Experimental Investigation of Cooling Performance of Metal-Based Microchannels

Metal-based microchannel heat exchangers (MHEs) are of current interest due to the combination of high heat transfer performance and improved mechanical integrity. Efficient methods for fabrication and assembly of functional metal-based MHEs are essential to ensure the economic viability of such devices. Al- and Cu-based high-aspect-ratio microscale structures (HARMS) have been fabricated through molding replication using metallic mold inserts. Such metallic HARMS were assembled through eutectic bonding to form Al- and Cu-based MHEs, on which heat transfer tests were conducted to determine the overall cooling rate and time constants. Electrically heated Cu blocks were placed outside the MHEs and provided a constant flux, and water flowing within the microchannels acted as the coolant. Experimental results show a great influence of the type of metal, flow rate, and the surrounding conditions on the overall cooling performance of the MHEs.

• A 1 constant
• A 2 constant
• D h microchannel hydraulic diameter (μ m)
• T temperature (°C)
• t time (s)
• Greek Symbols
• τ 1 cooling time constant corresponding to the heat flux from the heater (s − 1)
• τ 2 cooling time constant corresponding to the heat flux from the surroundings (s − 1)

Science and technology related to new and more environmentally friendly ways of producing and distributing energy are being intensely investigated. New technologies capable of increasing the energy efficiency of transportation, housing, and appliance systems, if successfully commercialized, will reduce energy consumption and positively impact the global environment [[1]].

In 1981, Tuckerman and Pease pointed out that decreasing liquid cooling channel dimensions to the micrometer scale will lead to increases in heat transfer rates, and demonstrated a 40-fold improvement in heat-sinking capability experimentally in Si-based microchannels anodically bonded to Pyrex cover plates [[2]]. Since then, intense research on microchannel heat exchangers (MHEs) has ensued [[3]]. Many experimental heat transfer investigations have been conducted on Si-based microchannels [[4]]. The choice of Si-based microchannels does not result from the fact that Si is the optimal material for cooling devices in terms of its thermal and mechanical properties, but rather from the fact that fabrication technology for Si-based high-aspect-ratio microscale structures (HARMS) is the most mature and most widely available. Indeed, metals such as Al and Cu possess higher bulk thermal conductivities as compared to that of Si [[7]]. In addition, Al- and Cu-based MHEs promise much higher mechanical robustness over similar Si-based devices.

Various techniques for fabricating metal-based HARMS have been studied. Serial subtractive techniques, include micromilling [[8]], micro electrical discharge machining (μ EDM) [[9]], and laser-beam direct writing [[10]], are time-consuming and may encounter difficulties when fabricating microscale features in geometrical proximity. Micro powder injection molding requires multiple heat treatment steps, and suffers from incomplete mold filling and part shrinkage during heat treatment [[11]]. Micro casting also requires multiple heat treatment steps, and is an inefficient "double lost mold" technique [[13]]. These disadvantages make the techniques just mentioned less suitable for large-scale production of metal-based HARMS needed for building metal-based MHEs.

Molding replication offers an important alternative toward efficient fabrication of metal-based HARMS. The process of replication involves the use of primary HARMS, produced by a combination of lithography, etching, deposition, and other techniques, as a mold insert to create the negative of the insert pattern in metals by molding. One important example of molding replication is contained within the LiGA protocol for HARMS fabrication, combining deep x-ray/ultraviolet (UV) lithography (Lithographie) on polymeric resists, metal electrodeposition (Galvanoformung) into developed resist recesses, and molding replication (Abformung) of secondary HARMS [[14]]. Metal-based HARMS made by lithography and electrodeposition are expensive due to the high equipment and process costs of lithography and the low speed of electrodeposition. The importance of the molding replication step is that HARMS can be produced in different engineering materials without repeating the lithography/etching/deposition steps, and thus at low cost and high throughput.

