.
The second series was conducted
without the LN, cold walls, thereby providing a higher ambient heat leak condition
(51-54
W)
and reducing test costs. Detailed descriptions of the boiloff tests and procedures are reported
by Flachbart et al.
[3].
After the heat leak testing was completed for each tank
fill
level, the tank was locked up
and allowed to self-pressurize until the ullage pressure attained the maximum tank pressure
set point of
138
kPa. Upon reaching this pressure, the recirculation pump was activated, and
mixing continued until the ulllage pressure reached the minimum set point
(13
1
Pa). After
the pump was turned
off
at the minimum set point, the tank was allowed to self-pressurize
and the cycle began again. This automated operation cycle, with a control band of
k3.45
Wa,
continued until the
LH,
saturation condition attained the lower set point. The J-T vent valve
was then opened and the spray bar heat exchanger was utilized to extract thermal energy
from the bulk liquid. Thereafter, the J-T valvehpray bar heat exchanger was used each time
the pump cycled on.
MODEL OVERVIEW
A
transient one-dimensional analytical model
of
the TVS was formulated to characterize
the TVS performance within the MHTB LH, tank. The TVS performance formulation is
comprised
of
four combined thermal/fluid models
-
the heat exchanger, spray manifold and
injector tubes, recirculation pump, and tank. The heat exchanger model calculates the quality
and two-phase pressure loss at the vent exit. The spray manifold and injection tube model
determines the pressure drops within the manifold and tubes along with the spray flow rates
and velocities leaving the injection orifices. The recirculation pump model calculates the
pump head increase from the pump speed and the head coefficient curve provided by the
pump manufacturer. The tank model
is
a lumped model consisting of four control volumes-
the ullage, tank wall, tank wall liquid, and bulk liquid. The ullage is modeled to contain
gaseous hydrogen (GH,) and/or gaseous helium (GHe). Since the major thrust of the current
effort has been to correlate this analytical model/code with the
MHTB
test data, the TVS
analytical model is not provided here.
A
detailed description of the TVS analytical model
is
given by Nguyen
[4].
7
RESULTS AND DISCUSSIONS
The TVS testing was conducted in
1996
and 1998; seven tests were modeled for
comparison with the TVS performance code. The tests were selected to encompass the range
of tank fill levels, ambient heat leaks, operational modes, and ullage pressurants. The model
is in good agreement with the test data in the early stages of self-pressurization after tank
lockup; however, the analytical pressure begins to deviate and rises more rapidly than the
measured values. It is believed the analytical modeling, which assumes that the liquid and
ullage are each represented by a single node, did not accurately simulate the complex energy
exchange that actually occurred at the liquid-vapor interface. Therefore, the analytical
pressure rise rate after tank lockup is conservative relative to the measured data and the
modeled mixing cycles begun earlier.
With mixing, the stratification effects are minimized and the energy exchange across
the liquid surface is more predictable during the relatively short self-pressurization periods
between mixing cycles. Therefore, once the mixing and pressure rise cycles began, the
analytical and measured data closely matched. However, it was noted that the measured
pressure rise rates were slightly steeper than analytically modeled, whereas the pressure
reduction rates were practically identical. Correlations for the bulk liquid saturation pressure
and temperature indicated relatively good agreement for the entire range of conditions tested.
A summary of test segments compared with the TVS performance code is shown in TABLE
1.
Detailed model predictions and comparisons for test segments P263968G, P26398 lD,
and P263981X are discussed in the following sections.
Test Segment P263968G-Low Heat Leak, Mixing and Venting Mode, 90% Fill Level
Correlations
for
the venting and mixing operation were performed using test segment
€2639686, with a
90%
fill level and
20.2-W
tank heat leak. The ullage pressure prediction
by the model is in good agreement with the test data as depicted in
FIG
2.
The model ullage
pressure predictions are within the prescribed control interval of
131-138
kPa. The data
comparison indicates that, initially, the analytical model tracks the measured data very
well; however, the model calculates the cycling time to be slower than that of test data. The
calculated
and measured cycling times are
1.58
and
1.06
hr, respectively. The calculated
and measured ullage temperatures are shown in
FIG
3.
The predicted ullage temperature
TABLE
1.
TVS Analytical Correlation Cases
~~~~~ ~~~
Cycle
Fill Heat Pressure Rate
Level Leak Year,Test Rise Error
Ullage
(%)
(W)
SegmentNo. Operation Mode (Modemest)
(%)
~ ~ ~~~ ~ ~
GH,
90 54.1
1998, P263981D
Tank lockup/mixing 2 4
GHZ
90 20.2
1996, P263968G Mixinglventing
NA 33
GHZ
90 20.2
1996, P263968E&F
Tank lockup/mixing 2
11
GHZ
50 51.0
1998, P263981T Tank lockup/mixing/venthg
5
14 w/o venting
GH,& GHe 50 51.0
1998, P263981X MiXindvenhg
NA NA
GH,
50
18.7
1996, P263968K Mxhdvenhg
NA
.
28
GH,
25 18.8
1996, P263968L Mixhgvenhg
NA 26
0
wlventing
13t
132
-
(0
B
134
Y)
?!
a
132
130
I
Test
-c
Model
20,000 40,000 60,000 80,000
IO(
Time
(s)
FIGURE
2.
Ullage pressure, test P263968G.
24
*Test,
95%
Fill
-Test,
98%
Fill
+Model
23.5
1
-
y
23
t
E
22.5
n
22
21.5
B
30
0
20,000 40,000
60,000
80,000
10C
Time
(s)
FIGURE
3.
Ullage temperature, test F263968G.
300
is
21.2-22.3
K,
which is within the measured ullage temperatures
(21.2-23.1
K).
FIGS
4
and
5
show good agreement between the predicted and measured bulk liquid saturation pressure
and temperature values.
Test Segment
P263981D
-High Heat Leak, Self-pressurization and Mixing Mode,
90%
Fill Level
Model correlations with test data for self-pressurization and venting with a
90%
fill
level and a
54.1-W
heat leak were performed.
As
illustrated in FIG
6,
tank lockup occurred
at
10,380
s
and self-pressurization proceeded until the mixing cycles began and continued
throughout the remainder
of
the test segment without venting. The model is in good
agreement with the test data in the early stages
of
self-pressurization; however, the analytical
pressure data begins to deviate after
-14,500
s
and rises more rapidly than the measured
values.
A
similar comparison occurred with the
low
heat leak test series. The computed and
measured ullage pressures reached the upper pressure limit
of
137.9
kPa after about
20,000
s
and
30,000
s,
respectively. However, once the mixing cycles began, both the pressure rise
and reduction rates correlated very well. The deviation of ullage pressure prediction during
the tank lockup could
be
attributed
to
stratification effects that are not addressed by the
analytical model. FIG
7
represents comparison between calculated and measured ullage
135
.C
Model
134
1
11
Test
+Model
0
20,000 40,000 60,000
80,000
100,000
0
20,000 40,000
60,000
80,000 1Oi
Time
(s)
Time
(5)
100
FIGURE
4.
Bulk liquid saturation pressure,
FIGURE
5.
Bulk liquid saturation temperature,
test 12639686. test P263968G.