Flow Field and Thermal Management Analysis of an
Armored Vehicle Engine Compartment
Robert F. Kunz
Penn State University Applied Research Laboratory, University Park, PA
16804. rfk@wt.arl.psu.edu
Nameer Salman
ICEM-CFD Engineering, Livonia, MI 48152, nsalman@icemcfd.com
Abstract
Computational Fluid Dynamics (CFD) analyses were performed for an armored tank engine compartment cooling flow. Large hybrid unstructured
meshes (2.5-3.0x106 cells) were constructed using the ICEM-CFD grid generator. The flow and convective heat transfer field were computed using an inhouse CFD code, NPHASE. The commercial software package,
RADTHERM was utilized to incorporate radiation heat transfer within the
simulations.
Two steady operating conditions and one engine-off cool-down transient
were analyzed. Specifically, the conditions analyzed were open throttle (hereafter OT), Tac Idle (TI) and engine off soakback (SB). OT and TI were run
with and without convection heat transfer employed in the radiation assessments to provide best-estimate and conservative peak temperature predictions
respectively. SB was run transiently using fixed heat transfer coefficients obtained from NPHASE analysis.
Results are presented for the simulations performed, with emphasis placed
on peak temperatures of several design critical elements.
Software Tools
ICEM-CFD (2000) is a commercial geometric modeling and mesh generation
package that has been widely employed in the automotive industry to accommodate the very complex geometries associated with underhood thermal management analysis.
The NPHASE CFD code was developed by the author and several colleagues and is described in detail in Kunz et. al. (2001), Antal et. al. (2000)
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R.F. Kunz and N. Salman
and Yu et. al. (2001). The code is fully unstructured and supports arbitrary
element types (the meshes employed here utilize hexahedra, tetrahedra, prisms
and pyramids). A parallel implicit, pressure-based segregated solution procedure is employed. The code can predict steady state and time dependent flows
and employs higher order temporal and spatial discretization. A range of
physical models are implemented in the code. Those implemented in the present work are:
• High Reynolds number k-e turbulence model
• Perfect gas compressibility (with buoyancy)
• Porous media
• Turbomachinery capability (including body force modeling).
• Specified temperature, heat flux, heat transfer coefficient and various specialized conjugate heating boundary conditions
NPHASE does not contain an on-board radiation heat transfer modeling
capability. For this, the commercial package RADTHERM
(http://www.thermoanalytics.com) was employed. This software is widely used
for underhood thermal management analysis in the automotive industry. As
described below, heat transfer coefficients and fluid temperatures, obtained
from NPHASE analyses, were imported into RADTHERM. Based on specified material properties, RADTHERM employs a view-factor based algorithm
to determine the convection-conduction-radiation heat balance on all solid
surfaces in the domain, resulting in the final temperature predictions of interest.
ICEM-CFD generates a “CGNS” file that includes all vertex, edge, face and
element data defining the hybrid unstructured mesh, as well as “patch family”
designations that can be used to define boundary conditions, as well as volume
family identifiers that can be used for localized element based treatments
within the flow solver (e.g., body forces within the fan). A CGNS file reader is
available for NPHASE, which accommodates this richness afforded by the
CGNS format.
The ICEM/NPHASE meshes employed for the CFD analyses carried out
here have just under 5.0x105 triangular wall faces. This very fine resolution is
consistent with fine grid requirements for the fluid-thermal CFD analysis.
However, this is far more than necessary for requisite accuracy in the
RADTHERM analysis, and would require processor weeks to even run in
RADTHERM (which is not currently parallelized). Accordingly, a procedure
was developed to “coarsen” the CFD surface so as to the reduce fidelity of
viewfactor and radiation patch simulation. This process involves: 1) Using
ICEM to coarsen surface meshes (while retaining good resolution on important parts), and to generate a Patran Neutral file defining the new surface
mesh, 2) Running the Thermoanalytics (vendors of RADTHERM) tool, mapconvbc, to “interpolate” (closest point) the fine mesh CFD surface solution
(heat transfer coefficients and film temperatures) onto the coarser model, and
to output a Patran Neutral file with these interpolated values, 3) Importing
these interpolated H, Tfilm in the RADTHERM analysis to define convection.
