The objective is to optimally design the thermal components of a system that uses carbon dioxide (CO2) from the Martian atmosphere to produce oxygen (O2) for spacecraft propulsion and/or life-support. Carbon dioxide is thermally decomposed into carbon monoxide (CO) and O2 followed by the electrochemical separation of O2. The design of the overall system and its various individual components depends on, among other things, the fraction of the stoichiometric yield of O2 that can be realized in the system and the temperature of operation of the electrochemical separation membrane. The analysis indicates that a substantial reduction could be obtained in the mass and power requirements of the system if the unreacted CO2 were to be recycled. The concepts of an optimum temperature of the zirconia cell and impracticality of plant operation at low cell efficiencies are also discussed. The design of the thermal equipment is such that the mass and power requirements of the individual components and of the overall system are optimized

Iyer, Venkatesh A.

Sridhar, K. R.

English

NASA Technical Reports Server

N9 1 - 2 4 :2
Thermal Analysis, Optimisation and Desi.qn of a Martian
Oxygen Production Plant
Venkatesh A. lyer and K.R. Sridhar
Department of Aerospace and Mechanical Engineering
The University of Arizona
o
, _.. _.
Abstract
The objective of this work is to optimally design the thermal components of a system that uses
carbon dioxide (COz) from the Martian atmosphere to produce oxygen (02) for spacecraft
propulsion and/or life-support. CO 2 is thermally decomposed into carbon monoxide (CO) and 02
followed by the electrochemical separation of Oz. The design of the overall system and its various
individual components depends on, among other things, the fraction of the stoichiometric yield
of 02 that can be realised in the system and the temperature of operation of the electrochemical
separation membrane. The analysis performed indicates that a substantial reduction could be
obtained in the mass and power requirements of the system if the unreacted CO 2 were to be
recycled. The report also discusses the concepts of an optimum temperature of the Zirconia cell
and impracticality of plant operation at low cell efficiencies. The design of the thermal equipment
would be such that the mass and power requirements of the individual components and of the
overall system would be optimised.
Introduction
There has been tremendous interest in recent years on the need for in-situ-propellant-processing
(ISPP) [1,2]. More than 80% of a spacecraft's mass is due to the propellant. Hence production
of propellants at locations remote from the Earth is essential for frequent and extended space
explorations. Here, the production of Oxygen from the predominantly CO 2 atmosphere of Mars
is explored. The system that would be used on Mars would have a filter at the CO 2 intake end to
remove the dust and particulates from the supply gas. It would also have a liquefaction and
storage unit at the Oz production end. These components are not being designed by UA/NASA
https://ntrs.nasa.gov/search.jsp?R=19910015062 2020-03-19T18:26:14+00:00Z
IB-IO
SERC presently. However, they will be part of our future investigations. The system currently being
designed incorporates the sub-systems necessary for supplying the CO z at Martian pressures and
temperatures. This system will be designed, fabricated and tested extensively on our ground-
based facilities. To distinguish it from the future flight-tested system for Mars, the ground-based
testing system will be referred to as the "Test Bed".
Description of the Test Bed
The Test Bed, shown in figure 1, can roughly be divided into 4 sub-systems:
1. The simulation sub-system consists of CO z supply and a Cryo-Vacuum Chamber cooled by a
Cryo-cooler. This simulates the Martian ambient conditions of 6.4mbar pressure and a
temperature of 200K. Though the Martian atmosphere contains only 95.3% CO 2, presently the
supply gas is composed entirely of CO 2.
2. The compressor, heat exchanger and heater form the CO 2 preparatory sub-system. In order
• to avoid a very bulky system at high CO 2 mass flow rates, the supply gas is pressurized from 6.4
mbar to higher pressures. At the present time CO 2 is assumed to be pressurized to 1bar. The
reasons for this are two-fold. The Test Bed is operated on the Earth and operating the system at
1 bar minimises the probability of gas leaks. Secondly, electrochemical separation membranes
have been tested extensively with other gases at 1 bar and higher pressures and its performance