Since 2003, successful HARMS replication by direct molding has been demonstrated in Pb and Zn [[15]], Al [[16]], and Cu [[17]]. Two critical elements were needed for successful molding replication of reactive metals such as Al and Cu. The first one is to control the near-surface chemical/mechanical interactions between the mold insert and the molded metal. This was achieved via conformal deposition of suitable ceramic coatings over HARMS mold inserts [[18]]. The second one is to improve the mechanical properties of the mold insert bulk at elevated molding temperatures. This was achieved by fabricating mold inserts out of refractory metals and alloys with μ EDM [[17], [22]]. Efficient fabrication of refractory mold inserts containing geometrically complex microfeatures was achieved through parallel μ EDM, in which complex micropatterns were transferred simultaneously onto surfaces of refractory metals/alloys using lithographically defined electrodes [[17], [23]].

Using these surface-engineered refractory mold-inserts, microchannels with complex geometries have been replicated successfully in metals such as Al and Cu by compression molding. Al-based microchannels were successfully bonded to other Al microchannels and plates by eutectic bonding with vapor-phase co-deposited Al–Ge [[24]] thin-film intermediate layers [[25]]. Multilayered microchannel arrays were successfully assembled to form metal-based MHE prototypes, with measured tensile bond strengths exceeding 75 MPa and reaching as high as 165 MPa [[26]].

Heat transfer characteristics of Cu-based MHEs have been measured previously by Lee et al. in Cu microchannels fabricated by mechanical cutting and sealed with a polymeric top plate by gluing [[27]]. In this paper, the process of fabrication and assembly of simple, entirely Al- and Cu-based, MHE prototypes is described.

Heat transfer in microchannels depends significantly on channel geometry. Extensive studies of microchannel flow and heat transfer have been carried out in the past two decades. The focus of these studies has been on characterizing the flow and thermal transfer at a fundamental level, while application development has been a more recent focus.

The first experimental results of fluid behavior in microchannels were obtained for trapezoidal shaped microchannels with hydraulic diameters 50 μ m [[28]]. The results agreed with smooth macro-channel theoretical values in both laminar and turbulent regimes. Gas and liquid flow in circular microtubes and measured friction factors also agreed well with theoretical values based on macro-channel situations [[29]]. It was also indicated that the laminar–turbulent transition occurred at a Reynolds number of around 2000. Several studies presented results for microchannels with hydraulic diameters ranging from 100 to 500 μ m, and showed good agreement with theoretical results [[31]]. Other studies provided some insight into gas and liquid flow inside microchannels and observed that friction factor results agree with macroscale friction laws [[35]]. Visualization of liquid flow through microchannels with diameters ranging from 244 to 974 μ m aimed at verifying onset of turbulence concluded that the development of turbulent flow in microchannels is retarded as compared to the macroscale systems [[39]]. Experimental investigations were conducted to explore the validity of classical correlations based on conventional channels for predicting the thermal behavior in single-phase flow through rectangular microchannels with widths ranging from 194 to 534 μ m, and with channel depth being nominally five times the width in each case [[40]]. The experiments were conducted for flow Reynolds numbers ranging from ∼ 300 to ∼ 3500 and used deionized water as the working fluid. A "checkmate" design of a micro heat exchanger with cross-flow was the first attempt at heat transfer enhancement in microchannels [[41]]. A micro pin fin heat exchanger fabricated using the LiGA process [[42]] indicated slightly higher enhancement as compared to a traditional macro pin fin channel, and indicated that there is need for optimization of the pin fin placement in the channel to further improve the level of enhancement. Lee et al. [[27]] studied a variety of rectangular microchannels made of Cu, ranging from 194 μ m to 534 μ m in width. They compared their experimental results against conventional correlations, indicating wide disparities between experiments and correlations.