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387
Modeling Details
The general configuration and flow path of the engine compartment is shown
in figure 1. In figure 2, two views of the ICEM model are shown. Note that
for TI and OT an artificial exit extension was installed to accommodate the
highly recirculating fan exhaust flow within the cooling fan housing.
Grids
For OT and TI, a 2,951,279 element mesh was constructed. The mesh employed 2,804,579 tetrahedra and 146,700 prisms. A detail of the ICEM mesh
in the vicinity of the engine is shown in figure 3a, illustrating the complexity
and high geometric resolution of the model. Further indication of the high
resolution of the present mesh is observable in the part surface grid plots given
in figure 13. For the SB case, all elements downstream of the top of the heat
exchanger pack were deleted and the face on the top of the heat exchanger
pack itself was re-designated as a wall (see figures 1 and 8 for reference). This
resulted in a somewhat smaller mesh of 2,424,411 tetrahedra and 240 prisms.
By virtue of the relatively low Reynolds numbers encountered in the engine
compartment, tetrahedra were deemed adequate for resolution of most wall
layers. An average wall spacing (i.e., wall adjacent element volume centroid to
wall face centroid distance) for all wall faces was 1.6mm. This gave rise to an
average y+ value for all wall adjacent cells of 73 and 39 for OT and TI runs respectively, consistent with the high-Reynolds number shear and thermal wall
function turbulence modeling employed. For the SB case, the average value of
y+ was approximately 7. In all simulations, those elements for which y+ drops
below 10 employ a two-layer wall-function treatment consistent with the presence of the cell centroid in the laminar sublayer.
Domain Decomposition
ICEM outputs a CGNS file which is read into a sequence of front-end utilities
which, among other tasks, implements domain decomposition using the freely
available METIS (2001) partitioning software. All of the NPHASE simulations were performed on a LINUX cluster of 1GZ Pentium IV processors, using 24 processors. The 24 domain METIS partitioning for the OT and TI
simulations is shown in figure 3b.
General Flow Modeling Parameters
The following physical parameters were common to all analyses performed:
1. Perfect gas air: g=1.4, R=287 J/kg*oK
2. Thermal wall-functions for convection heat transfer (Prturbulent=0.91)
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3. Fixed exit pressure = 1 atm
4. Buoyancy terms included in momentum equations
5. Constant molecular viscosity, m=1.5x10-5 kg/m*s
All simulations were performed employing second order accurate convection and diffusion discretization. All simulations exhibited a several order of
magnitude residual reduction within 500-1000 iterations. Figure 4 shows a
representative residual and global mass conservation history for an OT simulation. In these analyses, residuals all eventually begin to level off due to the inherent small amplitude unsteadiness in regions of very low flow. Nevertheless,
global mass conservation and various solution monitoring scalars (incremental
pressure and temperature drops) converged to three significant digits within
2000 iterations. All simulations were run out to at least this level of convergence.
OT and TI NPHASE Modeling
For all cases, an inlet ambient temperature was specified. An inflow density
was specified corresponding to this temperature and assumed standard atmospheric pressure. Inlet axial velocities were specified based on inlet crosssectional flow area, density and prescribed mass flow rates. Inlet values of turbulence kinetic energy and turbulent dissipation rate were determined based
on an assumed turbulence intensity level of 3% and length scale of 5% of the
inlet duct width. In each of these three cases the pressure at the exit plane was
specified as ambient.
