Flow Field and Thermal Management Analysis of an Armored ...extras.springer.com/2004/978-3-642-53586-4/385.pdf · based on nominal air-box cross-sectional area and given engine air

Flow Field and Thermal Management Analysis of anArmored Vehicle Engine Compartment

Robert F. Kunz

Penn State University Applied Research Laboratory, University Park, PA16804. [email protected]

Nameer Salman

ICEM-CFD Engineering, Livonia, MI 48152, [email protected]

Abstract

Computational Fluid Dynamics (CFD) analyses were performed for an ar-mored tank engine compartment cooling flow. Large hybrid unstructuredmeshes (2.5-3.0x106 cells) were constructed using the ICEM-CFD grid gen-erator. The flow and convective heat transfer field were computed using an in-house CFD code, NPHASE. The commercial software package,RADTHERM was utilized to incorporate radiation heat transfer within thesimulations.

Two steady operating conditions and one engine-off cool-down transientwere analyzed. Specifically, the conditions analyzed were open throttle (here-after OT), Tac Idle (TI) and engine off soakback (SB). OT and TI were runwith and without convection heat transfer employed in the radiation assess-ments to provide best-estimate and conservative peak temperature predictionsrespectively. SB was run transiently using fixed heat transfer coefficients ob-tained from NPHASE analysis.

Results are presented for the simulations performed, with emphasis placedon peak temperatures of several design critical elements.

Software Tools

ICEM-CFD (2000) is a commercial geometric modeling and mesh generationpackage that has been widely employed in the automotive industry to accom-modate the very complex geometries associated with underhood thermal man-agement analysis.

The NPHASE CFD code was developed by the author and several col-leagues and is described in detail in Kunz et. al. (2001), Antal et. al. (2000)

386 R.F. Kunz and N. Salman

and Yu et. al. (2001). The code is fully unstructured and supports arbitraryelement types (the meshes employed here utilize hexahedra, tetrahedra, prismsand pyramids). A parallel implicit, pressure-based segregated solution proce-dure is employed. The code can predict steady state and time dependent flowsand employs higher order temporal and spatial discretization. A range ofphysical models are implemented in the code. Those implemented in the pre-sent 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 spe-cialized conjugate heating boundary conditions

NPHASE does not contain an on-board radiation heat transfer modelingcapability. For this, the commercial package RADTHERM(http://www.thermoanalytics.com) was employed. This software is widely usedfor underhood thermal management analysis in the automotive industry. Asdescribed below, heat transfer coefficients and fluid temperatures, obtainedfrom NPHASE analyses, were imported into RADTHERM. Based on speci-fied material properties, RADTHERM employs a view-factor based algorithmto determine the convection-conduction-radiation heat balance on all solidsurfaces in the domain, resulting in the final temperature predictions of inter-est.

ICEM-CFD generates a “CGNS” file that includes all vertex, edge, face andelement data defining the hybrid unstructured mesh, as well as “patch family”designations that can be used to define boundary conditions, as well as volumefamily identifiers that can be used for localized element based treatmentswithin the flow solver (e.g., body forces within the fan). A CGNS file reader isavailable for NPHASE, which accommodates this richness afforded by theCGNS format.

The ICEM/NPHASE meshes employed for the CFD analyses carried outhere have just under 5.0x105 triangular wall faces. This very fine resolution isconsistent with fine grid requirements for the fluid-thermal CFD analysis.However, this is far more than necessary for requisite accuracy in theRADTHERM analysis, and would require processor weeks to even run inRADTHERM (which is not currently parallelized). Accordingly, a procedurewas developed to “coarsen” the CFD surface so as to the reduce fidelity ofviewfactor and radiation patch simulation. This process involves: 1) UsingICEM to coarsen surface meshes (while retaining good resolution on impor-tant parts), and to generate a Patran Neutral file defining the new surfacemesh, 2) Running the Thermoanalytics (vendors of RADTHERM) tool, map-convbc, to “interpolate” (closest point) the fine mesh CFD surface solution(heat transfer coefficients and film temperatures) onto the coarser model, andto output a Patran Neutral file with these interpolated values, 3) Importingthese interpolated H, Tfilm in the RADTHERM analysis to define convection.