Recently there have been several studies that have focused on various aspects of microchannel geometry to enhance heat transfer. Liu et al. [[43]] studied convective heat transfer in a quartz microtube with three different inner diameters of 242, 315, and 520 μ m. They indicated that the experimental values for Nusselt number matched well with the laminar flow heat transfer correlation. They also indicated laminar–turbulent transition in the range of Reynolds number between 1500 and 5500 for the microtubes. Steinke and Kandlikar [[44]] presented a comprehensive review of friction factor data in microchannels with liquid flow. They indicated that entrance and exit losses need to be accounted for while presenting overall friction factor losses in microchannels. Most of the data that accounted for friction factor loss shows good agreement with conventional theory. They also provided a new procedure for correcting measured pressure drop to account for inlet and exit losses.

In the present work, fabrication of metallic MHE by molding replication is discussed. Additionally, heat transfer tests have been performed to investigate the overall cooling performance of such devices using infrared thermography. Cooling rates are evaluated for both Al- and Cu-based MHE devices.

Microscale mold inserts were fabricated from nickel-based superalloy Inconel X750 plates. As-received Inconel plates were machined to square insert blanks, with an active area of ∼ 15,000 μ m × ∼ 15,000 μ m, ∼ 3200 μ m in height. The top surface of the blank was mechanically polished with SiC abrasive papers down to 1200 grit size. Insert fabrication from the blanks involved three main steps: μ EDM of the active area, electrochemical polishing (ECP) of as-machined microscale Inconel features, and deposition of a conformal amorphous silicon nitride (a-Si:N) coating over electrochemically polished microscale features. A SARIX high-precision micro-erosion machine (model SR-HPM-B) was used for insert μ EDM. Flat molybdenum (Mo) sheets with a thickness of 500 μ m were used as blade electrodes. A series of parallel cuts was performed with the Mo electrodes, resulting in an array of parallel rectangular microprotrusions on the insert active area. As-machined Inconel blanks were electrochemically polished for a total of 10 min with current control in a mixed acid solution of HClO4 (70%)/CH3COOH (80%) at a volume ratio of 1:1. Following ECP, a conformal a-Si:N was deposited over the Inconel inserts in a radiofrequency (rf) inductively coupled plasma (ICP)-assisted hybrid chemical/physical vapor deposition system [[45]]. A scanning electron microscopy (SEM) image of one such a-Si:N-coated Inconel insert is shown in Figure 1. The active surface of this insert contains 19 parallel rectangular microprotrusions, with an average width of ∼ 154 μ m, a height of ∼ 400 μ m, and a center-to-center spacing of ∼ 750 μ m. The thickness of the conformal a-Si:N coating is about 0.6 μ m. Further details on insert fabrication, coating deposition, and characterization of conformal a-Si:N coatings have been reported previously [[18], [21]].

Graph: Figure 1 SEM image of a Si:N coated Inconel insert.

Al 6061 (90 at.% Al, 4 at.% Si, 2 at.% Mg) and Cu 110 (99.9+ at.% Cu) coupons, with the same geometry of 35.5 mm × 35.5 mm square and 6.4 mm in thickness, were molded at high temperatures with one single Inconel insert coated with an a-Si:N coating. The molding process replicates the array of parallel rectangular microprotrusions on the insert into an array of parallel rectangular microchannels on the molded Al and Cu coupons. Before molding, the top surfaces of Al and Cu coupons were mechanically polished to less than 1 μ m in surface roughness. Compression molding was carried out in an MTS858 single-axis testing system interfaced to a custom-built high-vacuum, high-temperature, instrumented molding apparatus. A linear actuator could be programmed to move according to prescribed load forces in the force-controlled mode or actuator displacements in the displacement-controlled mode. Total axial force and axial displacement were measured continuously during the entire molding and demolding process. All molding runs included five segments: a fast approach followed by a slow approach in displacement control, a constant-loading-rate molding followed by a 15-s constant-force hold in force control, and a constant-rate demolding in displacement control. Further details on the molding apparatus and the molding protocol are given elsewhere [[16], [21]].