OT and TI simulations were carried out in similar fashion to one another,
with the only differences being the inflow velocity values, and the values of
fixed engine temperatures assigned as illustrated in figure 5 (values adapted
from engine manufacturer). For OT and TI, the transmission was modeled in
a fashion that accommodates the nearly constant transmission fluid temperature on the inside of its housing. As illustrated in figure 6, a locally onedimensional conjugate heat transfer condition is applied by considering the
wall thickness and material conductivity, with the inner wall temperature set
to the design operating transmission oil temperature. For OT and TI, the airbox was modeled in a fashion that accommodates the near-ambient temperature flow of engine air on the inside of the box. As illustrated in figure 7, a locally one-dimensional conjugate heat transfer condition is applied by
considering the wall thickness and material conductivity. The treatment here
differed from that applied to the transmission in that rather than a fixed inner
wall temperature, an inner wall film temperature (ambient) was defined, and a
heat transfer coefficient was determined from a standard Nusselt number correlation (Kreith, 1973):
Nu ≡
HL
= .029 Re.8 Pr1 / 3
k
(1)
where L is taken as half the airbox duct length, k is the thermal conductivity of
air at the inlet temperature, Pr is the Prandtl number of air (.72), and the Rey-
Flow Field and Thermal Management Analysis
389
nolds number is determined based on L and the bulk velocity of the engine air
based on nominal air-box cross-sectional area and given engine air mass flow
rates at OT and TI respectively.
The exhaust duct is treated as a constant temperature surface for OT and
TI. A number of the compartment parts are “2-sided”, that is, they have flow
on both sides. These “sheet-metal” pieces are treated as infinitely thin internal
boundaries. A conjugate heat transfer boundary condition is applied for these
pieces where heat flux and wall temperature are constant on each side of such
faces. These pieces include the cooling fan housing and the heat exchanger
housings. Adiabatic boundary conditions were employed for all other surfaces.
Bulk modeling is employed for the heat exchangers and cooling fan. As illustrated in figure 8, design values of pressure drop and temperature rise are
available for each of the three heat exchangers. Each of these devices is meshed
independently using tetrahedra, such that each element is uniquely defined as
being within one of the three coolers. Body forces are added to the momentum
equations to establish the correct pressure drop and flow straightening. Specifi˙ localx (Vref /2) , where
cally, in the flow direction a force is added: Fy = m
& local is the local cell’s mass flow rate and x is a loss coefficient. Generally, x is
m
determined from empirical correlations for loss mechanisms associated with
inlet, core, acceleration, and exit losses, using a suitable definition of reference
& local are known, x was iteravelocity, Vref. In the present work, since Dp and m
tively determined through several solution restarts such that the desired Dp was
matched to requisite accuracy. In the two cross flow directions, the loss factor
was increased by a factor of 10 to “straighten” the flow. This is illustrated with
a velocity vector plot in figure 9. The accuracy with which the design pressure
drops were matched is shown in figure 11.
Heat addition to the air flow in the three coolers was accommodated in a
consistent fashion. Specifically, local heat addition sources were added to the
˙ localEC PDTHX ,
enthalpy equation based on a local energy balance: q local = m
where DTHX ≡ Tair - Tcoolant , CP is the specific heat of the air and E is an unknown cell “efficiency”. For each exchanger we have available approximate
values for qtotal and D THX, which when substituted into the equation above
yield estimates for E for each of the three coolers. As with the loss coefficients,
E was then iteratively refined through several solution restarts such that the desired DT was matched to requisite accuracy. The accuracy with which the design temperature rises were matched is shown in figure 12.
The fan was also modeled using a bulk representation. Specifically, approximate machine rotation rates and pressure rise across the fan were available. Since the mass flow rate and flow path are fixed, a suitable body force
distribution could be designed. A design code at Penn State Applied Research
Lab was utilized to generate tangential, axial and radial forces through the meridional plane of the fan. These were then distributed onto the NPHASE grid
using bilinear interpolation. Elements of this procedure are illustrated in figure
10. The accuracy with which the design pressure rises were matched is shown
in figure 11.
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Soakback NPHASE Modeling
Soakback is the engine cool-down transient. This cool-down process has the
potential to be design limiting since at engine shut off, all engine and transmission cooling flow ceases. The large thermal capacitance of the engine and
transmission can lead to local increases in radiative heating within the compartment since convective cooling within the compartment will be significantly diminished. The cool-down transient has a time scale on the order of an
hour, whereas at engine shut-down, convection cooling is lost (or with a small
auxiliary soakback fan operating, greatly reduced) in a matter of seconds. Accordingly, this disparity in scales led to a soakback modeling approach involving a steady state analysis with the small soakback fan flow to provide heat
transfer coefficients for the transient radiation analysis. This approximation
should be valid, since heat transfer coefficients will be a weak function of component temperatures, and, as indicated above, little accuracy will be lost by neglecting the short duration convective heat transfer transient right at engine
shut-down.