at lower pressures is not known currently. The effect of lower pressures on the electrochemical
separation is the subject of a related investigation at UAJNASA SERC. After the compressor, the
pressurized CO 2 enters a waste heat recovery heat exchanger to gain energy from the gases
exhausting from a Zirconia (ZrOz) Cell, which will be discussed later. The CO z supply gas then
enters a heater where it reaches the same temperature as the ZrO z electrochemical separation
membrane unit. Thermal decomposition of CO z occurs in the heater. Since the temperature of the
gases entering the separation unit is the same as that of the unit itself, the separation unit is not
subjected to any thermal shocks.
3. The mixture of CO, 02 and CO z passes through the ZrO z cell, which electrochemically
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separates the 0 2 from the mixture. This O z is then fed to a storage device for later use. The cell
forms the oxygen separation sub-system.
4. The exhaust from the Cell, after passing through the heat exchanger, loses energy in a radiator
before passing through a polymeric membrane separator which separates the unreacted CO 2.
The radiator is essential because polymeric membrane separators cannot withstand the high
exhaust temperatures. There is a substantial pressure drop across the membrane separator,
which requires a second compressor to re-pressurize the CO 2 to 1 bar. The CO 2 then re-enters
the loop just before the heat exchanger. The system is depicted in figure 1.
Figure 2 illustrates the system without recirculation. The exhaust gas, after passing through the
heat exchanger, is expelled from the system. Henceforth, the system wit_h.hrecycling is referred to
as case 1, and the one without, as case 2. The number of moles of CO, CO z and 0 2 flowing at
various points in the system for both cases 1 and 2 is shown in table 1.
Nomenclature
e = Fraction of 0 z produced that is electrochemically separated by the cell.
I_sep = Membrane Separation Factor =
This is only for case 1.
Amount of CO.2 at Point 9
Amount of CO 2 at Point 8
,, = Cell Factor = moles of unreacted CO z for each mole of 0 2 produced
This determines the cell efficiency _¢eLt as:
_eett = moles of 0 z actually produced = 2/(2+,_)
theoretical max. no. of moles of 0 z possible
The Cell reaction is:
[(2+a)/e]CO 2 = [2/elCO + [l/e]O z + [ale]CO z
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Table 1: For Oxygen Production of 1 mole
Point Case 1
1 2+c  , co 2
e
2 2+(.__._CO z
e
3 2+(_2_t__CO 2
¢
4 2+(___._)COz
¢
5 2+(g._.)COz
@
6 [We]CO+ [(1 -e)/e] 02 + [,,/e] CO 2
7 [2/e]CO + [(1-e)/e]O2+ [,_/e]CO 2
8 [2/e]CO+ [(1-e)/e]O 2+ [,,/e]CO 2
9 [l_sepOC/e]CO2
10 [ IAsep¢/e] CO 2
Case 2
2+(__#__3CO 2
e
2+(__#__)CO 2
e
2+(&t_3co2
¢
2+(__#__)CO 2
e
2+(_____)0o2
@
[2/e] CO + [(1 -e)/e] 0 2+ [=/e]CO 2
[2/e]CO + [(1 -e)/e] 02 + [=/e]CO 2
Assumptions
The thermodynamic analysis has been performed based on the following assumptions:
1. The Membrane Separator fillers out all the CO and Oz from the exhaust gases. It is known that
the purity of the separated gases is very high. Hence this is a valid first-order approximation.
2. Only a steady state analysis has been performed.
3. Pressure drop and heat losses in the pipes, valves and bends are negligible.
4. The Compressors have been assumed to be isentropic, with specified mechanical and
isentropic efficiencies. Radiation losses from their surfaces have been neglected since they are
small in comparison.
The analysis gives the total heat transferred from the exhaust to the fresh CO z. Using an iterative
procedure, the cold side outlet temperature and the Logarithmic Mean Temperature Difference
(LMTD) are determined. Knowing the mass flow rates, the heat exchanger can be designed. The
power requirement of the ZrO 2 cell - the Nernst potential and the ionic component of the current
IB-15
corresponding to the oxygen ion flow can be calculated precisely. For a fixed 0 z production rate,
it is a constant.
Discussion of Results
The analysis was performed with the help of an interactive program written in Pascal. The
program has pull-down menus using which the user can change any of the input variables. This
was necessary because the optimal cell temperature, cell efficiency and membrane separation
factor is presently not known. Concurrent research is being done at UAJNASA SERC to investigate
the optimum system performance conditions. Some graphs have been obtained from the data
