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Thermal Management of Airborne Early Warning


Thermal Management of Airborne Early Warning and Electronic Warfare Systems Using Foam Metal Fins

Thermal Management of Airborne Early Warning and Electronic Warfare Systems Using Foam Metal Fins
John F. Klein, Noe Arcas, George W. Gilchrist, William L. Shields, Jr., Richard Yurman, and James B. Whiteside
Northrop Grumman Integrated Systems
This program addresses the need for thermal management of increasingly powerful and densely packaged electronic devices. Open-celled foam metals promise much better heat transfer between the coolant and the solid, compared with advanced fin heat exchangers. Potential benefits include reductions in the heat-exchanger-core volume and weight, as well as electronic device junction temperature. Equally important is the possibility of direct-attach cooling of high-power electronics. We developed and demonstrated the application of foam metals as heat exchanger cores in a laboratory-scale test model. To do this, we measured the effective heat-transfer coefficient and airflow resistance of foam-metal fins, and applied the results to the design and sizing of a demonstration laboratory-scale test heat exchanger. We developed a theoretical model of foam metal fins and used it to scale foam metal fins for sizing of a naval aircraft heat exchanger. Thermal testing of a six-element subpanel agrees within 3.2% of expected performance. A foam metal core capable of removing 100 kW of heat, in real time, from an aircraft flying at Mach 0.28 at 15,000 ft would weigh just 22 lb. The counter-flow heat exchange rescaling is linear, so a similar foam metal core capable of removing 1 MW, in real time, would weigh just 220 lb. The procedures we developed to prepare the test items, conduct the tests, and the results database we created can be applied to other applications. We are currently evaluating foam-metal heat exchangers for cooling electronics racks on submarines and radar power-amplifier modules. Foam-metal heat exchangers could also be considered for real-time heat rejection from airborne laser applications.

Introduction
Military and civilian electronic, avionic and power applications need to package increasing power in less space, using less weight, and/or lowering electronic device operating temperatures. Furthermore, heat must be removed from the platforms and rejected to the surrounding environment. Conventional cooling systems for our naval aircraft applications are becoming larger and heavier than desired. In support of Northrop Grumman Integrated Systems’ major System Development and Demonstration program for a naval aircraft, we successfully evaluated and developed a lightweight foam metal heat exchanger as a candidate for significant weight savings. Benefits of Foam Metal in Advanced Aircraft Heat Exchangers. A liquid-to-air, foam metal heat exchanger is considered here which was found to provide significant weight savings.
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Heat is rejected from liquid within the aircraft to air outside the aircraft. Air flows through a pod located outside the fuselage. To meet that requirement, heat exchangers using advanced conventional fin technology would have to be larger and heavier than desired. Foam metals have the potential to provide significantly better extended surface heattransfer for use in light, compact heat exchangers. Significant heat exchanger core size and weight savings have been projected by applying foam metal technology. The naval aircraft Thermal Design Group motivated consideration of the liquid cooling system heat exchanger for application of foam metal technology. Initial projections of foam metal benefits for the liquid cooling system (LCS) heat exchanger were based on our calculations, extrapolation of performance data for foam metals, and information from foam metal vendors. Our efforts focus on reticulated foam metals. Reticulation is a foam manufacturing process which produces foams well suited to our requirements as discussed later. Foam-metal heat transfer data available in the literature are for conventional off-the-shelf reticulated foam metals (7% to 11% dense, illustrated in Figure 1). These data are also in flow regimes that do not correspond to candidate application flow conditions. Foam metals with the desired heat transfer performance are compressed foams (15% to 30% dense), which are required for small, lightweight heat exchangers at high altitude airflow conditions. Our approach was to measure the air-side heat transfer and pressure drop characteristics of compressed foam metal under conditions that could be of interest for a variety of applications. We applied the results to the sizing and design of a laboratory-scale heatexchanger thermal-performance demonstration unit. Foam-metal heat exchangers are used to best advantage where space and weight are at a premium. However, they tend to be somewhat more expensive than conventional and advanced conventional fins. For new applications, a short fabrication development phase may be needed to produce new configurations. Our testing and analysis produced significant accomplishments, detailed in the sidebar on the next page. Most notably, we achieved significant weight savings for a naval aircraft heat exchanger. We also established a database for air-cooled foam-metal extended surfaces and used the data to evaluate alternative foam-metal heat exchanger designs for a naval aircraft. We designed, fabricated, and verified three test rigs: ? Kays & London airflow heat-transfer apparatus ? Foam-metal heat exchanger panel demonstration apparatus ? Kays & London liquid-flow heat-transfer apparatus (design and fabricate only)

Figure 1. Reticulated foam-metal extended surface 104 Technology Review Journal ? Spring/Summer 2005

Thermal Management of Airborne Early Warning and Electronic Warfare Systems Using Foam Metal Fins

Summary of Significant Accomplishments
We achieved the following significant accomplishments in our evaluation of a foam-metal heat exchanger technology: ? Demonstrated the thermal performance of a foam-metal heat exchanger core with significant potential weight savings for the radar liquid cooling system (LCS) for a U.S. naval aircraft ? Demonstrated the viability and advantages of foam-metal heat exchangers to the naval aircraft Thermal Design Group engineers and enlisted their support ? Established the foam-metal heat exchanger as a top naval aircraft weight-saving candidate ? Developed packaging concepts for the naval aircraft LCS foam-metal heat exchanger ? Developed naval aircraft LCS foam-metal heat exchanger assembly concepts to fit the LCS side pod ? Developed a “rules and tools” design methodology for foam-metal heat exchangers ? Demonstrated the capability of foam-metal liquid cooling to reduce junction temperature in naval aircraft radar power-amplifier-module (PAM) transistors by about 10°C ? Developed electronic-rack and cold-plate concepts ? Identified the potential of foam-metal liquid-cooled heat exchangers as part of a thermal electric system for cooling submarine electronic racks ? Established a Kays & London–type database for air-cooled foam-metal extended surfaces ? Designed, fabricated, and verified three test rigs: —Kays & London airflow heat-transfer apparatus —Foam-metal heat exchanger panel demonstration apparatus —Kays & London liquid-flow heat-transfer apparatus (design and fabricate only) ? Established a go-forward plan for naval aircraft implementation of the foam-metal heat exchanger as a naval aircraft weight saver