Engine shut-down is specified to occur while the engine is at TI. Accordingly, NPHASE was run with all boundary conditions as specified above for
TI, with the exception of inlet mass flow rate, which was set to a small value
corresponding to a notional soakback cooling fan. Also, no flow was admitted
through the heat exchangers for SB operation, as described above.
OT and TI RADTHERM Modeling
The engine compartment contains a variety of materials including metals and
alloys, insulation and lubrication oils. In general, the emmissivity, conductivity, density and specific heat of these materials had to be obtained. Each of the
surfaces in the model had to be assigned to one of these materials and given a
nominal thickness for conduction heat balance purposes.
The heat transfer coefficients and film temperatures predicted by NPHASE
were input into RADTHERM as described in the “Software Tools” section
above. The ICEM coarsening process left several of the key engine compartment parts under-resolved (i.e., too few triangle elements), so ICEM’s refinement feature was then applied to hull, engine, airbox, and transmission components to recover requisite surface resolution. Figure 13 illustrates example
NPHASE and RADTERM surface meshes for the airbox and generator components.
RADTHERM is brought up and the model imported through the
PATRAN neutral file generated from the NPHASE solution file interpolated
onto the coarsened surface mesh (using the mapconvbp utility provided by
Thermoanalytics, Inc., vendors of RADTHERM). The RADTHERM display
appears as shown in figure 14. Each of the 91 parts was then assigned appropriate material definitions and thickness as indicated above. All parts except
the engine components and exhaust duct are specified to employ the imported
H and Tfilm values. Engine components are assigned fixed temperatures per
Flow Field and Thermal Management Analysis
391
figure 5. The exhaust duct is assigned a fixed temperature of Tref for both OT
and TI.
Appropriate values for the “back” (i.e., not-flow-facing) side of every part
must also be specified. For the 2-sided “sheet-metal” parts, H and Tfilm values
are available from the NPHASE solution and are employed for each side. For
the steady state OT and TI simulations, the engine and exhaust duct back side
boundary conditions are not important since the engine is almost a completely
closed surface and the outer surface temperature is specified. Accordingly,
there H is set = 0. The transmission back side is set to a constant temperature
of T/Tref = 0.75, consistent with the NPHASE analysis by setting Tfilm/Tref =
0.75 and H = 1000 (i.e., so high that Twall = Tfilm). The airbox backside H and
Tfilm are specified as in the NPHASE analysis discussed above. The generator
backside is specified as adiabatic. For all hull pieces, including the bulkhead
and engine cover, H=0 is specified on the backside, which still allows radiative
transfer away from the engine compartment. All other pieces are set to adiabatic on the backside.
Emmissivities are set for the front and back faces of all parts. These are set
based on the material of the part, or, if the part is painted, an appropriate emmissivity corresponding to the paint is set. All components are painted except
the compartment cover, bulkhead, rear door, and the engine itself.
For OT and TI, RADTHERM was run as specified above for 400 iterations. This was sufficient to converge the runs to within 0.05oF. For comparison, another pair of OT and TI runs was performed with all internal engine
convection cooling “turned-off”, that is, all heat transfer coefficients imported
from NPHASE were overwritten as equaling zero. These “no-cooling-flow”
runs were performed to provide upper bound conservative estimates on peak
component temperatures.