Flow Field and Thermal Management Analysis 387

Modeling Details

The general configuration and flow path of the engine compartment is shownin figure 1. In figure 2, two views of the ICEM model are shown. Note thatfor TI and OT an artificial exit extension was installed to accommodate thehighly recirculating fan exhaust flow within the cooling fan housing.

Grids

For OT and TI, a 2,951,279 element mesh was constructed. The mesh em-ployed 2,804,579 tetrahedra and 146,700 prisms. A detail of the ICEM meshin the vicinity of the engine is shown in figure 3a, illustrating the complexityand high geometric resolution of the model. Further indication of the highresolution of the present mesh is observable in the part surface grid plots givenin figure 13. For the SB case, all elements downstream of the top of the heatexchanger pack were deleted and the face on the top of the heat exchangerpack itself was re-designated as a wall (see figures 1 and 8 for reference). Thisresulted in a somewhat smaller mesh of 2,424,411 tetrahedra and 240 prisms.By virtue of the relatively low Reynolds numbers encountered in the enginecompartment, tetrahedra were deemed adequate for resolution of most walllayers. An average wall spacing (i.e., wall adjacent element volume centroid towall face centroid distance) for all wall faces was 1.6mm. This gave rise to anaverage y+ value for all wall adjacent cells of 73 and 39 for OT and TI runs re-spectively, consistent with the high-Reynolds number shear and thermal wallfunction turbulence modeling employed. For the SB case, the average value ofy+ was approximately 7. In all simulations, those elements for which y+ dropsbelow 10 employ a two-layer wall-function treatment consistent with the pres-ence of the cell centroid in the laminar sublayer.

Domain Decomposition

ICEM outputs a CGNS file which is read into a sequence of front-end utilitieswhich, among other tasks, implements domain decomposition using the freelyavailable METIS (2001) partitioning software. All of the NPHASE simula-tions were performed on a LINUX cluster of 1GZ Pentium IV processors, us-ing 24 processors. The 24 domain METIS partitioning for the OT and TIsimulations 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*oK2. Thermal wall-functions for convection heat transfer (Prturbulent=0.91)


3. Fixed exit pressure = 1 atm4. Buoyancy terms included in momentum equations5. Constant molecular viscosity, m=1.5x10-5 kg/m*s

All simulations were performed employing second order accurate convec-tion and diffusion discretization. All simulations exhibited a several order ofmagnitude residual reduction within 500-1000 iterations. Figure 4 shows arepresentative residual and global mass conservation history for an OT simula-tion. In these analyses, residuals all eventually begin to level off due to the in-herent small amplitude unsteadiness in regions of very low flow. Nevertheless,global mass conservation and various solution monitoring scalars (incrementalpressure and temperature drops) converged to three significant digits within2000 iterations. All simulations were run out to at least this level of conver-gence.

OT and TI NPHASE Modeling

For all cases, an inlet ambient temperature was specified. An inflow densitywas specified corresponding to this temperature and assumed standard atmos-pheric pressure. Inlet axial velocities were specified based on inlet cross-sectional flow area, density and prescribed mass flow rates. Inlet values of tur-bulence kinetic energy and turbulent dissipation rate were determined basedon an assumed turbulence intensity level of 3% and length scale of 5% of theinlet duct width. In each of these three cases the pressure at the exit plane wasspecified 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 offixed engine temperatures assigned as illustrated in figure 5 (values adaptedfrom engine manufacturer). For OT and TI, the transmission was modeled ina fashion that accommodates the nearly constant transmission fluid tempera-ture on the inside of its housing. As illustrated in figure 6, a locally one-dimensional conjugate heat transfer condition is applied by considering thewall thickness and material conductivity, with the inner wall temperature setto the design operating transmission oil temperature. For OT and TI, the air-box was modeled in a fashion that accommodates the near-ambient tempera-ture flow of engine air on the inside of the box. As illustrated in figure 7, a lo-cally one-dimensional conjugate heat transfer condition is applied byconsidering the wall thickness and material conductivity. The treatment herediffered from that applied to the transmission in that rather than a fixed innerwall temperature, an inner wall film temperature (ambient) was defined, and aheat transfer coefficient was determined from a standard Nusselt number cor-relation (Kreith, 1973):

3/18. PrRe029.=≡k

HLNu

(1)

where L is taken as half the airbox duct length, k is the thermal conductivity ofair at the inlet temperature, Pr is the Prandtl number of air (.72), and the Rey-

nolds number is determined based on L and the bulk velocity of the engine airbased on nominal air-box cross-sectional area and given engine air mass flowrates at OT and TI respectively.