Applying the air-cooled foam metal database, we designed a counter-flow, foam-metal heat exchanger core for a naval aircraft liquid cooling system (LCS) heat exchanger. This core uses foam metals as the extended surface on both air and liquid sides of the heat exchanger. The resulting core is significantly smaller and lighter than both conventional and advanced conventional fin heat exchangers. Foam-metal core volume and weight can be as little as half that of advanced conventional fins for the same overall performance parameters. Thermal testing of both single- and six-element cores verified our design/ sizing expectations. Naval aircraft thermal engineers are encouraged by the foam-metal heat-exchanger-core heat transfer demonstration. As a result, we have begun to develop a heat-exchanger assembly packaging concept. The foam-metal heat exchanger is being considered as a top candidate in a Northrop Grumman naval aircraft weight savings program. Public data were also lacking on liquid-cooled foam-metal heat transfer. Therefore, we designed and fabricated a test apparatus to establish a Kays & London–type database for compressed foam metals. Test setup, installation, and data acquisition system modification were supported in parallel by additional Northrop Grumman capital investment. Actual testing will be started in 2005.
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Heat transfer and pressure-drop test results for compressed aluminum foam metals are reported here. Results were positive. One foam demonstrated airside performance similar to that needed for application to the naval aircraft LCS. The results were used to size candidate foam-metal heat exchangers. Test results for single-element and six-element subpanel air-to-liquid foam-metal heat exchangers verified our thermal performance expectations. Other Applications of Foam Metal Technology. Other potential applications of foam-metal heat exchanger technology in Northrop Grumman have been identified. Its use in cooling electronic racks on submarines is being evaluated with our Newport News submarine division, under a program sponsored jointly by Northrop Grumman and the Office of Naval Research (ONR). Liquid-to-liquid foam-metal heat exchangers will be demonstrated as part of that program. Liquid-cooled foam metal fins are being evaluated for use by the naval aircraft to cool high-power radar module transistors. Initial analysis and tests have demonstrated that at least a 10°C reduction in power-transistor-junction temperature is feasible. Another airborne application under consideration is real-time waste-heat removal from airborne laser (ABL) directed-energy weapon (DEW) aircraft platforms. Although foam metal fins are being evaluated as a replacement for conventional fins, foam metal cannot directly replace conventional fin materials because of the higher pressure drop per unit length associated with the higher heat transfer of foam metals. A different heat exchanger design is required. Foam-metal heat exchangers tend to be thin, flat panels. Counter-flow heat exchanger configurations have the advantage of simple linear scaling of total heat exchanged with panel face area. For example, a foam metal core weighing 22 lb, with a face area of about 5 ft2, and removing 100 kW of heat can be linearly scaled to a foam metal core weighing just 220 lb, with a face area of about 50 ft2, and removing 1000 kW of heat. Such a 220-lb core could be of interest to some ABL DEW platforms flight conditions similar to what we analyzed. Other potential applications for foam-metal heat transfer technology include cold plates for electronics, heat sinks at the device and junction levels, power electronics, avionics pods, avionics cooling racks, and dual thermal/structural applications. A heat exchanger has been designed and sized for a naval aircraft based on the foam-metal performance characteristics we have measured. Heat exchanger thermal performance has been verified by laboratory testing of both single heat exchange elements and a sixelement panel. Conceptual designs based on these foam-metal cores offer the potential to save significant weight. Application of foam-metal heat exchangers on a naval aircraft is being evaluated.

Test Apparatus
Overview of Test Apparatus
There was no heat transfer data available in the literature for compressed aluminum foammetals of the type we were considering using at the start of this program. Therefore, we built our own heat transfer test apparatus, which we have used to measure the heat transfer characteristics of various foam-metal fins. Figure 2 presents an exploded view of the heat transfer test apparatus. Filtered air flows through a settling chamber, into the test section, and exhausts into the room. The aluminum foam-metal test sample is brazed to two aluminum attachment plates, as Figure 3

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Insulation Blocks

Insulation Blocks Air Out

Plexiglas Downstream Airflow Duct Insulation Housing Air In Adjustable Airflow/Duct Height Foam Metal Sample

Airflow Transition Section from Setting Chamber

Figure 2. Exploded view of heat transfer test apparatus

Heat Addition

Airflow

Heat Addition

Figure 3. Heat transfer test sample

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shows. Air flows through the foam-metal test sample (Figure 3). Heat is applied to the test sample using electric heaters mounted on its upper and lower surfaces. The foam-metal test sample and its heaters are mounted into a thermally insulated test section (Figure 2). The thermal insulation ensures that the heater power goes into the test sample. The temperature of the sample attachment plate is controlled to 65.5°C by a proportional integral differential (PID) controller. Air temperature is measured upstream and downstream of the foam sample. The effective heat-transfer coefficient Heff is determined based on the heater power, foam mounting-surface temperature, air temperatures, and sample geometry. All test data were recorded on a digital computer.