One fluid supply channel and one fluid drain plenum were cut into replicated arrays of parallel Al and Cu HARMS replicated by molding, which were further fabricated using μ EDM to cut inlet and outlet channels and using a drilling machine to make connection holes. Then these holes were typed to connect plastic adaptor. Surface morphologies of molded HARMS were examined on a Hitachi S3600N scanning electron microscope (SEM). Figure 2a and Figure 3a show the fabricated Cu and Al HARMS, respectively. Two inlets are connected by a μ EDM-ed channel, which also connected to the microchannels. On the other end, a wide channel is μ EDM-ed as a pool to connect three outlets, one in the center of the active area and the other two on its sides. Figure 2b and Figure 3b show SEM views of one of the inlets. The μ EDM-ed inlet channel of each structure has sharp sidewalls and typical μ EDM-ed surface morphologies, and well-connected microchannels and inlet holes. For Al structure, some dark spots can be found in the μ EDM-ed channel. These spots were produced during the EDM process due to the relatively lower melting temperature. Figure 2c and Figure 3c show high-magnification views of molded Cu/Al microchannels. The microchannel has vertical sidewalls and sharp sidewall-to-bottom transitions. It is clear to see the typical sidewall and bottom surface roughness due in part to the sidewall roughness of the rectangular microprotrusion on the Inconel insert. Because of the same insert, both of them show similar surface morphologies and the same surface roughness in size.

Graph: Figure 2 Close views of the copper micro heat exchanger base [[46]].

Graph: Figure 3 Close views of the aluminum micro heat exchanger base.

Al/Cu HARMS and flat coupons were mechanically polished again to get a smooth surface for better bonding quality by using Al–Ge intermediate layers, which were deposited in the ICP assisted hybrid system. Al–Ge composite films were co-deposited in pure Ar (99.999%), with a total pressure of ∼ 1.3 mtorr, from two separate sputter sources, one for pure Al (99.99%) and the other for pure Ge (99.99%). The polished substrates, Al/Cu HARMS and flat coupons, were ultrasonically cleaned in acetone and methanol before being mounted on a rotatable holder situated in the middle of the deposition zone. The deposition sequence consisted of a substrate surface etch followed by co-deposition of Al and Ge. Substrate etching occurred in a pure Ar ICP with a total rf input power of 1000 W, a substrate bias of –100 V, and an etch duration of 20 min. Sputtering of Al and Ge cathodes commenced immediately after substrate surface etch. Substrates were rotated continuously at about 12 rpm during both etching and deposition. All Al–Ge depositions were carried out with a fixed Al cathode current of 1.0 A and a fixed Ge cathode current of 0.45 A, respectively. This deposition composition resulted in a Ge to Al composition ratio close to that corresponding to the Al–Ge eutectic composition of Al70Ge30. The substrates bias during deposition was fixed at –50 V. The deposition duration was 60 min, corresponding to a film thickness of ∼ 2.0 μ m. Further details on Al–Ge sputter co-deposition, structural characterization of Al–Ge nanocomposite thin films, and evaluation of bond strength have been reported in our previous papers.

Bonding experiments were carried out using the MTS858 single-axis testing system interfaced to a custom-built high-vacuum chamber. One heating station was mechanically attached to the bottom of the vacuum chamber. A second heating station was mechanically attached to the top linear actuator of the MTS858 testing system through a bellow-sealed motion feedthrough. The two heating stations were heated separately by resistive heating cartridges and the temperatures were measured by two separate K-type thermocouples. Further details on this high-vacuum, high-temperature bonding system have been given in our previous paper.

Flat Al/Cu coupons and HARMS with Al–Ge composite films deposited on the bonding surfaces were placed face-to-face on the lower heating station. The chamber was evacuated, and both heating stations were heated. Al HARMS/coupon bonding was carried out at the temperature of 500°C with an applied pressure of ∼ 1.5 MPa, while Cu HARMS/coupon was bonded at the temperature of 540°C with an applied pressure of ∼ 3 MPa. Figure 4 shows an optical overview of bonded Al and Cu devices connecting with plastic adaptors.

Graph: Figure 4 Assembled view of the copper and aluminum MHE devices [[46]].

