AIAA 2010-7129
46th AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit
25 - 28 July 2010, Nashville, TN
Design of a Catalytic Nitrous Oxide Decomposition Reactor
Bradley D. Hitch* and David T. Wickham†
Reaction Systems, LLC, Golden, CO 80401
In an SBIR Phase I project for AFRL, Reaction Systems completed the conceptual design
of a scramjet missile ignition system using stored nitrous oxide for fuel atomization
barbotage gas and/or a pilot ignition torch. Exothermically decomposing N2O to N2 and O2
can deliver over 845 kW of thermal power per pound per second of N2O flow to aid the
engine ignition process. We found that purely thermal decomposition of N2O was
impractical from the standpoint of both required reactor volume and the product
distribution obtained based on both computer simulations and lab experiments. We also
synthesized and demonstrated highly-active, thermally-stable N2O decomposition catalysts
with good low temperature activity and high selectivity for the formation of O2 and N2.
Finally, we designed an N2O storage and supply system with a regeneratively-cooled N2O
heat exchanger/reactor featuring a wall-mounted catalyst to minimize both pressure drop
and catalyst temperature. This paper discusses our N2O storage and supply system design as
well as our conceptual design of the N2O catalytic decomposition heat exchanger/reactor.
Nomenclature
Cf
St
Pr
jH
jm
Sc
Nu
Re
h
hm
D
k
Le
ρ
cp
P
V
=
=
=
=
=
=
=
=
=
=
=
=
=
=
=
=
=
wall friction coefficient
Stanton number
Prandtl number
Chilton-Colburn “j” factor for heat transfer
Chilton-Colburn “j” factor for mass transfer
Schmidt number
Nusselt number
Reynolds number
heat transfer coefficient
mass transfer coefficient
diameter
thermal conductivity
Lewis number
fluid density
fluid specific heat
pressure
volume
I. Introduction
High speed flight in the atmosphere presents a number of challenges, such as very high air recovery temperatures
and aerodynamic heating loads, high aerodynamic drag, and large internal flow total pressure losses. Supersonic
combustion ramjets (scramjets) are required above Mach 6, where the majority of the air flowing through the engine
is maintained supersonic to avoid very large engine internal pressures and obtain acceptable fuel heat release
efficiencies. Storable hydrocarbon fuels suitable for use under these conditions generally have both high thermal
stability and low volatility. Due to the operating complexity and increased weight of providing multiple air
breathing engines to propel a vehicle across the entire speed range, present designs for scramjet-powered hypersonic
missiles anticipate employing simple rocket boosters to bring them up to a minimum operating speed where a dualmode ram/scram engine can take over. Unfortunately, igniting scramjet engines at altitude is difficult. Low air
pressure, low air temperature, short residence time all combine to make reliable ignition and stable flame holding a
*
†
Chief Engineer and AIAA Senior Member.
President and AIAA Member.
1
Copyright © 2010 by Reaction Systems, LLC. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.
II.
Static Pressure
(psia)
Static Temperature
(F)
challenging problem. In addition, these missiles
would probably be launched from platforms such
Static Temperature and Pressure
as the B-52 or F-22 and therefore could spend a
considerable amount of time cold-soaking at
(Standard Atmosphere)
100
80
altitude before they are fired. Figure 1 shows the
Probable
temperature range expected during capture carry
Loiter Carriage
60
Altitude Range
is between -20 to -70°F. All of these conditions
10
40
can profoundly affect the ability of the scramjet
engine to start properly after the burnout of the
20
rocket booster.
1
0
Current scramjet engine designs generally
0
50000
100000
150000
use shallow side wall cavities for flame
-20
stabilization.