The exhaust duct is treated as a constant temperature surface for OT andTI. A number of the compartment parts are “2-sided”, that is, they have flowon both sides. These “sheet-metal” pieces are treated as infinitely thin internalboundaries. A conjugate heat transfer boundary condition is applied for thesepieces where heat flux and wall temperature are constant on each side of suchfaces. These pieces include the cooling fan housing and the heat exchangerhousings. Adiabatic boundary conditions were employed for all other surfaces.

Bulk modeling is employed for the heat exchangers and cooling fan. As il-lustrated in figure 8, design values of pressure drop and temperature rise areavailable for each of the three heat exchangers. Each of these devices is meshedindependently using tetrahedra, such that each element is uniquely defined asbeing within one of the three coolers. Body forces are added to the momentumequations to establish the correct pressure drop and flow straightening. Specifi-cally, in the flow direction a force is added: Fy = ˙ m localx(Vref /2) , where

localm& is the local cell’s mass flow rate and x is a loss coefficient. Generally, x isdetermined from empirical correlations for loss mechanisms associated withinlet, core, acceleration, and exit losses, using a suitable definition of referencevelocity, Vref. In the present work, since Dp and localm& are known, x was itera-tively determined through several solution restarts such that the desired Dp wasmatched to requisite accuracy. In the two cross flow directions, the loss factorwas increased by a factor of 10 to “straighten” the flow. This is illustrated witha velocity vector plot in figure 9. The accuracy with which the design pressuredrops were matched is shown in figure 11.

Heat addition to the air flow in the three coolers was accommodated in aconsistent fashion. Specifically, local heat addition sources were added to theenthalpy equation based on a local energy balance: qlocal = ˙ m localECPDTHX ,where DTHX ≡ Tair - Tcoolant , CP is the specific heat of the air and E is an un-known cell “efficiency”. For each exchanger we have available approximatevalues for qtotal and D THX, which when substituted into the equation aboveyield 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 de-sired DT was matched to requisite accuracy. The accuracy with which the de-sign temperature rises were matched is shown in figure 12.

The fan was also modeled using a bulk representation. Specifically, ap-proximate machine rotation rates and pressure rise across the fan were avail-able. Since the mass flow rate and flow path are fixed, a suitable body forcedistribution could be designed. A design code at Penn State Applied ResearchLab was utilized to generate tangential, axial and radial forces through the me-ridional plane of the fan. These were then distributed onto the NPHASE gridusing bilinear interpolation. Elements of this procedure are illustrated in figure10. The accuracy with which the design pressure rises were matched is shownin figure 11.



Soakback NPHASE Modeling

Soakback is the engine cool-down transient. This cool-down process has thepotential to be design limiting since at engine shut off, all engine and transmis-sion cooling flow ceases. The large thermal capacitance of the engine andtransmission can lead to local increases in radiative heating within the com-partment since convective cooling within the compartment will be signifi-cantly diminished. The cool-down transient has a time scale on the order of anhour, whereas at engine shut-down, convection cooling is lost (or with a smallauxiliary soakback fan operating, greatly reduced) in a matter of seconds. Ac-cordingly, this disparity in scales led to a soakback modeling approach involv-ing a steady state analysis with the small soakback fan flow to provide heattransfer coefficients for the transient radiation analysis. This approximationshould be valid, since heat transfer coefficients will be a weak function of com-ponent temperatures, and, as indicated above, little accuracy will be lost by ne-glecting the short duration convective heat transfer transient right at engineshut-down.

Engine shut-down is specified to occur while the engine is at TI. Accord-ingly, NPHASE was run with all boundary conditions as specified above forTI, with the exception of inlet mass flow rate, which was set to a small valuecorresponding to a notional soakback cooling fan. Also, no flow was admittedthrough the heat exchangers for SB operation, as described above.