To get exact dimensions of bonded Al and Cu devices, we cut these two devices after we finished the heat transfer experiments. Figure 5a and Figure 6a show the cross-section views of portions of the bonded microchannel structures, obtained by mechanical cutting perpendicular to the microchannel arrays. The entire structure contains 19 microchannels in each device. However, there is one channel blocked in the Al device, meaning the Al device has 18 through microchannels. Figure 5b and Figure 6b show higher magnification views of two typical bonded microchannels. The bonding interface is not discernable in this higher magnification view, indicative of the quality of bonding achieved. From the SEM picture, the average dimension of each microchannel can be obtained, as shown in Table 1.

Graph: Figure 5 Cut away view of the Cu microchannel after top plate bonding [[46]].

Graph: Figure 6 Cut away view of the Al microchannel after top plate bonding.

Table 1 Dimensions of Cu and Al devices

Figure 7 shows a schematic of the testing apparatus designed and built to study the thermal behavior of the micro heat exchanger fabricated using the already discussed procedure. The apparatus consisted of three sections: pressuring section, test section, and data acquisition section. For the experiments, the micro heat exchanger was used as a cooling device with water as the coolant. The first section consisted of a water holding tank connected to a compressed air bottle and was used to obtain sufficient pressure to push the coolant through the micro channels. This approach was preferred to the use of pump and provided a smooth and stable flow. A needle valve placed at the bottle neck helped release the required amount of air from the bottle and that placed downstream of the tank exit was used to make fine adjustments to the flow rate and attain smooth and stable flow through the channels. A pressure relief valve was also connected to the tank for safety measures.

Graph: Figure 7 Schematic of the experimental setup [[46]].

The test section consisted of either an Al- or a Cu-based specimen enclosed in a rigid polyvinyl chloride (PVC) foam insulation with the heater in contact with the top face of the specimen. The Al specimen consisted of 18 parallel channels with an average cross-sectional dimension of 291 μ m × 114 μ m, while the Cu specimen had 19 parallel channels with an average cross-sectional dimension of 319 μ m × 150 μ m. Figure 8 shows a detailed view of the test section. Two rectangular heaters were used. Each rectangular heater was fabricated by inserting a super watt cartridge heater (200 W) into a single machined copper block (5/8 × 5/8 inch) of same length(1/4 inches). The two copper blocks were bonded together, then attached to the specimen with the help of highly conducting arctic silver glue and then the cartridge heaters were inserted. Before attaching, the surface of the heater was polished and smoothed. Also, the arctic silver glue provided a uniform contact and eliminated the air gaps between the heater and specimen surfaces. Holes were drilled into the insulation, near the bottom surface of the specimen, for the inlet and outlet tube connections. In order to measure the inlet and the outlet temperatures, 36 gauge K-type thermocouples were inserted into the inlet and the outlet plenum through the inlet and the outlet tubes (Figure 9). Other K-type thermocouples were placed at the top surface, bottom surface, and heater surface. An Instrunet data acquisition system connected to a PC (personal computer) was used to collect the thermocouple read-outs. The differential pressure across the specimen was measured using a Dwyer digital manometer with a least count of 0.1 psi.

Graph: Figure 8 Detailed views of the test section (not to scale): (a) bottom view, (b) top view, and (c) cross-sectional view.

Graph: Figure 9 Schematic showing the placing of the inlet thermocouple.

Constrained by the sample dimensions, a maximum of two super watt heaters were used with a maximum input flux of 37 W/cm2. The volumetric flow rate was obtained by using a calibrated flow meter as well as by timing the amount of water collected at the exit over a fixed period of time. The inlet and the outlet fluid temperatures were measured with the help of five thermocouples—two at the inlet and three at the outlet.

The two specimens were tested under similar conditions of temperature and pressure and the same experimental procedure was followed. First thermocouples were inserted into each inlet and tube with the help of a T-fitting and then sealed with epoxy (Figure 9). In the same fashion, thermocouples were inserted into each outlet tube but the epoxy sealing was not required. The tubes along with the thermocouples were then connected to the inlet and outlet of the device. Another thermocouple was placed at the bottom surface of the device, which was then placed in the insulation. Four other thermocouples and the heater were placed on the top surface of the device. Further, connections to measure the differential pressure and flow rate were made. These settings were maintained the same for both the specimens.