Achieving
ignition
and
0.1
-40
flameholding in these cavities has proven to be
difficult, particularly with cold hardware and
-60
low-volatility fuels such as JP-7. Even under the
-80
0.01
best of conditions flameholding is marginal
within the sidewall cavities due to very high
Altitude (ft)
turbulence levels and short particle residence
times. The minimum scramjet operational speed Figure 1. U.S. Standard Atmosphere: variation of static
can be estimated by comparing estimated 1 temperature and pressure as a function of altitude.
millisecond ignition delay temperatures to inlet
air total temperature as a function of flight Mach number, as shown in Figure 2. This shows that while scramjet
operation down to Mach 4 should be viable with hydrogen fuel, the slower ignition process of kerosene-type fuels
requires a higher minimum flight Mach number to achieve stable flameholding. These operational Mach numbers
can be expected to be higher when trying to ignite a cold engine. Techniques to improve both ignition and
flameholding during the cold start process are therefore of great interest to minimize the flight Mach number
scramjet engine takeover speed and reduce the size of the required low Mach number vehicle boost system.
N2O Storage and Delivery System
T tot (K)
In this project we completed conceptual designs of both an N2O storage and delivery system and a catalytic heat
exchanger/reactor (HX/R) N2O decomposition system to supply hot gas for both barbotage fuel injectors and a pilot
ignition flame. A nominal engine air flow of 100
lbm/sec was used to be consistent with the midsize Robust Scram studies, assuming an overall
Constant Q Trajectory
stoichiometric conditions. Figure 3 shows a
Inlet Total Temperature
schematic for the N2O storage and delivery
4500
system. In this system the nitrous oxide is stored
Q=4000 psf
as a compressed liquid and we will use on/off
4000
valves and metering orifices, which will provide
Q=2000 psf
3500
rapid and accurate control and regulation of flow.
Q=1000 psf
3000
Even though liquid N2O has high vapor pressures
Q=500 psf
at low temperatures, two phase flow could occur
2500
as the liquid passes through the metering orifices
2000
and drops pressure. To avoid this potential
1500
problem, a high pressure helium bottle will be
Kerosene
used to pressurize the N2O liquid in its reservoir.
1000
H2 ignition
The N2O is contained in a bladder or bellows to
500
deliver compressed liquid regardless of the
0
vehicle orientation. The line connecting the high
0
2
4
6
8
10
pressure helium bottle to the N2O reservoir
contains an on/off valve and a pressure regulator
Flight Mach Number
for a constant N2O supply pressure during system
Figure 2. Hydrogen and kerosene 1 millisecond ignition
operation.
delay times plotted against constant-Q trajectory inlet air
total temperatures.
2
Helium
Pressurant
100 lbm/sec Engine Air Flow, φoverall = 0.9
Pressurization
Control Valve
N2O Liquid
(in bladder)
1000 psia
-60 F
Liquid
Supply
Valve
High Flow
Orifice
High Flow
Valve
N2O Catalytic
Decomposition
Chamber
Low Flow
Orifices
Pilot HX/R
1.84 lbm/sec
2428 F
33% O2
-59 F
500 psia
Liquid
0.239 lbm/sec
2428 F
33% O2
JP-7
Fuel
Pilot Flame
Supply Valve
Pilot
Orifice
Pressurization
Valve
High Pressure
Helium Bottle
(6000 psi)
to φ=3 Pilot
5.97 lbm/sec
500 psia, -60 F
Barbotage
Injector
0 F, 4 wt% O2+2N2
- or 111 F, 3.81 wt% CO2 + 3.14 N2
to Combustor
Figure 3. N2O supply system schematic and conditions for cold-soaked start.