Soakback RADTHERM Modeling
As discussed above, the soakback runs were performed transiently using the H
and Tfilm field obtained from steady state SB NPHASE simulation. In order to
accommodate the critical thermal inertia physics associated with soakback, several changes were made to RADTHERM part specifications. These changes
involved the engine, transmission, recuperator, and exhaust duct. Specifically,
the transmission was redefined as a 3-layer part as illustrated in figure 15a. The
reasoning for this is as follows: For soakback, the transmission was originally
modeled as a large chunk of alloy, of nominal thickness to match the dry
weight of the transmission and thereby mimic its thermal capacitance. This
non-conservative assumption allowed all incident radiative flux to be conducted away very efficiently into the transmission, thereby rapidly “smoothing” hotspots on the surface of the transmission since the thermal conductivity
of the alloy is high. Moving to the 3-layer model is more physically realistic
because the presence of the air and transmission oil layers (above and below
the sump line, respectively, as illustrated in figure 15b), that exist in the real
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configuration, greatly inhibits conduction normal to the transmission surface,
while the thick third layer still accommodates the thermal capacitance of the
system. The nominal core thickness of the transmission was determined using
an estimate for the volume of the transmission, the transmission’s known dry
weight, and the density of the alloy.
A similar 3-layer model is employed for the engine, as illustrated in figure
16. The nominal core thickness of the engine was determined using an estimate for the volume of the engine, the engine’s known dry weight, and the
density of steel. The Tinit/Tref = 1.24 core initial temperature of the engine was
estimated based on engine output, fuel consumption and engine air flow rate.
The recuperator and exhaust duct were also treated as three-layer parts with
outer surfaces of painted sheet metal, a layer of insulation and an inner layer of
sheet metal. A zero heat transfer coefficient was set on the inside of the recuperator and exhaust duct for soakback.
Except as noted above for the engine, all parts were given an initial temperature distribution from the steady TI simulation.
Results
As already indicated, three sets of runs were made: OT, TI, and SB. In this
section we summarize the results obtained. First, details of the CFD simulations are presented, followed by the RADTHERM results.
OT and TI NPHASE Results
Figures 17 through 27 contain elements of the NPHASE simulations performed. In figure 17, selected streamlines emanating from the inlet are shown
for OT. These are shaded by temperature. A temperature isosurface (T/Tref =
0.82) is also plotted. Though most streamlines are seen to follow a fairly direct
path from the inlet to the heat exchanger pack, a good deal of the flow is seen
to divert to either side, and this gives rise to a fairly complex compartment
flow field. The streamlines that transit the compartment toward the heat exchanger pack, come in close proximity to the exhaust duct and are thereby
heated through convection from the exhaust duct. The T/Tref = 0.82 isosurface
is seen to envelop the engine, heat exchanger pack, and exhaust duct in this
view, as expected. The strongly swirling exit flow induced by the fan is also observable in this view (2 counter-rotating vortices).
Figures 18 shows two views of numerous streamlines seeded at the inlet
and/or the top of the heat exchanger for OT. A very complex threedimensional flow field is seen to exist throughout the engine compartment.
Figures 19 and 20 show near engine views of the predicted velocity field for
OT and TI. The velocity vectors shown are projected into the viewing plane
and are shaded by magnitude of total velocity. It is observed that in the immediate proximity of the hottest components of the engine, velocities of V/Vref >
0.7 are encountered for OT (compare with inlet bulk velocity of V/Vref = 2.4).
Flow Field and Thermal Management Analysis
393
Peak values for TI are somewhat smaller than OT, as expected. Figure 21
shows two cross-sectional views of density contours, illustrating the weak but
non-negligible thermally induced perfect gas density variations within the engine compartment.
Soakback NPHASE Results
Elements of the NPHASE soakback solutions are provided in figures 22-27. In
figure 22, predicted contours of pressure are shown, for a compartment crosssection, illustrating the gravity head most responsible for driving the flow in
this case. Figure 23 illustrates the complex nature of the buoyancy driven flow
within the engine compartment. There, numerous streamlines, shaded by
temperature, are plotted. In general, the entire compartment is subject to free
convection flow. Temperature contours in the vicinity of the generator are
plotted in figure 24. Clearly, free convection from the hot exhaust duct impacts the engine cover and generator. Figure 25 shows a similar plot for a slice
taken in a nominal engine-airbox plane. Here free convection heat transfer
from the engine burner region and recuperator are seen to clearly impact the
airbox. Predicted velocity vectors in a cut plane through the transmission and
exhaust duct are shown in figure 26. These vectors are resolved into the x-y cut
plane and shaded by the magnitude of the resultant velocity in that plane,
|Vxy|. Values of |Vxy |/Vref @ 0.05 are observed compared to inlet velocities of
Vinlet/Vref = 2.4, 1.52 and 0.034 for OT, TI and SB, indicating a weak but nonnegligible flow. Free-convection induced velocity vectors in the y-z plane in
the vicinity of engine and transmission are shown in figure 27, illustrating significant free convection in the hot region between transmission and engine.