OT and TI RADTHERM Modeling

The engine compartment contains a variety of materials including metals andalloys, insulation and lubrication oils. In general, the emmissivity, conductiv-ity, density and specific heat of these materials had to be obtained. Each of thesurfaces in the model had to be assigned to one of these materials and given anominal thickness for conduction heat balance purposes.

The heat transfer coefficients and film temperatures predicted by NPHASEwere input into RADTHERM as described in the “Software Tools” sectionabove. The ICEM coarsening process left several of the key engine compart-ment parts under-resolved (i.e., too few triangle elements), so ICEM’s refine-ment feature was then applied to hull, engine, airbox, and transmission com-ponents to recover requisite surface resolution. Figure 13 illustrates exampleNPHASE and RADTERM surface meshes for the airbox and generator com-ponents.

RADTHERM is brought up and the model imported through thePATRAN neutral file generated from the NPHASE solution file interpolatedonto the coarsened surface mesh (using the mapconvbp utility provided byThermoanalytics, Inc., vendors of RADTHERM). The RADTHERM displayappears as shown in figure 14. Each of the 91 parts was then assigned appro-priate material definitions and thickness as indicated above. All parts exceptthe engine components and exhaust duct are specified to employ the importedH and Tfilm values. Engine components are assigned fixed temperatures per

figure 5. The exhaust duct is assigned a fixed temperature of Tref for both OTand TI.

Appropriate values for the “back” (i.e., not-flow-facing) side of every partmust also be specified. For the 2-sided “sheet-metal” parts, H and Tfilm valuesare available from the NPHASE solution and are employed for each side. Forthe steady state OT and TI simulations, the engine and exhaust duct back sideboundary conditions are not important since the engine is almost a completelyclosed surface and the outer surface temperature is specified. Accordingly,there H is set = 0. The transmission back side is set to a constant temperatureof 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 andTfilm are specified as in the NPHASE analysis discussed above. The generatorbackside is specified as adiabatic. For all hull pieces, including the bulkheadand engine cover, H=0 is specified on the backside, which still allows radiativetransfer away from the engine compartment. All other pieces are set to adia-batic on the backside.

Emmissivities are set for the front and back faces of all parts. These are setbased on the material of the part, or, if the part is painted, an appropriate em-missivity corresponding to the paint is set. All components are painted exceptthe compartment cover, bulkhead, rear door, and the engine itself.

For OT and TI, RADTHERM was run as specified above for 400 itera-tions. This was sufficient to converge the runs to within 0.05oF. For compari-son, another pair of OT and TI runs was performed with all internal engineconvection cooling “turned-off”, that is, all heat transfer coefficients importedfrom NPHASE were overwritten as equaling zero. These “no-cooling-flow”runs were performed to provide upper bound conservative estimates on peakcomponent temperatures.

Soakback RADTHERM Modeling

As discussed above, the soakback runs were performed transiently using the Hand Tfilm field obtained from steady state SB NPHASE simulation. In order toaccommodate the critical thermal inertia physics associated with soakback, sev-eral changes were made to RADTHERM part specifications. These changesinvolved the engine, transmission, recuperator, and exhaust duct. Specifically,the transmission was redefined as a 3-layer part as illustrated in figure 15a. Thereasoning for this is as follows: For soakback, the transmission was originallymodeled as a large chunk of alloy, of nominal thickness to match the dryweight of the transmission and thereby mimic its thermal capacitance. Thisnon-conservative assumption allowed all incident radiative flux to be con-ducted away very efficiently into the transmission, thereby rapidly “smooth-ing” hotspots on the surface of the transmission since the thermal conductivityof the alloy is high. Moving to the 3-layer model is more physically realisticbecause the presence of the air and transmission oil layers (above and belowthe sump line, respectively, as illustrated in figure 15b), that exist in the real



configuration, greatly inhibits conduction normal to the transmission surface,while the thick third layer still accommodates the thermal capacitance of thesystem. The nominal core thickness of the transmission was determined usingan estimate for the volume of the transmission, the transmission’s known dryweight, and the density of the alloy.

A similar 3-layer model is employed for the engine, as illustrated in figure16. The nominal core thickness of the engine was determined using an esti-mate for the volume of the engine, the engine’s known dry weight, and thedensity of steel. The Tinit/Tref = 1.24 core initial temperature of the engine wasestimated based on engine output, fuel consumption and engine air flow rate.