As shown, the system contains both high and low flow paths. The N2O in the high flow path, about 1.84 lbm/s,
will be used in the pilot ignition system, while that going through the low flow pathway (about 0.239 lb N2O/s) will
be used in a barbotage fuel injection system. Each pathway has both low and high flow orifices. Initially in each
path, the N2O flows through the low flow orifice, which limits flow to about 10% of the full value. This flow is
directed into a catalytic heat exchanger reactor that contains a calrod heater coated with the N2O decomposition
catalyst. A small amount of electrical power is used to preheat the catalyst to 350°C. At that temperature N2O will
decompose exothermically on the catalyst into O2 and N2, generating much higher temperatures and heating the
reactor substantially. Once the reactor has been heated to its full operating temperature the main valves are opened
and the full N2O flows through each leg are decomposed exothermically into oxygen and nitrogen with a potential
equivalent thermal power of 1.75 MW to aid engine ignition.
Current scramjet barbotage fuel injection systems are supplied with ~500 psig air at about 4% by weight of the
fuel flow to assist in fuel atomization and injection into the combustor, so those values were specified here. Under
the cold-soak conditions assumed in Figure 3, the N2O remains in the liquid state until it enters the HX/R, where it is
then boiled by the decomposition heat release. Mixing the hot barbotage gas with the cold fuel results in an overall
fuel temperature rise of ~60°F in the absence of combustion. The fuel/gas mixture temperature could be raised to
~111°F if some of the fuel is used to burn the oxygen in the decomposed N2O stream, assuming only CO2 and liquid
H2O as combustion products. This temperature rise would significantly improve fuel atomization and penetration
performance compared to simply injecting very cold liquid fuel using a pressure atomizer.
The amount of N2O required depends on whether it will be used to only supply barbotage gas or if it will also be
used to provide pilot flame ignition. If we assume that 10% of the overall fuel flow will be consumed in a pilot
3
Equation 1
N2O Mass Flow
2.5
Pilot + Barbotage
Total Flow Rate, lb m /s
burner with N2O at the theoretical sooting limit
(i.e. equivalence ratio ~3), an ignition pilot flame
running for five seconds would consume 9.2
pounds of N2O for a 100 lbm/s combustor air
flow. The barbotage gas flow for this same
engine would consume only 7.7 pounds of N2O
while operating over 30 seconds with a gas mass
fraction of 4%, as seen in Figure 4, and operates
until cracked fuel products are available from the
combustor cooling system. The total usable N2O
inventory required would be 16.4 pounds with a
tank volume of around 0.4 ft3 to accomodate
ullage and unusable propellant. A spherical tank
of this volume made of 120 ksi titanium and
designed for a 9000 psi maximum working
pressure would weigh about 12 lbs. For a carbon
composite tank, weight can be estimated from the
relation:
2
mN2O = 16.4 lbm
VN2O = 0.334 ft3 @ 70 F
1.5
1
0.5
Barbotage Only
0
0
10
20
30
40
Time, seconds
Figure 4. N2O mass flow rate versus time.
PV
≈ 106
m
where P is the pressure in psi, V is the volume in cubic inches, and m is the mass in pounds. This relation indicates
that a carbon composite N2O tank would only weigh about 6 pounds, for a fully loaded weight of approximately 22
lbs. Weight associated with the helium pressurization system, valves, and delivery piping can be expected to add
another 10 - 15 pounds to the overall ignition system weight.
III.
Catalytic Decomposition Heat Exchanger/Reactor Design
Our initial objective in this project was to combine the catalytic N2O decomposition with gas phase reaction
because catalysts reduce the temperature needed for a reaction to begin, but once the temperature is high enough,
gas phase reactions can be much faster than those that occur on the surface of a catalyst. To investigate this
approach, we modeled the gas phase reaction assuming an adiabatic perfectly-stirred reactor (PSR) as shown in
Figure 5, with an existing gas phase detailed kinetic
mechanism for MMH/RFNA combustion that contained
the N2O thermal decomposition mechanism1.
Partially
In these calculations we started with very high N2O
τ
Reacted
N
O
res
2
inlet temperatures to ensure ignition, then gradually
Mixture
decreased the inlet temperature to observe where blowout
occurs, as shown in Figure 7.