Detailed Description of Test Apparatus
Airflow. Airflow is supplied by a compressor system that cools and dehumidifies the air after it leaves the compressor. Air flows through the foam-metal test sample in the oneinch direction (Figure 3). Each test sample is sealed in the airflow duct to ensure that the air does not bypass it. Sample thickness transverse to the airflow direction can be varied from 0.95 to 2.54 cm. The height of the airflow passage is adjusted to match the sample thickness. One wall of the airflow duct is fixed in space; the other wall is adjustable (Figure 2). The airflow duct and sample width are constant (15.24 cm) for all tests. Air temperatures were measured upstream and downstream of the foam sample. The airflow rate was measured using a digital flow meter connected to the data acquisition system. Airflow pressure drop across the foam is measured using pressure taps, located in the duct wall upstream and downstream of the sample, and a differential pressure transducer. Foam-Metal Test Sample Assembly. The foam-metal test sample is brazed to two aluminum mounting plates on the 2.54 × 15.24 cm attachment surface (Figure 3). Five thermocouples are drilled into each mounting plate at the interface between the foam and the foam attachment surface. These five thermocouples are used to determine the foam-sample/ mounting-plate interface temperature. Four thermocouples are located in each corner, and a fifth thermocouple is located near the middle of the 2.54 × 15.24 cm attachment surface. A sixth thermocouple, for PID control, is located symmetrically with the fifth thermocouple. Heater assemblies are attached to the surfaces indicated as Heat Addition in Figure 3. Thermal conduction losses from the sample were minimized by providing thermal isolation from the sample mounting-attachment surfaces and surrounding the sample with thermal insulation. Electric Heaters/Determination of Sample Power. Two independent electric heater assemblies apply heat to opposite sides of the foam metal sample. Each heater assembly consists of a mica heating element attached to an aluminum base plate by squeezing the heater between a compliant thermal insulating pad and a 0.635-cm-thick stainless steel plate. Each heater assembly is capable of delivering about 26.3 W/cm2 to the sample (i.e., 850 W per heater.) Indium foil was used as a thermal interface between the heater assembly and the sample mounting-plate surfaces. Each foam-sample/mounting-plate interface temperature is controlled independently using a two-zone PID controller. The PID controller is tuned and programmed to maintain the mounting-plate interface temperature at about 65.5°C. This temperature was stable to ±0.08°C during the time that a data set was taken. Heater power is determined by measuring the heater current, voltage, and duty cycle. The heater power measurements were confirmed two different ways. First, data were recorded offline during the data collection period and analyzed using a frequency analyzer. Second,
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a heat balance was performed, using air mass flow rate together with inlet and outlet air thermocouple temperature. The three power measurement methods agreed within ±3%. The test sample is mounted in the test section so that it is enclosed in thermal insulation. The conduction heat loss from the sample was determined by turning off the airflow at the end of the day, sealing the outlet duct with insulation, and recording the cool-down rate of the sample assembly. The conduction-heat-loss rate was determined based on the cooldown rate and the weight of the sample assembly, including heater assembly. Based on these test results, the conduction heat loss rate was less than 3 W. Total heater power for the samples during these tests ranged from about 150 to 850 W.

Test Procedure
Test Sequence
Airflow was started at a low flow rate and increased in steps to the highest flow rate for each test series. Airflow was set to the desired level for the first test point. Heater power was turned on after the airflow had stabilized. Both sample mounting-plate interface temperatures were slowly ramped up to 65.5°C using the PID controller. Airflow and heater power were allowed to stabilize (typically in 1 to 2 hr). After airflow and heater power were stable, a data set was taken over about 10 min. Generally, the data rate was 10 readings per minute. All readings—such as thermocouple temperature, airflow rate, pressure drop, and heater power—were recorded by a digital computer. The data acquisition system recorded the readings electronically in a Microsoft? Excel spreadsheet format. Values for each 10-min data set were averaged for the duration of the recording interval. Airflow was then adjusted to the next flow rate, and the system was allowed to restabilize. The data acquisition system calculated a running value of parameters that helped determine when airflow and thermal steady states were reached. Data were taken 20 to 30 min after steady state was reached. For reference, the definitions of all terms in the equations below are listed in the “Nomenclature” sidebar, page 110.

Determination of Effective Heat-Transfer Coefficient
The effective heat-transfer coefficient Heff was determined using the power for each heater, the heater/foam-metal interface area, and the temperature difference between the air and the mounting-surface/foam-metal interface. The interface temperature was uniform within ±2% for most data points. A log mean temperature difference (LMTD) was used [1]. Average air inlet Tair,in, exit Tair,out, and plate temperature Tp were used to form the differences in the LMTD, as follows: ?Tin = Tp – Tair,in , ?Tout = Tp – Tair,out , LMTD = (?Tout – ?Tin)/ln(?Tout/?Tin) , Heff = Qheater/(LMTD(interface area) , where Tp = foam mounting surface temperature, °F Tair,in = air inlet temperature, °F Tair,out = air outlet temperature, °F Heff = effective heat-transfer coefficient at the foam-metal/aluminum interface, W/in2-°C
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Nomenclature
Af – airflow blockage factor As – foam-metal surface area per unit volume b – half-separation between liquid passages cp – specific heat Dh – hydraulic diameter Dlig – average diameter of the foam ligament dtin – temperature of liquid in minus Tair, in f – friction factor, ?p/[4(L/Dh)(ρVint2)] G – mass velocity, ρVint Heff – effective heat-transfer coefficient at the foam/aluminum interface htrue – local heat-transfer coefficient between the foam ligament surface and the fluid keff – foam effective thermal conductivity j – Colburn j factor, (Heff/cpG)Pr2/3 L – airflow length of foam sample PID – proportional integral differential controller ppi – starting foam nominal pores per inch Pr – Prandtle number Qheater – heater power (W) ρair – air density ρfoam – foam metal density, % solid Re – Reynolds number, GDh/? scfm – standard cubic feet per minute Tair, in – air inlet temperature Tair, out– air outlet temperature Tp– foam mounting-surface temperature T60/40 – temperature of 60/40 mix of eythlene glycol/water Vinf – approach velocity to the foam Vint – internal velocity within the foam, Vinf/Af ?Tin = Tp – Tair, inF ?Tout = Tp – Tair, outF ?p – air-side pressure drop across foam sample length L

Qheater = heater power, W The small conduction loss from the samples has been subtracted from Qheater for these calculations. The effective heat transfer coefficient Heff is measured in W/in2 -°C. The conversion to English units is 273.3 Btu/hr ft2 -°F, and to metric units is 1551.9 W/m2 -°C.