generated by using the program for an oxygen production rate of 10 kg/day, which corresponds
to the production rate needed for an unmanned sample return mission [2]. Figure 3 shows the
variation of the specific heats of CO 2, CO and 0 with temperature. Empirical relationships [3]
have been used to determine these curves.
This variation in the specific heat (Cp) precludes
treating Cp as a constant, say, across the heat
exchanger. Over the entire range of
temperatures considered, it is seen that Cp
increases with temperature for all three gases,
with the curve for CO z increasing most sharply.
Figure 4 shows the graph of heater power with
ceil temperature O's). The power consumed by
the heater increases with Ts as expected, for
t400.O
O_
,_1200.0
..,_-
0
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V1 800.0
_,_,-N_ COz
02
CO
600.0 i Izoo.o soo.o loo_.o 14oo.o
Temperature, K
Fig 3: Variation in Cp
low values of T5. It reaches a maximum, and then, contrary to expectations, starts dropping.
Figures 6 and 7 show plots of temperature of the cell ('1"5)with the outlet temperature T4 of the
cold CO 2 from the heat exchanger. Though not apparent from the figures, (Ts-T4) actually
increases with T 5. The anomalous behaviour of figure 4 can be explained by the fact that though
Cp itself increases, and (Ts-T4) also increases, the integral of CpdT over the temperature range
IB-16
200.0
180.0 ,
h2
o
_ 160.0 -
o
--e-
l 40.0 -
120.0
600.0
Cell Effic =O.6
Cell Efflc=O 5
Cell Effic=O.6
i
800.0 1000.0 1200.0 1400.0
Temperature Of Cell, K
Fig 4: Heater Power - Case 1
1600.0
320.0"
280.0 "
• 240.0 -
o
n
.*_200.0 -
0
"1"
160.0 -
i _ Cell Effic=0.4
Cell Eff_c=O.5
6
120.0
I 8J.0_ J J 1 , [6OO.O 1000.0 1200.0 1400.0 1600.0
j Temperature Of Cell, K
Fig 5: Heater Power - Case 2
v
C_1400.0
O
C_ 1200.0
2 1ooo.o
E
_._ 800.0
0
600.0
600.0
Cell Effic=0,4
Cell Effic=0.5
Cell Effic=0,6
800.0 ' _ ' 'I0OO.O I2o0.o 14o0.0
Temperature of Cell, K
v
(_1400.0
"10
O
C_ 1200.0
2 1_o.0
E
_ 8OO.O
23
0
600.0
600.0
Cell Effic =0,4
Cell Effic=0.5
Cell Effic=0.6
O_ , i , i i8 .0 1000.0 1200.0 1400.0
Temperature. of Cell, K
Fig 6: T4 v/s Ts - Case 1 Fig 7: T4 v/s T5 - C_se 2
T 4 to T5 actually decreases after a point. This is opposed to the almost linear increase of heater
power with temperature in case 2 as shown in figure 5. In this case, though the mass flow rates
of CO z at points 4 and 5 are the same as in case 1, T3, and hence T4, are lower for the same
value of Ts. This is because the recirculating CO 2 is at a higher temperature than that of the CO 2
at point 2. Hence, the increased (15-'1"4)term counters the decrease in the slope of the Cp curve
IB-17
enoughto causethe heaterworkto increase.Here,andin discussionslaterin this report,we
freely use the '(Ts-T4)" argument. Although the empirical relations used for Cp were fourth-order
polynomials in temperature, terms such as (T54-T44)etc. depend essentially on (Ts-T4). Note that
the magnitude of heater work is greater for case 2 than for case 1, for a fixed T5. It may be
expected that at cell temperatures higher than those considered, figure 5 behave similar to figure
4. However, in practice, cell temperatures higher than 1500K are not of interest.
:=60.0 1
 oooI
o 40,0
_ 30.0
- I09.04
- g9.04 ,_
-8g.04 .
"6
- 79.04 C_
3
O
- 69.04 L=.
-59.04 0
-49,04
20.0 I 39.040.0 0:1 o'2 0.'_ 0:4 0._ o'e 0/7 0.0
Cell Efficiency
Fig 8: CO 2 Intake - Case 1
160.0
140.0
_ 120o0
0
_c ao.o
60,0
_ 40.0
20.0
0.0 0.'1
-299,04
-279.04
- 259,04
- 239.04
- 219,04 .
- 199.O4
- 179.04 Q_
- %59.04
- _ 39.04 EL.
- 119.04 _C_
-99.04
_"59,04
J 39,040.'2 0.'_ 0.'4 0._ o.'s 0.'7 0.,,
Cell Efficiency
Fig 9: CO 2 Intake - Case 2
Figures 8 and 9 show the CO 2 intake required for cases 1 and 2 respectively. Clearly, the intake
required increases sharply as the cell efficiency approaches zero. The intake in cubic-feet per
minute (cfm) at the inlet pressure of 6.4 mbar gives us an idea of the size of pumps that we will
require. Note that for a given cell efficiency, the intake required in case 2 is greater than that in
case 1, especially at lower efficiencies. Since the slope of the curve decreases rapidly with
increasing efficiency, and keeping in mind that the intake determines the capacity of the
Compressor C 1, the conclusion is that the compressor capacity becomes impractical at very low
efficiencies. For a start, we take an efficiency of 0.2 which is reasonable at this point of time. The
pumping speed required is about 280 cfm for case 2 and about 115 cfm for case 1. If we take an
IB-18