The compositions
dYk
ω& w
1
associated with the points identified in Figure 7 are shown
= − (Yk − Yk* ) + k k
dt
τ
ρ
in Figure 6. Point #1 (0.01 sec residence time) had the
Figure
5.
Perfectly
stirred
reactor
model.
best conversion of N2O to O2, but a very high reactor
operating temperature (1788 K, 1515°C) due to the high
inlet N2O temperature. A substantial shift in the exit temperature was observed to occur just prior to Point #2 due to
much higher levels of endothermic NO being formed, as seen in Figure 6. As we continued to decrease the inlet
N2O temperature we started to see breakthrough of N2O (Point #3), and finally the reaction in the PSR blew out and
the exit temperature became the inlet temperature (Point #4). At 733 K this temperature is still well above the
expected N2O temperature that could be supplied to the reactor from a missile N2O tank.
The same calculation was repeated for a 0.1 second residence time with similar results (solid line with blue
diamonds), resulting in a predicted inlet temperature of 524 K at blowout, which is again much higher than we
would expect to be able to deliver N2O from the storage and supply system. Although it is possible to preheat the
N2O to some extent by regeneratively cooling the reactor with the inlet supply, there is no way to obtain the required
reactor operating temperatures indicated here without adding a significant amount of heat to the N2O flow from an
outside source.
4
The required minimum reactor volume was
obtained from the definition of PSR residence
time and the ideal gas law:
3500
Neat N2O
Sizing Pt.
Equation 2
Equation 3
Equation 4
τ res
m
=
m&
m& τ res = m =
PV
RT
m& τ res RT
Vreactor =
P
Exit Temperature, K
3000
Residence Time
0.1 Second
0.01 Second
2500
Neat N2O
P = 1 atm
1
2000
1500
Blowout
2
1000
3
500
4
0
0
0.2
0.4
0.6
0.8
1
Calculating the required reactor volume at
Fractional Simulation Time
this point from Equation 4 for a total flow of 1
Thermal decomposition of N2O: exit
lbm/sec indicated a minimum required reactor Figure 7.
temperature
vs
simulation time with decreasing inlet
volume of 7.3 cubic feet. While this could be
reduced by a factor of ten by operating the reactor temperature for two different residence times
at 10 atmospheres, we also knew that the actual
required volume will be several times that of the perfectly stirred reactor estimate.
These modeling results indicated that gas phase N2O decomposition would not only require very large reactor
volumes, but would also produce large amounts of NO, an endothermic product. These modeling results also agreed
with laboratory experiments where we demonstrated that gas phase N2O decomposition required very high
temperatures (ca. 1000°C) and produced very little oxygen. In addition, our initial catalyst results indicated that the
Figure 6: PSR exit compositions at the points shown in Figure 7.
5
reaction rate was much faster over a catalyst compared to gas phase decomposition and was probably limited only
by mass transfer rates. Consequently, we redesigned of the catalytic portion of the N2O reactor and sized it to
convert the entire N2O flow catalytically.
Figure 8 shows the conceptual layout of a counter flow heat exchanger/reactor consisting of two concentric
tubes surrounding a high-temperature capability Calrod-type electric heater. The outer tube is 1 inch OD while the
inner tube has a 0.75 inch outside diameter. Both of these tubes are located outside the Calrod heater, which has an
OD of 0.50 inch. The inner-most surfaces of the HX/R would be coated with our high temperature N2O
decomposition catalyst, allowing backside cooling of the catalyst and inner tube wall by the flow of cold N2O
entering the reactor. The Calrod will bring the catalyst up to its minimum operating temperature before N2O is
introduced at the low flow rate using a power level of about 500 watts for ~150 seconds to bring the Calrod up to a
temperature of 250°C needed to light off the decomposition reaction. When the low flow contacts this hot catalyst it
decomposes into N2 and O2, producing substantially higher temperatures. After the HX/R has come up to operating
temperature the high flow valve in the center of the system schematic of Figure 3 would be opened to provide the
full gas flow rate required by the barbotage injectors.
