MANY OF THE micro heat engines mentioned in Part I

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1 326 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 A Silicon Microturbopump for a Rankine-Cycle Power-Generation Microsystem Part II: Fabrication and Characterization Changgu Lee, Mokhtar Liamini, and Luc G. Fréchette, Member, IEEE, Member, ASME Abstract In Part I of this two-part paper, the design approach for a microturbopump was presented. This second part describes the fabrication and experimental characterization of the demo microturbopump device, which includes hydrostatic bearings, a spiral-groove viscous pump, and a multistage microturbine. The device is composed of five wafers: one glass wafer, one siliconon-insulator (SOI) wafer, and three silicon wafers. The silicon and SOI wafers are patterned using shallow and deep reactive ion etching (total of 14 masks), while the Pyrex glass wafer was ultrasonically drilled. Anodic bonding, fusion bonding, and manual assembly with alignment structures were then used to complete the device and enclose the 4-mm-diameter rotor. The approach allowed the microfabrication of unique interdigitated blade rows in the microturbine and interchangeable parts for flexible testing. After completion of the device, bearings were first tested in static and dynamic conditions. Then, the turbine was characterized with compressed air only and spun up to r/min, producing 0.38 W of mechanical power. The pump performance map was also completely characterized for speeds up to r/min, showing a maximum pump flow rate of 9 mg/s and maximum pressure rise of 240 kpa. In a turbopump system performance test using compressed air to the turbine and water in the pump, the rotor was spun up to r/min, which corresponds to 25 m/s in tip speed. At this condition, the turbine produced W of mechanical power with 41 kpa of differential pressure and 24 mg/s of flow rate, and the pump pressurized water by 88 kpa with a flow rate of 4 mg/s, maintaining constant efficiency of 7.2% over the operating range. Out of the total power produced by the turbine, 10% was consumed by the viscous pump, while the rest was dissipated by other components through viscous drag. The system-level predictions by models introduced in Part I also match the measured performance, suggesting that a valid design basis has been established for this type of rotating micromachine. [ ] Index Terms Heat engines, microfluidics, micropumps, power generation, turbines. Manuscript received March 19, 2010; revised September 10, 2010; accepted October 24, Date of publication December 20, 2010; date of current version February 2, This work was supported in part by the NASA Glenn Research Center Alternate Fuels Foundation Technologies Program under Contracts NAS and NAS , monitored by Dr. Glenn Beheim, and in part by the Basic Science Research Program through the National Research Foundation of Korea funded by the Ministry of Education, Science and Technology under Grant Subject Editor L. Lin. C. Lee is with the Department of Mechanical Engineering, Sungkyunkwan University, Suwon , South Korea ( peterlee@skku.edu; cl2051@columbia.edu). M. Liamini and L. G. Fréchette are with the Department of Mechanical Engineering, Université de Sherbrooke, Sherbrooke, QC J1K 2R1, Canada ( Mokhtar.Liamini@USherbrooke.ca; lucf@alum.mit.edu). Color versions of one or more of the figures in this paper are available online at Digital Object Identifier /JMEMS I. INTRODUCTION MANY OF THE micro heat engines mentioned in Part I of this paper [1] are adopting deep reactive ion etching (DRIE) and anodic- or fusion-bonding techniques to form quasi 3-D structures. These techniques enable moving components with wafer-level thicknesses and flow structures with high aspect ratios such as bearings, channels, deep holes, pistons, and rotors. Single crystal silicon as a building material offers great structural characteristics for high-speed rotating machinery due to its high strength-to-density ratio and can be micromachined to high precision [2]. Among the power-microelectromechanical-systems (MEMS) devices fabricated using silicon microfabrication technology, several turbomachinery-based microengines with configurations similar to the current microturbopump have been experimentally demonstrated to date. Fréchette et al. [3] have developed a radial-flow single-stage microturbine with a 4.2-mm-diameter rotor and operated up to 1.4 million r/min, producing about 5 W of mechanical power. Microturbochargers that combine a compressor driven by a turbine have been demonstrated by Kang et al. [4] up to r/min and Savoulides et al. [5] up to r/min. A turbomachinerybased microturbopump for high flow rates has also been developed, reaching r/min and delivering 200 mg/s of water at 75 kpa, although with unsteady behavior. The performance was mainly limited by sealing limitations, bearing issues, and pump-flow instabilities [6] [8]. In all these devices, only singlestage turbomachines were used, limiting the potential pressure ratio in the cycle. In this study, microturbopump devices based on a spiralgroove viscous pump and a multistage turbine were fabricated using the silicon microtechnology for the Rankine Microturbine power-generation application. The turbopump system consists of a turbine, pump, thrust bearings (TBs), journal bearing (JB), and seals. It was designed in Part I of this paper to produce 4.7 W of mechanical power at 260-m/s tip speed, pressuring water by 4.5 atm. The device is made of five wafers of glass and silicon, which are patterned through ultrasonic drilling or DRIE and are anodically or fusion bonded. The fabricated device, which has three parts, is then manually assembled using alignment structures. This paper presents the fabrication process for the microturbopump and addresses its unique challenges. The fabricated device was also completely characterized and demonstrated at the component and system levels. Since /$ IEEE