off-the-shelf vacuum pump which can fulfil these requirements, then for case 2, an S Series Rotary
Vane Pump Model $630F manufactured by Leybold Vacuum Products Inc. will suffice. It weighs
1450 Ib and has a power requirement of 25 hp. For case 1, where the CO 2 is recycled, a much
smaller pump such as the Leybold Model $160C, which weighs 310 Ib and consumes 7.5 hp can
do the job. Also required in case 1 is a much smaller pump (compressor C2) to compensate for
the pressure drop across the membrane separator. For this a small pump like the Leybold TRIVAC
Rotary Vane Pump Model D2A will do. This pump weighs 41 Ib and consumes 0.33 hp. This
clearly indicates that at this point, though case 2 appears compelling from the points-of-view of
system simplicity and reliability, we must recycle the CO 2 in order to design a reasonably compact
system. Note that we must also consider the mass added to the system by the radiator and the
membrane separator. However, the reduction in compressor mass and power obtained by
recycling offsets the increase in mass due to the radiator and membrane separator. Having
decided on recycling, we take an optimistic look at what our requirements would be if we were
successful in obtaining a cell efficiency of, say, 0.5, which gives us a CO z intake of about 70 cfm.
To take care of this requirement, a first approximation would be a pump like the Leybold S100C,
with a mass of 220 Ib and a power requirement of 5.0 hp. This makes the system design very
feasible as far as its mass requirements go. Obviously, for an actual mission, a much more
240.0
220.0
1
_ 2_.o -
o
n 180.0 -
_1_.0 -
140.0 - e.ee.e-e Cell Temp=14OOK
Cell Temp-1200K
Cell Temp= 10OOK
Cell Ternp- 9OOK
120.0 I
o.o o.'1 0.'2 0.3 o.'4 03 o.'e 0.'7
Cell Efficiency
0.8
500.0 ,
400,0 l
• 300.0
o
n
 ,oo.ot
@ .
"T -4
J100.0
| _Nt.e-,w=Cell Temp=14OOK
Cell Temp-1200K
_ Cell Temp=lOOOK
Cell Ternp- 9OOK
0.0 o.o o.'1 0.'2 0.3 o.'4 o.'s
Cell Efficiency
i
0.6 O,r7 0.8
Fig 10: Heater Power v/s "Ccett- Case 1 Fig 11: Heater Power v/s "_cett " Case 2
IB-19
sophisticated pump with better materials and design would be used. But it would depend on
whether an efficient Zirconia cell can be designed. Figure 10 is a graph of heater power versus
cell efficiency for case 1. In the light of information obtained from figures 4 and 10, we realise that
at low efficiencies the heater power increases because of increased CO 2 mass flow rate. But the
interesting point made by this graph is that at higher cell temperatures and lower efficiencies, the
effect of the cell temperature prevails, and causes the slope of the curve to decrease.In fact, the
curve for a cell temperature of 1400K is almost flat, suggesting that there could be a temperature
T5 between 1400K and 1200K at which the load on the heater is almost constant. Also important
is the fact that it is a lower temperature difference 0"5-'1"4)that 'overcomes' the increase in the
slope. Figure 11, for case 2, is clearly a follow-up of figure 5. Here, the temperature difference (T5-
1"4) does not fall sufficiently to cause the slope to decrease.
Plans For The Future
1. The system design depends, to a large extent, upon the system operating pressure. The
optimum system pressure needs to be examined. If the system pressure were low, then the option
of using a blower (lower mass) instead of a compressor will be examined.
2. The heat exchanger design presently limits the exhaust gas temperature to 70OK, due to fears
of carbon deposition. Suitable changes will be made in the design when a related investigation
by Prof. D.C. Lynch, Department of Materials Science, University of Arizona, provides results on
the seriousness of carbon deposition.
3. The pressure and heat losses in the system components and piping will be incorporated in the
analysis and
4. The individual components will be designed and tested.
Acknowled.qements
The graphics routines of the Pascal program used in the analysis were written by Mario Rascon
and Kirby Hnat. Their help is gratefully acknowledged.
IB-20
References
1 K Ramohalli, E. Lawton & R Ash, 'Recent Concepts in Missions to Mars: Extraterrestrial
Processes', Journal of Propulsion and Power, Vol 5, pp 181-187, 1989
2. R L Ash, W.L. Dowler & G Varsi, • Feasibility of Rocket Propellant Production on Mars', Acta
Astronautica, Vol 5, pp 705-724, 1978
3. Thermophysical Properties of Matter, The TPRC Data Series, Vol 6, p 50, 145 & 154, IFI/Plenum,
1970


Thermal analysis, optimization and design of a Martian oxygen production plant

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