Discussion of Results
Compressed-Foam-Metal Heat Transfer and Pressure Drop
Test Samples and Test Matrix. Table 1 lists all samples tested, arranged by sample density. Sample characteristics included are nominal pores per inch (ppi), foam thickness (actual and nominal), density in terms of percentage solid aluminum, compression ratios along all three axes, surface area per unit volume, hydraulic diameter Dh, ligament diameter
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Table 1. Foam Metal Sample Characteristics
Compression Ratioa 1
Sample Number

Calculated Propertiesb 5
Specific Surface Area, As (ft2/ft3)

2

3

4

6
Airflow Passage Hydraulic Diameter, Dh (in.)

7
Ligament Diameter, Dlig (in.)

8
Airflow Blockage Factor, A f

Compression Direction

1-in. Airflow Direction, DL/L (Y)

6-in. Sample Direction, DL/L (X)

Release Density (%)

40 ppi c KL-16 Y KL-15 Y KL-2A Y Y KL-6 KL-12 Y, X 17.5 22.9 23.7 18.5 27.4 2.160 2.840 2.932 1.860 2.025 0 0 0 0 2.000 1.072 1.061 1.010 1.081 1.041 20 ppi c KL-11 Y KL-14 Y, X KL-13 X 32.1 32.3 33.0 4.097 2.055 0 0 2.000 4.171 1.031 1.034 1.041 2119 2132 2178 0.023 0.026 0.015 0.012 0.012 0.012 0.347 0.470 0.230 1584 2061 2133 1665 2446 0.018 0.016 0.016 0.018 0.016 0.009 0.009 0.009 0.009 0.009 0.430 0.374 0.366 0.432 0.417

a For compression directions, see Figure 4. b B. Ozmat, Ozer Engineering/ERG, Inc., private communications. c Vendor data

Dlig, and airflow blockage factor Af. Foam starting ppi, dimensions, density, compression ratios, and ligament diameter are physical properties determined and supplied by the foam vendor. Figure 4 defines the compression direction axes. The ability to compress foams fabricated by direct-foaming reticulation is a major advantage for heat transfer optimization. Foam Metal Extended Surface Characteristics. Reticulated foam ligament diameter is not uniform. Rather, it is a representative average value, determined by inspection under a microscope. It is smallest near the center of a ligament and largest near the ligament intersection nodes. The cross-section of each ligament is triangular, with the sides of the triangle slightly concave toward the centerline. Ligament cross-section shape is thought to increase local fluid turbulence and increase heat transfer with fluids flowing through the foam. The foams used in our study are fabricated by a direct-foaming/reticulation process. The cross-sectional area of ligaments for those foams is solid—an advantage for heat conduction. In foams fabricated by other methods such as precursor sintering processes, the ligament cross section is hollow in the center, which limits heat transfer along the ligaments and is thus a disadvantage for heat transfer. Such foams are manufactured by coating precursor polymer foam ligaments with a thin-layer slurry of the target foam material. Next, the coated precursor polymer foam is fired for sintering. The precursor burns off and a foam with somewhat porous ligaments remains. Ligament porosity is caused by polymer infiltration of the coating material during burn-off. Such foams tend to

Thickness Direction, DL/L (Z)

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Thickness Direction Z 1-in. Airflow Direction Y X 6-in. Sample Direction
Figure 4. Compression direction axes

be brittle. It is difficult if not virtually impossible to optimize the structure of such foams, for use as an extended heat-transfer surface by methods such as compression. The foam specific surface area, hydraulic diameter, and airflow blockage factor (Table 1) were calculated values supplied by the foam vendor. High values of specific surface area are advantageous for heat transfer. Compressing the foam and increasing the number of ligaments per unit volume results in high specific surface area. Compression also increases flow resistance. Trade-off between heat transfer and pressure drop due to compression are discussed below. Hydraulic diameter is determined based on the compressed-foam cell size, ligament diameter, and compressed-foam cell density, ppi. The airflow blockage factor Af determines the internal fluid velocity Vint within the foam. The internal velocity in the foam is calculated using the airflow blockage factor and the approach velocity Vinf . That is, Vint = Vinf/Af , where V inf = approach velocity to the foam, V int = internal velocity in the foam, Af = airflow blockage factor. Approach to Foam Metal Heat Exchanger Sizing. Velocity in the foam is higher than approach velocity. Internal velocity is used in physical scaling parameters such as the Reynolds and Stanton numbers. Approach velocity is useful for determining heat exchanger face area required for packaging design of a given application. Face area can be determined by working backward from the pressure drop requirement. Pressure drop requirement determines the maximum internal velocity. Using the area ratio, the approach velocity can be calculated. Face area is determined using mass flow rate, density, and approach velocity. That is, Aface = mdot/(air density)(Vinf) where Aface = heat exchanger airflow face area, mdot = total air mass flow. Heat Transfer and Pressure Drop Test Results. Most foam samples were 40 ppi, compressed in the airflow direction (Y on Figure 4). Some samples had biaxial compression in the X and Y directions. There is no intentional compression in the Z direction (between the sample mounting faces). The small values listed in Table 1 are due to slight changes in dimension during fabrication. Compression in the Z direction would increase pressure
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(4)

(5)

Thermal Management of Airborne Early Warning and Electronic Warfare Systems Using Foam Metal Fins