OT and TI RADTHERM Results
Selected RADTHERM results for OT are shown in figures 28 and 29. There,
temperature contours are shown for the RADTHERM runs carried out with
the NPHASE H and Tfilm and with no convection cooling for the engine compartment cover, generator and airbox. These results, as well as those of other
engine components and for the TI operating condition illustrate that significant hot spots can occur opposite the engine and transmission surfaces. Also
found is that convection cooling provided significant reduction in these hot
spot peak temperatures as expected.
Soakback RADTHERM Results
As mentioned above, the soakback transient could be limiting for several critical engine components due to the sudden loss of internal convective cooling to
the engine and transmission thermal masses. The desire to capture this was ac-
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R.F. Kunz and N. Salman
commodated in the RADTHERM modeling strategy outlined above.
RADTHERM was run in transient mode for the SB case. After some experimentation, it was found that time steps of 30 seconds, and per-time-step convergence tolerances of 0.5oC yielded requisite numerical accuracy. The results
in figures 30 through 36 summarize these runs.
Figure 30 illustrates the thermal history of the airbox. It is observed that the
temperature of the airbox hot-spot adjacent to the recuperator increases before
peaking after t/tref @ 4, and then dropping off. The figure includes a time history of the peak airbox temperature, including a comparison to an
NPHASE+RADTHERM simulation with no soakback flow. It is seen that
that the impact of soakback cooling flow is very small.
Figure 31 illustrates the thermal history of the bulkhead. The temperature
of the bulkhead hot spot opposite the engine increases only very slightly,
peaking after t/tref @ 1.5 and then dropping off. The time history of the peak
temperature shown in the figure demonstrates that the impact of soakback
cooling flow is small, however, it is seen to actually increase the bulkhead peak
temperature somewhat. The physical explanation for this counter-intuitive
finding can be gleaned from figure 32. There, predicted total velocity vectors
are plotted in an x-y plane midway between the smallest recuperator-bulkhead
gap. Clearly, the soakback flow has a significant effect on the local velocity
field there, with significant cross flow arising compared to the principally vertical buoyancy dominated field without soakback flow. Comparison with the
lower part of the figure shows that with the soakback fan on, cross flow gives
rise to higher local convection temperatures by virtue of the transport of hotter
fluid into the hot-spot region. This in turn gives rise to higher film temperatures when interpolated to RADTHERM. Accordingly convection
flux, q ¢¢ = H(Twall - Tfilm ) , is lower for the case without soakback flow.
Figure 33 illustrates the thermal history of the engine compartment cover.
The temperature of the cover hot spot opposite the exhaust duct increases,
peaking at t/tref @ 4 and then dropping off. Soakback cooling reduces the peak
predicted engine compartment cover temperature by a negligible amount.
The thermal history of the engine is illustrated in figure 34. There it is seen
that the engine cools quite rapidly from its initial condition (steady TI operation). The figure illustrates that all components of the engine+recuperator+exhaust duct cool with the exception of those parts in immediate contact with the burner, which retain higher temperatures through
conduction from the burner region. The peak engine temperature decay is
rapid in the first t/tref @ 10 after engine shut down, but levels off to a much
slower decay after that. Simulations with and without soakback fan cooling are
virtually indistinguishable for the engine on this scale. For comparison, in figure 34, the original engine model prediction (no-multilayer treatment – assumed pure metal) is included. It is observed that the improved physical modeling of the engine gives rise to a higher peak (and, as it were, bulk)
temperature in the slow heat decay region beyond t/tref @ 10.