Infrared thermography was used to determine the surface temperature of the MHE device to determine the overall cooling rate and time constant for the device. The surface was coated black to improve emissivity up to 0.96. The test device was initially heated to a known temperature and cooler water was suddenly introduced into the test channels. The surface temperature response was captured at 60 Hz capture rate with the FLIR SC500 camera. The embedded thermocouples were used to calibrate the infrared (IR) response in situ. In each IR image a rectangular area was marked and the average temperature of that area was used to plot the transient temperature profile.

Figure 10 shows the surface temperature drop as coolant is suddenly allowed into the MHE. Initially, the heater is turned on and the MHE is allowed to achieve a steady temperature around 100°C. The coolant is suddenly introduced into the MHE and the surface temperature is monitored. It is clear that there is a significant temperature drop in the first 2–3 s from 100 to 40°C. Then there is a gradual decrease from 40 to 30°C in the next 60 s. The first drop in temperature occurs when the coolant draws the heat away from the copper or aluminum. The second gradual drop occurs when the heat is drawn from the surrounding insulation. Since the insulation has high heat capacity, this results in a slower gradual decrease in temperature. However, the entire device cools down to acceptable temperatures within 60 s of coolant introduction.

Graph: Figure 10 Surface temperature response of the MHE in a transient cooling mode [[46]].

Based on the preceding hypothesis, the transient temperature profile for all the test cases were normalized and then approximated with the following decaying formula:

Graph

where T(t) is the normalized transient temperature, τ1 and τ2 are the cooling time constants corresponding to the heat flux from the heater and surroundings, respectively, and the constants A1 and A2 give an idea about the magnitude of these different heat fluxes.

Figure 11 shows the normalized cooling rate curve that has been fitted using Eq. (1). The perfect fitting of the cooling rate curve completely supports the existence of two different time constants. Similar trend was observed for both the devices and for all the test cases. The constants for all the test cases have been reported in Table 2 and Table 3.

Graph: Figure 11 Fitting of the normalized transient cooling curve [[46]].

Table 2 Constants for Al-based MHE

Table 3 Constants for Cu-based MHE

Figure 12a and Figure 12b present the time constant based on the transient cooling mode behavior of the MHE. The results show that the time constant for the device decreases with increasing cooling flow rate, as expected. Time constants corresponding to the heat flux from the heater are in the range of 1 to 2 s for the copper device and 2.5 to 2 s for the aluminum device. Those corresponding to the heat flux from the surrounding insulation show a similar trend but are one order higher in magnitude. Such reduced response is due to a large difference in the thermal conductivities of the insulation and the metallic device. Moreover, the aluminum device response is slower than the copper device due to the basic difference in thermal conductivity of the metals. Copper has higher thermal conductivity and thus results in lower time constants.

Graph: Figure 12 MHE time constant based on transient cooling mode results: (a) τ1 [[46]] and (b) τ2.

Figure 13 shows the transient cooling mode behavior of metallic MHE from the initial state to the steady state. At instant 1, the heaters were switched on to provide an input power of 370 W. The surface temperature was constantly monitored with help of thermocouples placed on the surface was allowed to rise. As soon as the surface temperature reached 100°C, the valve was opened to push the water through the microchannels (instant 2) and within a minute the system reached steady state (instant 4). Infrared images were taken for the whole duration at 60 Hz capture rate and processed to obtain the cooling curve (Figure 13). Figure 14 shows the thermographs corresponding to each instant. The thermographs at instants 2 and 3 give an idea of the cooling effectiveness of the metal-based microchannel device. The time difference between instant 2 and instant 3 is just about a second. The fourth thermograph indicates the rise in coolant temperature as well as even heating of the coolant as it passes through the microchannels.