The recuperator and exhaust duct were also treated as three-layer parts withouter surfaces of painted sheet metal, a layer of insulation and an inner layer ofsheet metal. A zero heat transfer coefficient was set on the inside of the recu-perator and exhaust duct for soakback.

Except as noted above for the engine, all parts were given an initial tem-perature distribution from the steady TI simulation.

Results

As already indicated, three sets of runs were made: OT, TI, and SB. In thissection we summarize the results obtained. First, details of the CFD simula-tions are presented, followed by the RADTHERM results.

OT and TI NPHASE Results

Figures 17 through 27 contain elements of the NPHASE simulations per-formed. In figure 17, selected streamlines emanating from the inlet are shownfor 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 directpath from the inlet to the heat exchanger pack, a good deal of the flow is seento divert to either side, and this gives rise to a fairly complex compartmentflow field. The streamlines that transit the compartment toward the heat ex-changer pack, come in close proximity to the exhaust duct and are therebyheated through convection from the exhaust duct. The T/Tref = 0.82 isosurfaceis seen to envelop the engine, heat exchanger pack, and exhaust duct in thisview, as expected. The strongly swirling exit flow induced by the fan is also ob-servable in this view (2 counter-rotating vortices).

Figures 18 shows two views of numerous streamlines seeded at the inletand/or the top of the heat exchanger for OT. A very complex three-dimensional flow field is seen to exist throughout the engine compartment.

Figures 19 and 20 show near engine views of the predicted velocity field forOT and TI. The velocity vectors shown are projected into the viewing planeand are shaded by magnitude of total velocity. It is observed that in the imme-diate 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).

Peak values for TI are somewhat smaller than OT, as expected. Figure 21shows two cross-sectional views of density contours, illustrating the weak butnon-negligible thermally induced perfect gas density variations within the en-gine compartment.

Soakback NPHASE Results

Elements of the NPHASE soakback solutions are provided in figures 22-27. Infigure 22, predicted contours of pressure are shown, for a compartment cross-section, illustrating the gravity head most responsible for driving the flow inthis case. Figure 23 illustrates the complex nature of the buoyancy driven flowwithin the engine compartment. There, numerous streamlines, shaded bytemperature, are plotted. In general, the entire compartment is subject to freeconvection flow. Temperature contours in the vicinity of the generator areplotted in figure 24. Clearly, free convection from the hot exhaust duct im-pacts the engine cover and generator. Figure 25 shows a similar plot for a slicetaken in a nominal engine-airbox plane. Here free convection heat transferfrom the engine burner region and recuperator are seen to clearly impact theairbox. Predicted velocity vectors in a cut plane through the transmission andexhaust duct are shown in figure 26. These vectors are resolved into the x-y cutplane 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 ofVinlet/Vref = 2.4, 1.52 and 0.034 for OT, TI and SB, indicating a weak but non-negligible flow. Free-convection induced velocity vectors in the y-z plane inthe vicinity of engine and transmission are shown in figure 27, illustrating sig-nificant 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 withthe NPHASE H and Tfilm and with no convection cooling for the engine com-partment cover, generator and airbox. These results, as well as those of otherengine components and for the TI operating condition illustrate that signifi-cant hot spots can occur opposite the engine and transmission surfaces. Alsofound is that convection cooling provided significant reduction in these hotspot peak temperatures as expected.

Soakback RADTHERM Results

As mentioned above, the soakback transient could be limiting for several criti-cal engine components due to the sudden loss of internal convective cooling tothe engine and transmission thermal masses. The desire to capture this was ac-



commodated in the RADTHERM modeling strategy outlined above.RADTHERM was run in transient mode for the SB case. After some experi-mentation, it was found that time steps of 30 seconds, and per-time-step con-vergence tolerances of 0.5oC yielded requisite numerical accuracy. The resultsin figures 30 through 36 summarize these runs.

Figure 30 illustrates the thermal history of the airbox. It is observed that thetemperature of the airbox hot-spot adjacent to the recuperator increases beforepeaking after t/tref @ 4, and then dropping off. The figure includes a time his-tory of the peak airbox temperature, including a comparison to anNPHASE+RADTHERM simulation with no soakback flow. It is seen thatthat the impact of soakback cooling flow is very small.