Assuming full conversion can be obtained and external heat losses are minimal, we can expect the temperature
N2O Liquid
Dryout Location
4
Catalyzed
Calrod
Begin Boiling
3
5
2
6
1
7
8
Hot N2/O2
Mixture
Figure 8. Catalytic N2O heat exchanger/reactor with model analysis stations.
of the N2/O2 mixture exiting the HX/R can be over 2400°F, which is beyond the melting point of typical super
alloys. However, less than full conversion combined with backside cooling of the inner tube with the incoming cold
N2O can prevent melting of the tube wall. The hexaaluminate catalyst can withstand the adiabatic decomposition
temperatures for periods on the order of hours without detrimental effect, but the Calrod may need to be made from
a refractory material such as silicon carbide, molydisilicide or one of the high temperature Kanthol alloys. Also, it
may be possible to use a standard Inconel-sheathed Calrod if radiant cooling of the surface to the inner tube wall is
large enough, or if the HX/R always operates in a transient mode.
If the catalyst is sufficiently active, all of the N2O reaching the surface is destroyed and the N2O mass fraction
profile in the catalyzed channel can be fairly easily estimated. This is done using standard heat transfer coefficient
correlations and the Chilton-Colburn analogy for the relationship between heat and mass transfer2:
Equation 5
Cf
2
= St Pr
2
3
= j H = j m = St m Sc
2
3
Where Cf is the skin friction coefficient, St is the Stanton number, Pr is the Prandtl number, jH is the ChiltonColburn ‘j-factor’ for heat transfer, jm is the j-factor for mass transfer, Stm is the mass transfer Stanton number, and
Sc is the Schmidt number. Further description of some of these non-dimensional parameters in terms of fluid
physical properties and their underlying meaning are given in Table 1.
6
Table 1. Non-dimensional groups important in heat and mass transfer
Group Definition Interpretation
StH
Pr
Stm
Sc
Le
Nu
h
ρuc p
cpμ
k
hm
u
μ
ρD AB
k
ρc p D AB
hD
k
Ratio of the heat transfer coefficient (h) to the freestream heat capacity flux.
Ratio of momentum to thermal diffusivity.
Ratio of species transport velocity to the wall to freestream convection velocity.
Ratio of momentum to the mass diffusivity of A in B (DAB).
Sc/Pr, ratio of thermal to mass diffusivity.
Nondimensionalized surface temperature gradient.
Using a standard turbulent flow heat transfer correlation in terms of the Nusselt number (Nu), such as DittusBoelter:
Nu D = 0.024 Re 0D.8 Pr 0.4 =
Equation 6
hD
k
allows us to estimate the heat transfer coefficient, h (W/m2-K). The Chilton-Colburn heat and mass transfer analogy
then gives the mass transfer coefficient as:
hm =
Equation 7
h
ρc p Le
2
3
Where hm is the mass transfer coefficient with units of m/s. The mass flux of species A to the catalyst surface is
then:
m& ′A′ = hm (ρ A,bulk − ρ A,surf )
Equation 8
The density of species A on the catalyst surface will be approximately zero if the reaction is mass transfer
limited. This, along with the mass balance for a differential element shown Figure 9, yields the first-order linear
ordinary differential equation governing the mass fraction of N2O along the length of the catalyzed channel of the
heat exchanger:
Equation 9
dy N 2O
dz
=−
hm Pwet y N 2O
uA flow
=−
hPwet y N 2O
2
ρc p Le 3 uA flow
=−
St h Pwet y N 2O
2
Le 3 A flow
As long as the Stanton and Lewis numbers are nearly constant over the length of the reactor, along with the wetted
perimeter and flow cross-sectional area, this ODE is separable and easily solved - indicating that the mass fraction of
N2O falls exponentially along the length of the catalyzed channel:
7
⎛ St P z
⎜
y N 2O (z ) = y N 2O , 0 exp⎜ − 2h wet
⎜ Le 3 A
flow
⎝
Equation 10
⎞
⎟
⎟⎟
⎠
Since the Stanton number usually has a value of about
0.002 and the Lewis number is usually not far from 1.0,
Equation 10 indicates that the N2O mass fraction profile
depends primarily on the length of the channel and the
.
ratio of the wetted perimeter to the flow area. This profile
r”
.