Graph: Figure 13 Temperature profile from initial state to steady state.

Graph: Figure 14 Thermographs at different instants: (a) t = 5 s, (b) t = 33.091 s, (c) t = 34.042 s, and (d) t = 60 s.

Fabrication of Al- and Cu-based high-aspect-ratio microscale structures (HARMS) through molding replication using metallic mold inserts has been demonstrated. Such metallic HARMS were assembled through eutectic bonding to form Al- and Cu-based MHEs. Heat transfer tests were conducted to determine the overall cooling rate and time constants. The time constant for the MHE device was found out to be lower for Cu channels with times around 1–2 s response. Aluminum is only slightly lower due to the lower thermal conductivity. Surrounding conditions were observed to play an important role by considerably increasing the cooling time.

The demonstration of fabrication and thermal performance for metal-based MHE devices has enormous potential in many heat transfer problems, primarily in avionics and electronic cooling. It is clear that the use of metal-based micro heat exchangers may result in significant improvement in cooling efficiency for microelectronic systems.

Graph

Pritish R. Parida received his B.Tech. degree from Indian Institute of Technology Guwahati (IITG), Guwahati, India, in 2006 and his M.S. degree from Louisiana State University, Baton Rouge, in 2007, both in mechanical engineering. He is currently a Ph.D. candidate in the Department of Mechanical Engineering, Virginia Tech. His research interests include thermal packaging for power electronics, cooling systems, investigation of enhanced heat transfer designs, and modeling and simulation of micro-scale conjugate heat transfer problems using the lattice Boltzmann method.

Graph

Fanghua Mei received the B.S. and M.S. degrees in materials science and engineering from Shanghai Jiao Tong University, Shanghai, China, in 2002 and 2005, respectively. He is currently a Ph.D. candidate in the Department of Mechanical Engineering, Louisiana State University, Baton Rouge. His research interests include thin film deposition, fabrication of metal-based MEMS, and testing of assembled MEMS devices.

Graph

Jing Jiang received the B.S. and M.S. degrees in the Department of Mechanics and Mechanical Engineering from the University of Science and Technology of China, Hefei, Anhui, China, in 1999 and 2002, respectively. He is currently a Ph.D. candidate in the Department of Mechanical Engineering, Louisiana State University, Baton Rouge. His research interests include fabrication, modeling, and simulation of metal-based MEMS.

Graph

Wen Jin Meng received the B.S. degree in physics with honors from California Institute of Technology (Caltech), Pasadena, in 1982 and the Ph.D. degree in applied physics from Caltech in 1988. He served as a postdoctoral research fellow in the Materials Science Division of Argonne National Laboratory from 1988 to 1989 and as a staff research scientist in the General Motors/Delphi Automotive Systems R&D Center from 1989 to 1999. Since 1999, he has been with Louisiana State University, Baton Rouge, where he is currently the Gerald Cire & Lena Grand Williams Professor in the Department of Mechanical Engineering. His research spanned topics concerning solid-state phase transformations, heteroepitaxial growth of metal and ceramic thin films, plasma-assisted vapor-phase deposition processes, nanostructured ceramic coatings, and surface engineering. His recent research focuses on the materials aspects of microscale fabrication and assembly, especially on metal-based microsystems.

Graph

Srinath V. Ekkad received his B.Tech. degree from JNTU in Hyderabad, India, in 1989 and then his M.S. from Arizona State University in 1991 and Ph.D. from Texas A&M University in 1995, all in mechanical engineering. He was a research associate at Texas A&M University and a senior project engineer at Rolls-Royce, Indianapolis, before he joined Louisiana State University as an assistant professor in 1998. He moved to Virginia Tech as an associate professor of mechanical engineering in fall 2007. His research is primarily in the area of heat transfer and fluid mechanics with applications to heat exchangers, gas turbines, and electronic cooling. He has written more than 100 articles in various journals and proceedings and one book on gas turbine cooling. His research focuses on enhanced heat transfer designs, with a variety of applications.