The generator soakback thermal history is illustrated in figure 35. The hot
spot on this part is seen to decay monotonically during soakback. This is due
to the comparatively low thermal mass of the exhaust duct which itself cools
Flow Field and Thermal Management Analysis
395
quickly (refer figure 34). The simulations with and without soakback cooling
are very similar.
Figure 36 shows views of the transmission over the t/tref = 30 after engine
shut down. Clearly observable is the decay in the engine-burner-facing hot
spot. Also observable is the relatively rapid cooling experienced in the sump
region compared with the transmission housing above the sump line. This of
course is due to the modeling discussed above. The figure also includes the
time history of the decay of the two transmission hot spots. The temperature
decays monotonically from engine shutdown for both hot spots, and simulations with and without soakback cooling are very similar.
Conclusions
The analyses carried out under the present effort involved the application of stateof-the-art grid generation, CFD and radiation heat transfer analysis software. Large
high quality hybrid unstructured meshes were constructed using ICEM-CFD. The
NPHASE CFD code was employed to solve the flow and convective thermal transport. The commercial radiation analysis package, RADTHERM, was employed in
assessing the radiation dominated peak temperatures in the system.
References
Kunz, R.F., Yu, W.S., Antal, S.P., Ettorre, S.M. “An Unstructured Two-Fluid
Method Based on the Coupled Phasic Exchange Algorithm,” AIAA Paper 2001-2672,
Proc. 15th AIAA Computational Fluid Dynamics Conference, Anaheim, CA, June,
2001.
Antal, S.P., Ettorre, S.M., Kunz, R.F., Podowski, M.Z. “Development of a NextGeneration Computer Code for the Prediction of Multicomponent Multiphase
Flows,” presented at the International Meeting on Trends in Numerical and Physical
Modeling for Industrial Multiphase Flow, Cargese, France, September 27, 2000.
Yu, W.S., Kunz, R.F., Antal, S.P., Ettorre, S.M. “Unstructured Rotor Stator
Analysis of Axial Turbomachinery Using a Pressure-Based Method”, ASME Paper presented at the International Mechanical Engineering Congress and Exposition, New
York, NY, November, 15, 2001.
Kreith, Frank Principles of Heat Transfer, Harper and Row, New York, 1973.
Schlicting, Hermann Boundary Layer Theory, McGraw-Hill, New York, 1968.
Touloukian, Y.S., DeWitt, D.P. (eds.) Thermal Radiative Properties: Metallic
Elements and Alloys, Vol. 7 of Thermophysical Properties of Matter, Plenum Press,
New York, 1970.
METIS Version 4.0 documentation, 2001.
ICEM CFD Software User Manual v4.1, ICEM CFD Engineering, Berkeley, CA,
2000.
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R.F. Kunz and N. Salman
Fig. 1. Sketch of engine compartment configuration and cooling air flow path for OT & TI.
Fig. 2. Two views of the ICEM model.
Fig. 3. a) View of the ICEM engine compartment model illustrating the complexity of the configuration. b) View of 24 domain METIS partitioning for engine compartment model.
Flow Field and Thermal Management Analysis
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mass flow
pressure
velocity
Fig. 4. Representative NPHASE convergence history for engine compartment analyses. Shown
are pressure and velocity residuals, and global mass flow rate through compartment (outflowinflow).
Fig. 5. Engine T/Tref map, specified in NPHASE for OT/TI operation.
Fig. 6. Illustration of transmission components and specialized conjugate heating boundary
condition employed in NPHASE.
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Fig. 7. Illustration of airbox and specialized conjugate heating boundary condition employed in
NPHASE.
Fig. 8a. Illustration of three heat exchangers and design specifications for their respective pressure drop, Dp/Dpref, at OT & TI operation.
Fig. 8b. Illustration of three heat exchangers and design specifications for their respective temperature rise, DT/DTref, at OT & TI operation.
Flow Field and Thermal Management Analysis
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Fig. 9. Predicted velocity vectors just above and through the three heat exchangers, illustrating
the straightening due to resistance modeling employed.