.
N2OPwetΔz
is therefore essentially independent of the mass flow rate
m(yN2O+ΔyN2O)
myN2O
or changes in the mixture temperature and composition
along the channel since these effects cancel each other out
in the nondimensional parameters. This also indicates that
Δz
most of the conversion occurs in the entrance to the
catalyzed channel where the backside cooling flow is at its
z
highest temperature, increasing the risk of burning or
Figure
9.
Mass transfer-limited performance
melting the inner heat exchanger/reactor tube at this point.
analysis
model.
The results of this analysis are shown in Figure 10,
showing predicted bulk fluid and inner tube wall
temperatures versus distance from the inlet (or exit) of the heat exchanger/reactor along with the predicted N2O mass
fraction profile along the catalyzed inner flow channel. This figure shows that compressed liquid N2O enters the
heat exchanger at a temperature of -60°F and is heated to 47°F by Station 2, where it reaches the saturation
temperature. Boiling and dryout occurs from Station 2 to 3, where the temperature remains fairly constant as all of
the liquid is converted to gas. At this point the gas temperature increases and reaches 662°F (350°C) by the time it
reaches station 4 and 5, which is the point where the flow is directed into the smaller inner tube and starts contacting
the catalyst. As pointed out in our experimental catalyst work, 350°C is hot enough to cause the N2O catalytic
decomposition reaction to occur rapidly. A maximum bulk fluid temperature of 2172°F was predicted at Station 6,
after which the bulk fluid temperature falls due to a decreasing reaction rate and heat absorbed by the incoming
liquid. An overall N2O decomposition efficiency of 87% with an exit temperature of 2050°F was predicted based on
the use of smooth tube walls. The predicted overall heat exchanger length was 28 inches using a 1” OD outer tube,
a ¾” OD inner tube, and a 0.5” diameter Calrod, with a weight of approximately 3.5 lbs. The overall heat duty is
266,000 Btu/hr with an average flux of nearly 600,000 Btu/ft2-hr on the inner tube wall. An overall total pressure
loss of 40 psi was predicted for the 500 psia exit static pressure design point.
The approximate global heat exchanger solution was checked locally by analyzing the complete thermal circuit
shown in Figure 11, which explicitly captures the chemical heat release on the catalyst surfaces as well as the
N2O HX/R Performance
N2O HX/R Performance
2500
1500
8
7
N2O Mass Fract. (Bulk Fluid)
Temperature, F
2000
6
87% N2O Conversion
tcat,i = 0.007"
1000
5
Bulk Fluid
Inner Tube Wall (e-Ntu)
500
4
2
3
0
1 0
5
1
5
10
15
20
25
30
0.9
0.24 lbm/s N 2O
Tin = -60 F
Tout = 2050 F
Pout = 500 psia
0.8
0.7
0.6
0.5
0.4
0.3
0.2
6
0.1
8
0
0
-500
Distance from Inlet or Exit, inches
7
5
10
15
20
25
30
Distance from Exit, inches
a
b
Figure 10. (a) Bulk fluid and inner tube wall temperatures versus distance from the HX/R inlet predicted
with the effective counterflow heat exchanger approach, and (b) the predicted N2O mass fraction profile
based on the mass transfer rate limitation
8
convective, conductive, and radiative heat transfer paths inside the HX/R. The rate of chemical heat release was
assumed to be mass transport limited in keeping with good catalytic reactor design practice, as well as the high
activation energy measured for the catalyst.