drop with little benefit in heat conduction to the mounting faces. Figures 5 and 6 present measured results of the relationship between airflow velocities in compressed 40 ppi foam for, respectively, effective heat-transfer coefficient and pressure drop. Heat exchanger core approach face velocity Vinf as a design parameter is useful in determining panel size and flow duct size required to supply and remove fluid. Therefore, approach velocity perspective is useful for heat exchanger system packaging considerations. For the purpose of understanding the heat transfer physics, the correct velocity to use is the internal velocity. Internal velocity, Vint, is determined from approach velocity and area factor (see Equation 4). Figure 5 shows heat transfer as a function of internal velocity. For the purpose of sizing the heat exchanger assembly approach velocity is the best parameter. Figure 6 shows pressure drop as a function of approach velocity. Below about 0.3 psi/in., heat transfer increases rapidly with greater pressure drop. Between about 0.3 and 0.6 psi/in., the rate of heat transfer increase declines. The curve bends over. Heat transfer tends to saturate as internal velocity increases. For our applications, the allowable pressure drop is in the 0.1 to 0.4 psi/in. range. We did not test pressure drops sufficiently high to determine whether heat transfer as a function of pressure drop continues upward at a moderate slope or reaches an asymptotic value. These trends could lead to a maximum in the relationship between heat transfer and pressure drop. Foam Metal Heat Exchanger Design Considerations. The test data shows that foam metals can provide high extended-surface heat transfer. However, pressure drop per unit length is relatively high. To manage the pressure drop per unit length, there is a prime trade-off for foam-metal heat exchangers between approach-face-area velocity and flow length though the foam. For a given coolant mass flow, the approach face area determines the coolant normal face velocity. Coolant normal face velocity determines the coolant flow resistance per unit flow path length in the foam metal. The total flow length determines the total pressure drop. Because of these considerations, foam-metal heat exchanger cores tend to be high frontal area and short flow length. These are the main reasons that a naval aircraft foam-metal core is a 1-inch-thick panel. For various extended surfaces, the trade-off considerations are quite different. Therefore, the optimal design for one type of extended surface will be very different from the optimal design for another extended surface. The best way to compare extended-surface technologies is to optimize, to the same performance requirements, the heat exchanger based on each extended surface. Heat exchanger technologies can be compared on an equal basis by using top-level performance parameters, such as total heat exchanged; inlet and outlet temperatures, and total pressure drop on each side. On that basis, foam-metal heat exchanger cores for our application have about half the volume and half the weight of optimized heat-exchanger advanced fins, such as lanced offset fins. Other practical consideration include the size of the passages relative to clogging by debris. Clogging by debris is an important consideration for microchannel heat exchangers. Colburn and Friction Factor Test Results. Results are presented in Figures 7 and 8 for the Colburn j factor and flow resistance f as a function of the internal Reynolds number, Re. Figures 7 and 8 present traditional nondimensional relationships for heat transfer data. The Colburn j factor (Figure 7), j = (Heff Pr2/3)/[(ρairVcp)air], is a comparison between surface heat removal Heff and the convective heat content of the fluid. At low Reynolds numbers, some of the data become independent of the Reynolds number. The Reynolds number is formed by the fluid properties, internal velocity, and foam hydraulic diameter (Table 1).
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5.0
Effective Heat-Transfer Coefficient, Heff (W/in.2 -°C)

4.5 4.0 3.5 3.0 2.5 2.0 1.5 1.0 0.5 0.0 0

KL-16 KL-6 KL-15 KL-2A KL-12

40 ppi foam

20

40

60

80

100

Internal Velocity, Vint (ft/s)

Figure 5. Effective heat-transfer coefficient as function of internal airflow velocity

0.5
KL-16 KL-6

0.4
Pressure Drop, Dp (psi)

KL-15 KL-2A KL-12

0.3

40 ppi foam

0.2

0.1

0.0 0 5 10 15 20 25 30
Approach Velocity, Vinf (ft/s)

Figure 6. Pressure drop per unit length as a function of approach airflow

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1
KL-16 KL-6 KL-15 KL-2A KL-12

Colburn j Factor

40 ppi foam

0.1 10

100
Reynolds Number, Re

1000

Figure 7. Colburn j factor as a function of Reynolds number

1
KL-16 KL-6 KL-15 KL-2A

Flow Resistance, f

KL-12

40 ppi foam 0.1

0.01 10

100
Reynolds Number, Re

1000

Figure 8. Friction factor as a function of Reynolds number

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The friction factor (Figure 8), f = ?p/[4(ρair V2/2)air (L/?h)], compares flow resistance per unit length scale with flow dynamic pressure. Data presented by Figure 8 show the foammetal friction factor to be relatively insensitive to the Reynolds number. At low Reynolds numbers, some data show increasing sensitivity to the Reynolds number. Heat transfer and flow resistance increase with foam density and internal airflow velocity for both dimensional and nondimensional relationships. The friction factor and Colburn j factor were used to scale between sea-level test data and high-altitude aircraft operating-point conditions for air: heat transfer coefficient, pressure drop, velocity, and density. Sea-level heat-transfer test data were scaled to high-altitude conditions at the same mass flow per unit heat-exchanger panel face area. The friction factor and Colburn j factor also can be used to scale in-service heat exchanger performance for operation at various altitudes. Expectations for heat exchanger performance must be adjusted based on aircraft flight conditions at different altitudes. Heat Transfer Pressure Efficiency. A measure of pressure efficiency can be formed by the ratio of heat transfer to pressure drop, expressed as a ratio of Colburn j factor to flow resistance f/2. Figure 9 presents results for those data. All but one of the samples tested have a maximum value of that ratio. The sample without a maximum may not have been tested at a Reynolds number sufficiently low to exhibit a maximum. Data in Figures 5 through 8 generally show higher heat transfer and pressure drop with increasing foam density. Data in Figure 9 show higher pressure efficiency with lower density material. Although the lower densities appear more pressure-efficient for heat transfer, the absolute value of the heat transfer is lower for the lower density samples. Foam-metal heat-exchanger core designs can be adjusted to achieve high heat transfer and avoid high-pressure drop. That approach was used to size a naval aircraft heat exchanger core discussed below. Comparison of Foam Metal Fin Compression Directions. Figure 10 presents the effect of foam-compression direction for 20-ppi foam samples. Compression-direction notations are indicated at the top of Figure 10. All three samples were compressed to about the same final density, 32% (Table 1). Compression in the y direction is parallel to the airflow. Compression in the x direction is transverse to the airflow. The highest heat transfer Heff was obtained with compression in the x or y directions. However, the highest pressure drop was obtained with compression in the x direction. High-pressure efficiency was obtained with both y, and equal x and y, compression. Compression in the y direction yields high heat transfer and moderate pressure drop. Equal x and y compression results in lower heat transfer and lower pressure drop, compared with y compression. Clearly, comparing such ratios can be misleading. One must also understand the application absolute drivers. The data suggest that a potentially rich design space exists for compressed-foam-metal heat exchangers, in terms of the relationship between pressure drop and heat transfer. To identify attractive combinations of heat removal, size, weight, and pressure drop, it was necessary to consider foam thicknesses other than those tested. A relationship for foamthickness scaling was derived by generalizing the method presented by Kern and Kraus [2]. The foam effective thermal conductivity was obtained from Ozmat [3]. Equations (6) and (7) present the resulting relationship for the effective heat-transfer coefficient:

Heff = keff m tanh (mb),
where m is defined as

(6)

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100
KL-16

Ratio Heat Transfer to Flow Resistance, j/(f/2)

KL-6 KL-15 KL-2A KL-12

40 ppi foam 10

1 10

100
Reynolds Number, Re

1000

Figure 9. Ratio of heat transfer to pressure drop

Surface with Heat Addition

Foam Compression Reference Directions Z Y Air, Vinf Surface with Heat Addition X

Foam Compression Direction Y, parallel to airflow Equal Y and X X

dPL 0.364 0.23 0.581

Heff 3.1 1.9 3.2

Heff/dPL 8.52 8.26 5.51

Figure 10. Effect of compression direction

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m=
where keff = foam effective thermal conductivity,

htrue As , keff

(7)

b = half-separation between liquid passages, htrue = local heat-transfer coefficient between the foam ligament surface and the fluid, As = foam-metal surface area per unit volume. The heat transfer coefficient between the airflow and ligament, htrue, was calculated based on Equation (3) and values of Heff, keff, As, and b. For comparison, the heat transfer coefficient between a cylinder of average ligament diameter Dlig and air velocity Vint was also calculated. The two values were within a factor of 2 of each other. Increasing cylinder diameter brings the two values for htrue closer together. Alternatively, a cylinder may not be the best representation of the ligaments inside the foam.

Foam Metal Heat Exchanger Thermal Performance Demonstration
Cross-flow and Counter-flow Foam Metal Heat Exchanger Core Configurations. Both crossflow (Figure 11a) and counter-flow (Figure 11b) were considered for foam-metal heat exchanger core configurations. Alternating sections of airflow foam-metal and liquid flow were used to make up the heat exchanger core. For a naval aircraft application, both crossflow and counter-flow heat exchanger cores have air flowing through the air-side foam metal in the 1-in. direction. Cross-flow liquid flow passages consist of extruded aluminum passages with integral fins. Liquid flows through the length of extrusion, in a direction across the air flow (Figure 11a). In contrast, in a counter-flow core, liquid flows through the liquid passage in the 1-in. direction, counter to air flow (Figure 11b). The counter-flow core we considered for a naval aircraft has compressed foam metal as its extended surface on the liquid side also. For application to a naval aircraft LCS, a counter-flow heat exchanger core resulted in a few pounds savings in core weight. Another advantage of a counter-flow core is that it allows more packaging flexibility than a cross-flow core. Changing the cross-flow liquid side length affects the total heat exchanged. Therefore, changing the shape for the crossflow core panel frontal area changes the total heat exchanged. In contrast, as long as the total air mass flow, approach velocity, and total frontal area are maintained for a foam-metal counter-flow core, the panel frontal area shape can adjusted to meet specific packaging requirements. Airflow design must be carefully done to ensure uniform airflow distribution across the entire heat exchanger face. Uniform airflow distribution for the naval aircraft application has been designed and verified by both CFD analysis and testing. Manufacturing Demonstration Panel. Figure 12 is a photograph of a 20-liquid-leg crossflow manufacturing demonstration core panel. The liquid legs are extruded fin stock channels. This demonstration panel is representitive of cross-flow type core illustrated by Figure 11a. A manufacturing demonstration panel representative of counter-flow core is in the process of being fabricated. Single-Element Core. A cross section of a counter-flow core element is shown in Figure 13. Liquid plenums and flow, as well as airflow, paths are illustrated. A counter-flow heat exchanger core consists of a stack of single elements and appears similar to the manufacturing demonstration panel shown in Figure 12.

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a. Cross-flow heat exchanger Air Out

b. Counter-flow heat exchanger

Liquid Plenums Air Out Liquid In

Air In

Air and Liquid Flow across Each Other Air In

Liquid Out

Air Flow Path Liquid Flow Path Air and Liquid Flow Counter to Each Other

Figure 11. Alternative approaches: Cross-flow and counter-flow cores

1 in. Liquid Flow Air Flow

11.8 in.

12 in.

Figure 12. Manufacturing demonstration panel

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Liquid Passage End Cap Liquid Passage Inlet Plenum Air/Liquid Separator Plate

Liquid Foam

Liquid Flow

Air Flow

Air Flow

Air Foam

Liquid Passage Outlet Plenum
Figure 13. Cross section of a single-element counter-flow heat exchanger core

Six-Element Subpanel. A subscale heat exchanger was designed and sized for a heat transfer operating point representative of the naval aircraft. The subscale heat exchanger consisted of six air-liquid counter-flow legs. A full-scale panel would have about 125 similar legs. For ease of fabrication, each leg was fabricated individually. Individual legs were mechanically stacked up to form a six-element heat exchanger subpanel (Figure 14). Figure 14a illustrates six elements being stacked up. Figure 14b illustrates the mechanical assembly. The mechanical assembly was used for thermal performance testing. Figure 15 is a photograph of our six-element performance demonstration test article, head on, as the air would approach the test article in the duct. Air foam passages and liquid plenum edges are shown in this photograph. Red silicon rubber gaskets seal the sixelement assembly to the Plexiglas? side mounting flanges. Cork gaskets, which seal the interface between individual elements, can be seen between adjacent half-height air foam passages. The cork gaskets are on a plane of heat transfer symmetry. The heat exchange section is about 11 in. high by 3.0 in. wide. The Plexiglas side frames and aluminum liquid headers at top and bottom would be replaced by a more compact design for an actual heat exchanger.