Figure 31 illustrates the thermal history of the bulkhead. The temperatureof 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 peaktemperature shown in the figure demonstrates that the impact of soakbackcooling flow is small, however, it is seen to actually increase the bulkhead peaktemperature somewhat. The physical explanation for this counter-intuitivefinding can be gleaned from figure 32. There, predicted total velocity vectorsare plotted in an x-y plane midway between the smallest recuperator-bulkheadgap. Clearly, the soakback flow has a significant effect on the local velocityfield there, with significant cross flow arising compared to the principally ver-tical buoyancy dominated field without soakback flow. Comparison with thelower part of the figure shows that with the soakback fan on, cross flow givesrise to higher local convection temperatures by virtue of the transport of hotterfluid into the hot-spot region. This in turn gives rise to higher film tempera-tures when interpolated to RADTHERM. Accordingly convectionflux, )TH(Tq filmwall -=¢¢ , 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 peakpredicted engine compartment cover temperature by a negligible amount.

The thermal history of the engine is illustrated in figure 34. There it is seenthat the engine cools quite rapidly from its initial condition (steady TI opera-tion). The figure illustrates that all components of the en-gine+recuperator+exhaust duct cool with the exception of those parts in im-mediate contact with the burner, which retain higher temperatures throughconduction from the burner region. The peak engine temperature decay israpid in the first t/tref @ 10 after engine shut down, but levels off to a muchslower decay after that. Simulations with and without soakback fan cooling arevirtually indistinguishable for the engine on this scale. For comparison, in fig-ure 34, the original engine model prediction (no-multilayer treatment – as-sumed pure metal) is included. It is observed that the improved physical mod-eling 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 hotspot on this part is seen to decay monotonically during soakback. This is dueto the comparatively low thermal mass of the exhaust duct which itself cools

quickly (refer figure 34). The simulations with and without soakback coolingare very similar.

Figure 36 shows views of the transmission over the t/tref = 30 after engineshut down. Clearly observable is the decay in the engine-burner-facing hotspot. Also observable is the relatively rapid cooling experienced in the sumpregion compared with the transmission housing above the sump line. This ofcourse is due to the modeling discussed above. The figure also includes thetime history of the decay of the two transmission hot spots. The temperaturedecays monotonically from engine shutdown for both hot spots, and simula-tions with and without soakback cooling are very similar.

Conclusions

The analyses carried out under the present effort involved the application of state-of-the-art grid generation, CFD and radiation heat transfer analysis software. Largehigh quality hybrid unstructured meshes were constructed using ICEM-CFD. TheNPHASE CFD code was employed to solve the flow and convective thermal trans-port. The commercial radiation analysis package, RADTHERM, was employed inassessing the radiation dominated peak temperatures in the system.

References

Kunz, R.F., Yu, W.S., Antal, S.P., Ettorre, S.M. “An Unstructured Two-FluidMethod 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 Next-Generation Computer Code for the Prediction of Multicomponent MultiphaseFlows,” presented at the International Meeting on Trends in Numerical and PhysicalModeling for Industrial Multiphase Flow, Cargese, France, September 27, 2000.

Yu, W.S., Kunz, R.F., Antal, S.P., Ettorre, S.M. “Unstructured Rotor StatorAnalysis of Axial Turbomachinery Using a Pressure-Based Method”, ASME Paper pre-sented at the International Mechanical Engineering Congress and Exposition, NewYork, 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.



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 con-figuration. b) View of 24 domain METIS partitioning for engine compartment model.

Fig. 4. Representative NPHASE convergence history for engine compartment analyses. Shownare pressure and velocity residuals, and global mass flow rate through compartment (outflow-inflow).

Fig. 5. Engine T/Tref map, specified in NPHASE for OT/TI operation.

Fig. 6. Illustration of transmission components and specialized conjugate heating boundarycondition employed in NPHASE.

pressure

velocity

mass flow



Fig. 7. Illustration of airbox and specialized conjugate heating boundary condition employed inNPHASE.

Fig. 8a. Illustration of three heat exchangers and design specifications for their respective pres-sure drop, Dp/Dpref, at OT & TI operation.