Local temperatures predicted for the inner tube wall and the Calrod surface using the thermal circuit of Figure
11 are shown in Figure 12. Fair agreement between this local heat transfer analysis and the initial global ε-Ntu
Rrad,o
Rrad,i
RCat
RBoil
T2
RBoil
RWall
T2W,o
T2W,i
R7
T7Cat
T7
R7
T7Rod
T7W
Bulk Fluid
Heating
q’Boil
q’Cat,tube
q’Cat,calrod
Figure 11. Thermal circuit for the prediction of local temperatures in the heat exchanger/reactor
approach is seen with respect to the inner tube wall temperature. The detailed circuit also allows us to predict the
Calrod surface temperature, which has a maximum value of 2485°F due to the lack of backside cooling plus the
catalytic surface reaction. This is beyond the melting point of most nickel-based superalloys, but is actually
obtainable with existing heater element materials. We could lower the temperature of this component to the bulk
fluid temperature by dispensing with its catalyst coating, but this step would also substantially lengthen the reactor
and would make the initial pre-operative heating of the unit more difficult.
IV.
Summary and Conclusions
Temperature, F
Overall this Phase I project was extremely successful as we confirmed the feasibility of using catalytic N2O
decomposition on board a scramjet-powered vehicle either as a pilot ignition system or as a source of gas for
barbotage fuel atomization. Our work developing and characterizing the N2O decomposition catalyst is reported in
our previous paper3. With respect to the N2O storage and delivery system and catalytic decomposition reactor, we
showed that a system sized to provide N2O for both barbotage and igniter torch operation for a nominal 100 pps
engine would require just over 27 lbs of N2O and
the storage tank would weigh approximately12
lbs constructed from titanium and 6 lbs if
N2O HX/R Predicted Temperatures
constructed from a carbon composite. We also
3000
87% N 2O Conversion
completed a conceptual design of a counter flow
tcat,i = 0.007"
catalytic heat exchanger/reactor that is about one
2500
inch in diameter by about 2.5 feet in length that
uses the energy generated from N2O
2000
5
6
decomposition to vaporize the liquid N2O and
1500
provide a hot O2/N2 mixture to the engine at about
7
8
2000°F. This heat exchanger also uses backside
1000
cooling from the liquid N2O to prevent internal
Inner Tube Wall (e-Ntu)
components from exceeding their respective
Inner Tube Wall (Local Circuit)
500
Calrod (Local Circuit)
temperature limits.
V.
0
Acknowledgements
The authors gratefully acknowledge the Air Force
SBIR office for funding this work under contract
number FA8650-09-M-2956. We also would like
to thank our contract monitor, Mr. David
0
5
10
15
20
25
30
Distance from Inlet or Exit, inches
Figure 12: Comparison of predicted tube wall and Calrod
surface temperatures obtained using global counterflow
heat exchanger and local thermal circuit approaches.
9
Buckwalter at Wright Patterson Air Force Base, for his thoughtful comments and direction over the course of this
project.
VI.
References
1
Hitch, B.D., Davidson, D.F., and Lynch, E.D., “Gaseous, Liquid, and Gelled Propellant Hypergolic Reaction
Mechanisms,” STTR Phase I Final Report No. RSLLC-F-7003, Reaction Systems, LLC, 19 March (2007).
2
Incropera, F.P. and DeWitt, D.P., “Fundamentals of Heat and Mass Transfer,” 4th ed., John Wiley & Sons, New
York, pg. 321 (1996).
3
Wickham, D.T., Hitch, B.D., and Logsdon, B., “Development and Testing of a High Temperature N2O
Decomposition Catalyst,” Paper No. AIAA-2010-xxxx presented at the 46th AIAA/ASME/SAE/ASEE Joint
Propulsion Conference & Exhibit, 25 - 28 Jul, Nashville, TN (2010).
10