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Liquid In Air Flow

Single Element Air Flow Liquid Plenums

Capped Off

Liquid In

Liquid Flow

Liquid Out a. Assembly of elements

Liquid Out

Capped Off b. Mechanical assembly

Figure 14. Six-element heat exchanger subpanel

Liquid Pressure Tap Holes

Liquid Plenums

Air Foam Passages

Silicon Rubber Gaskets (red)

Cork Gaskets (tan)

Mounting Flanges

Figure 15. Six-element performance demonstration test article: Head-on view

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Heat Transfer Verification Test Results. Initially, we tested a single element. Singleelement test results verified our sea-level performance expectations, based on our test data. The next step for our foam-metal heat exchanger was to complete fabrication, assembly, and testing of the six-element performance demonstration subpanel. First, we verified the overall pressure drop of the six-element assembly. Air-pressure drop for the six-element subpanel, as installed in the test apparatus, was measured to be 0.12 psi. Cold-flow pressure drop for the six-element thermal test unit was compared with that of the K&L foam-metal test data for 15.7 % dense foam. The two sets of data agreed very well at the same airflow. This agreement provided initial conformation that the airside foam density reported by our vendor was what we expected. Figure 16 presents heat transfer test results for the six-element subpanel. Calculated performance is presented for a heat exchanger based on the thermal characterization database for KL-8, a heat exchanger with 15.7% dense foam. At the baseline airflow, 159 standard cubic feet per minute (scfm), calculated and measured heat exchange values agree within 3.2%. Air and liquid side heat balances agree within 2.7%. That is excellent agreement for laboratory heat transfer testing. Comparison of the calculated value for a heat exchange with 23% dense airside foam and the test data for 15.7 % dense foam shows only a 7.2% reduction in total heat exchanged. Therefore, the 15% foam reduced pressure drop to almost half, while sacrificing only 7.2% heat transfer. All of the above comparisons are at a sea-level approach velocity of about 14.4 ft/sec. As shown in Figure 16, the measured heat exchange at higher airflow, 220 scfm, agrees with the calculated heat exchange within about 10%. Heat exchange values at lower airflow, 110 scfm, agree within less than 1%. Air-to-liquid thermal performance testing of the six-element unit agreed with calculated heat exchange within 3.2 %. Air and liquid side heat balances agreed with each other within 2.7%. The six-element subpanel test results validate our expected thermal performance and design/sizing methodology. Early in the program, manufacturing feasibility was verified by fabrication of a 21-element subpanel, shown in Figure 12 above.

Foam Metal Heat Exchanger Optimization Study
Figure 17 summarizes heat removal, foam-metal core weight, and airflow trades for two fixed aluminum foam-metal densities. A 100-kW, 23% dense foam-metal core would weigh about 20.5 lb and require about 240 lb/min airflow. By comparison, a 100-kW, 8% dense foam-metal core would weigh about 11 lb and require about 600 lb/min airflow. The 8% dense panel, although lighter, requires more air mass flow and would have a larger heatexchanger-panel frontal area. Total airflow available is related to higher level system trades, such as aircraft flight conditions and practical limits on inlet-scoop frontal area. Those are in turn related to considerations such as drag force on pod and strength/ weight/size of attachments to the fuselage. Our selection of 15% dense foam represents a compromise between panel frontal area (which drives size and volume), weight, and total air mass flow rate. Compared with our initial 23% dense foam-metal concept, the panel with 15% dense foam is a few pounds lighter and somewhat larger face area. Optimization, to find a maximum or minimum, can be done on just one item at a time. For multidimensional problems, such as the parameters mentioned above, an optimization metric can be formed of the various factors. For example, a function consisting of a weighted linear sum of the parameters can be formed. Frequently, limits can be placed on
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6 5
Heat Exchanged (kW)
T60/40 , in = 150?F Tair, in = 53?F Tin = 97?F

4 3 2 1 0
Nominal 22.9% dense Heat exchanger with K&L–8 15.7% dense foam 220 scfm 159 scfm 110 scfm

0

5

10 15 Approach Velocity, Vinf (ft/s)

20

25

Figure 16. Six-element panel performance verification test results

400 350 300
8% Foam, Case B, Heat 23% Foam, Case C, Heat 8% Foam, Case B, Airflow 23% Foam, Case C, Airflow

3000 2500

250 200 150

2000

1500

1000 100 50 0 0 10 20 30 Weight (lb) 40 50 500 0 60

Figure 17. Foam metal heat exchanger trade study results

Air Mass Flow (lb/min)
123

Heat Removal (kW)

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the range of acceptable values for each parameter, simplifying the analysis considerably. Several numerical methods exist for finding the minima or maxim of such an expression, such as the simplex method. For a naval aircraft, LCS weighting factors were not available. Therefore, we relied on limited-scope informal trade-offs, considering primarily weight, size, shape, total airflow, and pressure drop. These are constrained or driven by pod volume and shape, depending on whether or not the pod is predefined.

Conclusions
Results of this project have laid the groundwork for foam-metal heat exchangers to be a significant weight saver candidate on a naval aircraft. Currently, a foam-metal heat exchanger is being evaluated for application on a naval aircraft. Foam-metal heat exchangers are used to best advantage where space and weight are at a premium. Foam metals can also be advantageous for cooling high-heat fluxes, such as high-power electronics. Foam metal fins are being evaluated for cooling high-power radar amplifier transistors. They have potential to lower junction temperatures by 10°C. Compact foam-metal heat exchangers are also being evaluated as part of an ONR technology demonstration program for cooling submarine electronic equipment racks in conjunction with a system using thermoelectric coolers. Application of foam-metal heat exchangers to real-time heat rejection from airborne laser DEW platforms is under active consideration. A heat transfer and flowresistance database for design of compressed foam-metal heat exchanger has been established as a result of this program. Results comparing the various combinations of foam compression direction indicate a potentially rich foam-metal heat exchanger design trade-off space in terms of heat transfer and pressure drop. These results were applied to the design of a heat exchanger for potential application to a naval aircraft. Test results of single-element and six-element subpanels verified our design methodology and heat exchanger performance expectations. Both analysis and laboratory testing have validated thermal performance of a ram air-cooled foam-metal heat exchanger on a naval aircraft.

Acknowledgments
This work was sponsored jointly by Northrop Grumman Corporation, the Office of Naval Research, and the Defense Advanced Research Projects Agency (DARPA).