Fig. 10. Elements of fan body-force modeling employed. a) View of fan vicinity, b) output of
throughflow code.
Fig. 11. Errors in NPHASE vs. design pressure changes (Dp NPHASE-DpDESIGN) across heat exchangers and fan.
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R.F. Kunz and N. Salman
Predicted Temperature Contours
Fig. 12. Errors in NPHASE vs. design temperature changes
heat exchangers.
(DTNPHASE-DTDESIGN) across
Fig. 13. NPHASE and RADTHERM surface meshes for generator and airbox.
Fig. 14. RADTHERM interface with view of engine compartment model (grey-scale).
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Fig. 15a. 3-layer transmission part treatment for soakback RADTHERM analyses.
Fig. 15b. ICEM-CFD family split employed for the transmission to distinguish between regions
with sump oil and those without under soakback conditions.
Fig. 16. Sketch of 3-layer engine part treatment for soakback RADTHERM analyses.
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Fig. 17. Elements of OT NPHASE simulation. Streamlines shaded by temperature, and temperature (T/Tref = 0.82) iso-surface (lighter surface enshrouding engine region).
Fig. 18. Selected streamlines from OT NPHASE simulation.
Fig. 19. Rear-view in-plane velocity vectors, shaded by velocity magnitude (scale in V/Vref), for
OT (left) and TI (right) NPHASE simulations.
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Fig. 20. Top-view in-plane velocity vectors, shaded by velocity magnitude (scale in V/Vref), for
OT (left) and TI (right) NPHASE simulations.
Fig. 21. Two cross-sectional views of density contours (r/rambient), illustrating weak thermally
induced perfect gas density variations within the engine compartment.
Fig. 22. NPHASE SB simulation. Pressure contours, illustrating weak gravitational head rise
within engine compartment.
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Fig. 23. NPHASE SB simulation. Streamline field illustrating the complex thermally induced
free-convection field.
Fig. 24. NPHASE SB simulation. T/Tref contours in vicinity of generator and exhaust duct.
Fig. 25. NPHASE SB simulation. T/Tref contours in vicinity of engine and air box.
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Fig. 26. NPHASE SB simulation. Free-convection induced velocity vectors in x-y plane in vicinity of transmission, generator and exhaust duct.
Fig. 27. NPHASE SB simulation. Free-convection induced velocity vectors in y-z plane in vicinity of engine and transmission.
Fig. 28. RADTHERM OT temperature predictions on engine compartment cover. With
NPHASE H,Tfilm convection values (left) and no-convection (right).
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Fig. 29. RADTHERM OT temperature predictions on generator and on airbox. With
NPHASE H,Tfilm convection values (left) and no-convection (right).
Fig. 30. Sequence of predicted airbox surface temperature fields for SB case. Temperature contour range is from 120oF to 240oF. t/tref=0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peak predicted airbox surface temperature.
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Fig. 31. Sequence of predicted bulkhead surface temperature fields for SB case. Temperature
contour range is from 120oF to 225oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peak
predicted bulkhead surface temperature.
Fig. 32. NPHASE predicted total velocity vectors in an x-y plane midway between smallest recuperator-bulkhead gap for SB cases. NPHASE predicted temperature contours (T/Tref) on
bulkhead for SB cases.
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Fig. 33. Sequence of predicted engine compartment cover surface temperature fields for SB case.
Temperature contour range is from 115oF to 185oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time
history of peak predicted engine compartment cover surface temperature.
Fig. 34. Sequence of predicted engine and exhaust duct surface temperature fields for SB case.
Temperature contour range is from 120oF to 550oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time
history of peak predicted engine surface temperature.
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409
Fig. 35. Sequence of predicted generator surface temperature fields for SB case. Temperature
contour range is from 120oF to 290oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peak
predicted generator surface temperature.
Fig. 36. Sequence of predicted transmission surface temperature fields for SB case. Temperature
contour range is from 120oF to 280oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peak
predicted transmission surface temperature. Hot spots considered here are directly across from
the engine burner (left) and directly attached to the engine (right).