Fig. 8b. Illustration of three heat exchangers and design specifications for their respective tem-perature rise, DT/DTref, at OT & TI operation.

Fig. 9. Predicted velocity vectors just above and through the three heat exchangers, illustratingthe straightening due to resistance modeling employed.

Fig. 10. Elements of fan body-force modeling employed. a) View of fan vicinity, b) output ofthroughflow code.

Fig. 11. Errors in NPHASE vs. design pressure changes (DpNPHASE-DpDESIGN) across heat ex-changers and fan.



Fig. 12. Errors in NPHASE vs. design temperature changes (DT DT )- acrossheat exchangers.

NPHASE DESIGN

Fig. 13. NPHASE and RADTHERM surface meshes for generator and airbox.

Fig. 14. RADTHERM interface with view of engine compartment model (grey-scale).

Predicted Temperature Contours

Fig. 15a. 3-layer transmission part treatment for soakback RADTHERM analyses.

Fig. 15b. ICEM-CFD family split employed for the transmission to distinguish between regionswith sump oil and those without under soakback conditions.

Fig. 16. Sketch of 3-layer engine part treatment for soakback RADTHERM analyses.



Fig. 18 Selected streamlines from OT NPHASE simulation..

Fig. 17. Elements of OT NPHASE simulation. Streamlines shaded by temperature, and tem-perature (T/Tref = 0.82) iso-surface (lighter surface enshrouding engine region).

Fig. 19. Rear-view in-plane velocity vectors, shaded by velocity magnitude (scale in V/Vref), forOT (left) and TI (right) NPHASE simulations.

Fig. 20. Top-view in-plane velocity vectors, shaded by velocity magnitude (scale in V/Vref), forOT (left) and TI (right) NPHASE simulations.

Fig. 21. Two cross-sectional views of density contours (r/rambient), illustrating weak thermallyinduced perfect gas density variations within the engine compartment.

Fig. 22. NPHASE SB simulation. Pressure contours, illustrating weak gravitational head risewithin engine compartment.



Fig. 23. NPHASE SB simulation. Streamline field illustrating the complex thermally inducedfree-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.

Fig. 26. NPHASE SB simulation. Free-convection induced velocity vectors in x-y plane in vi-cinity of transmission, generator and exhaust duct.

Fig. 27. NPHASE SB simulation. Free-convection induced velocity vectors in y-z plane in vi-cinity of engine and transmission.

Fig. 28. RADTHERM OT temperature predictions on engine compartment cover. WithNPHASE H,Tfilm convection values (left) and no-convection (right).



Fig. 30. Sequence of predicted airbox surface temperature fields for SB case. Temperature con-tour range is from 120oF to 240oF. t/tref=0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peak pre-dicted airbox surface temperature.

Fig. 29. RADTHERM OT temperature predictions on generator and on airbox. WithNPHASE H,Tfilm convection values (left) and no-convection (right).

Fig. 31. Sequence of predicted bulkhead surface temperature fields for SB case. Temperaturecontour range is from 120oF to 225oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peakpredicted bulkhead surface temperature.

Fig. 32. NPHASE predicted total velocity vectors in an x-y plane midway between smallest re-cuperator-bulkhead gap for SB cases. NPHASE predicted temperature contours (T/Tref) onbulkhead for SB cases.



Fig. 33. Sequence of predicted engine compartment cover surface temperature fields for SB case.Temperature contour range is from 115o oF to 185 F. t/t =0, 2.5, 5, 10, 15, 20, 25, 30. Timehistory of peak predicted engine compartment cover surface temperature.

ref

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. Timehistory of peak predicted engine surface temperature.

Fig. 35. Sequence of predicted generator surface temperature fields for SB case. Temperaturecontour range is from 120oF to 290oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peakpredicted generator surface temperature.

Fig. 36. Sequence of predicted transmission surface temperature fields for SB case. Temperaturecontour range is from 120oF to 280oF. t/tref =0, 2.5, 5, 10, 15, 20, 25, 30. Time history of peakpredicted transmission surface temperature. Hot spots considered here are directly across fromthe engine burner (left) and directly attached to the engine (right).


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Flow Field and Thermal Management Analysis of an Armored ...extras.springer.com/2004/978-3-642-53586-4/385.pdf · based on nominal air-box cross-sectional area and given engine air