References
1. 2. 3. W.M. Rohsenow and H. Choi, Heat Mass and Momentum Transfer, Prentice-Hall, Englewood Cliffs, N.J., 1961. D.Q. Kern and A.D. Kraus, Extended Surface Heat Transfer, McGraw-Hill, New York, 1972. B. Ozmat, Ozer Engineering/ERG, Inc, private communications.

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Author Profiles
John F. Klein is a Project Lead Engineer in Advanced Early Warning/Electronic Warfare (AEW/EW), Advanced Technology and Development Center business area, at Northrop Grumman, providing Project/Technical leadership of foam metal Heat Exchanger (HX) development and demonstration program; application of foam metals as heat transfer media for Northrop Grumman military aircraft, electronics, and power systems and directed energy weapon mitigation program for aircraft protection. He has 36 years of experience in research and development of fluid flow and heat transfer for a wide range of applications. Other recent projects include development of a solidification apparatus and control system for experimentation with the GASAR gas-assisted foam metal solidification process; thermal management of control and power electronics for electron-beam lithography systems, thermal spray process development. He is the author of more than 50 technical reports and publications. He holds 12 U.S. and several foreign patents related to thermal applications and processes. He holds a BS in mechanical engineering from Clarkson College of Technology; an MS in mechanical engineering and Engineer of Mechanical Engineering from the Massachusetts Institute of Technology; an MS in computer science from Rensselaer Polytechnic Institute; and a Software Engineering Certificate from the Polytechnic University of New York. john.klein@ngc.com Noe Arcas is a Senior Technical Specialist in the AEW/EW Advanced Technology and Development Center business area at Northrop Grumman, providing support of laboratory testing, instrumentation, data acquisition, vibration and acoustics. He has over 40 years of experience in the development of software and instrumentation for measuring and analyzing vibroacoustic environments, aircraft noise reduction, aircraft loads, foam metal, fluid flow, and heat transfer testing. Recent projects include the development and feasibility demonstration of foam metals as heat transfer media for Northrop Grumman military aircraft, the application of active control technology to reduce noise in commercial aircraft engines, and evaluation of in-flight loads for electronic jamming systems of EA-6B and F-18 aircraft. He is the author of over 20 technical reports and publications. He holds 10 U.S. patents in acoustics technology. He is a member and former president of the New York chapter of the

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Institute of Environmental Science and Technology and served as a member of the AIAA Aeroacoustics Committee. He is a registered professional engineer in the state of New York. He holds a BS in mechanical engineering from the City University of New York, an MS in mechanical engineering from New York University, and an MS in applied mathematics from Adelphi University. noe.arcas@ngc.com George W. Gilchrist is a Senior Technical Specialist in the AEW/EW Advanced Technology and Development Center business area at Northrop Grumman. He has 23 years of experience in detailed thermal and infrared modeling at Northrop Grumman, providing thermal modeling support to a wide range of aerospace applications. Most recently, Mr. Gilchrist has been program manager on a series of programs that use thermal modeling capabilities in the development of exploitation algorithms used in the interpretation of remotely sensed infrared imagery. He holds three U.S. patents in the field of thermal sciences and infrared imagery. Mr. Gilchrist holds a Bachelor’s of Engineering in mechanical engineering from the State University of New York, Stony Brook, and an MBA from Adelphi University. george.gilchrist@ngc.com William L. Shields, Jr., is an Engineer II in the AEW/EW Advanced Technology and Development Center business area at Northrop Grumman. He provides engineering support for test hardware installation, instrumentation set-up and application, and test operation. He has over 30 years of experience in the fabrication of military aircraft and laboratory testing. Most recently, he has focused on high-energy fluid mechanics, operating and maintaining a shock tunnel and aerosol tube for Northrop Grumman. Recent projects include the testing and measurement of foam metals and their heat transfer properties, static and vibration measurements of electronic pods for the EA-6B Prowler’s Improved Capability III Airborne Electronic Attack weapon system (ICAP III) program, and the fabrication and testing of a variable-area nozzle for military aircraft engines. He has conducted extensive testing and evaluation in the field of thermal imaging and nondestructive inspection of military aircraft. For those technologies, he has supported the Porous Metal Heat Exchanger program, Strategic Environmental Research

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and Development program (SERDP), the Environmental Security Technology Certification program (ESTCP), and the DD(X) Advanced Destroyer program. He is a high school graduate. william.shields@ngc.com Richard Yurman is currently an Engineering Specialist in the Thermodynamics group on the Advanced Hawkeye program in Northrop Grumman’s Integrated Systems sector. He has worked in the areas of thermodynamics, heat transfer, and infrared phenomenology on a number of aircraft programs, including the E-2D, EA-6B, F-14, A-6F, and Joint Surveillance and Targeting Attack Radar System (JSTARS), as well as for various engineering, integration, and test assignments for the AEW/EW Advanced Technology and Development Center. He has performed fieldwork to test and verify numerous thermal and radiometric models and has used thermal modeling techniques to develop algorithms to extract intelligence information from infrared imagery. He holds three patents on various thermal methods and systems and has authored and co-authored numerous papers and reports in scientific and engineering journals. He holds both BS and MS degrees in chemical engineering from the Cooper Union School of Engineering, New York. richard.yurman@ngc.com James B. Whiteside leads the intellectual asset management team for the Airborne Early Warning and Electronic Warfare Systems business area at Northrop Grumman. He has 36 years of experience in structures and materials engineering, research, and development. Recent projects include a structural integrity prognosis system, leadership of a program to investigate structurally integrated thermal management of airborne early warning and electronic warfare systems, accelerated insertion of composite materials, a composites affordability initiative, leadership of processing-science-based manufacturing design procedures for an autoclave, and damage tolerance and design allowables for the B-1A composite horizontal stabilizer. He is the author of more than 35 technical publications and a registered professional engineer in Louisiana. He holds a BS in mechanical engineering from Tulane University and a PhD in mechanical engineering from the University of Sheffield, England. james.whiteside@ngc.com

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