2 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 327 bearing performance is critical for rotor operation, it was first characterized in static and dynamic conditions. Rotor tests at high speeds were performed to identify the operating range of the current device using only gas as the working fluid. Then, the pump was characterized with water at various speeds to define the pump maps. Performance of the whole microturbopump system was also extracted from additional tests. Each characterization process includes verification of the models developed in Part I by comparing them with the experimental data. Initial demonstration of this device and partial results were presented in [9]. II. MICROFABRICATION The complete device is composed of five wafers, as shown in Fig. 1: one Pyrex 7740 glass wafer (wafer A), one siliconon-insulator (SOI) wafer (wafer B), and three silicon wafers (wafers C, D, and E). The thickness of the silicon wafers is 450 μm, and that of the glass wafer is 500 μm. The SOI wafer consists of a 400-μm-thick substrate and a 50-μm-thick device layer divided by 0.5 μm of oxide. The process was done on 100-mm wafers, on which 12 devices (dies) were fabricated from the batch process. The first wafer (wafer A) has holes for the turbine inlet and outlet, pump outlet, pressure taps, and the optical probe. The second wafer (wafer B) has flow channels and the stator blades of the turbine. The rotor is formed in the third wafer (wafer C) by etching the turbine rotor blades and the JB. The fourth wafer (wafer D) has the TB, pump, seal, pressure tap hole, and flow channels. Finally, the fifth wafer (wafer E) has channels and holes to supply or drain the bearing and pump flows. A. Fabrication Process 1) Fabrication Process Overview: The microturbopump structures are fabricated using photolithography and etching techniques, with a total of 14 photomasks, 4 shallow silicon etching steps, 10 deep silicon etching (DRIE) steps, and 3 plasma-enhanced chemical vapor deposition (PECVD) oxide depositions. In the fabrication process, having proper recipes for deep etching is critical to achieve the design dimensions and best performance of the device. The DRIE processes (Unaxis 770) consist of three steps: one deposition and two sequential etches. In the deposition step, mainly C 4 F 8 is supplied to the etch chamber, and SF 6 is supplied for the first and second etch steps. The inductively coupled plasma (ICP) power for the whole etching process is fixed. Two DRIE recipes were used for the device fabrication: one (recipe A) for large open areas (aspect ratio of depth to width < 1) and the other (recipe B) for trenches with high aspect ratio (depth to width > 10). Details of the DRIE processes are found in Table I. The main differences between the two are the duration period and RF power. Recipe B produces deep trenches of straight vertical walls up to aspect ratios of around 30 due to longer etching duration and higher RF power for etching, which enhances directionality of ion etching. The etch verticality for recipe B was sufficiently high that there was less than 10% of feature-size difference between the etch opening and the bottom of the trench. Etch depths from DRIE processes were measured by a profilometer and had differences from the design of up to 5 μm for the deepest etches. The turbine-statorblade etching was however controlled more precisely (1 μm) since the etch stops on the buried oxide layer of the SOI wafer, so the etch depth uncertainty is defined by the silicon layer thickness precision (refer to wafer B, step 3 in the following). For shallow etches, error of the depths was controlled within 0.1 μm. The fabrication process is as follows and is shown in Fig. 2 (further details of the fabrication process can be found in [10]). Wafer A: The holes are formed by ultrasonic drilling (Sonic- Mill). Wafer B: Using a nested mask of oxide under photoresist (PR), the top side is deep etched at two different depths (recipe B). After the first deep etch, the PR mask is removed and etching continues with the oxide mask until the buried oxide layer is reached. The other side is then patterned and deep etched to the buried oxide layer (recipe A). Then, the oxide layer is removed by buffered oxide etchant (BOE) to complete the through-wafer channels. The nested etching technique is useful when multiple deep etching processes are required on one side of the wafer because lithography is difficult over the topology of a deep-etched surface. Wafer C: The top side receives a shallow etch at first and a deep etch to form the rotor blades (recipe A). Then, the oxide is deposited by PECVD to protect the surface during the through etch from the other side. Then, a glass handle wafer is attached with PR on the blade side to prevent the rotor from falling in the chamber when it is cut out by deep etching. The JB is defined by a high aspect ratio (20 : 1) through a wafer circular trench etched from the bottom side (recipe B). The rotor is extracted and kept separately from the wafer after removing the oxide layer by BOE. Wafer D: At first, the top side receives four photolithography and shallow etch steps to form the bearing restrictor, seal, and pump grooves. Subsequently, the bottom side is etched halfway through (recipe A), with the top side covered by an oxide layer for surface protection. Finally, the top side is patterned and etched until the auxiliary TB nozzles are completely etched through (recipe B). Wafer E: The wafer is etched halfway through from both sides to create channels and through holes (recipe A). After the wafers are completed, they go through cleaning processes before bonding by removing the oxide with BOE 6:1 followed by metal oxide semiconductor (MOS) (or RCA) clean (10 min of base clean and 10 min of acid clean). Glass wafer A and silicon wafer B are aligned and anodically bonded (25 N, 1000 V, 350 C, 10 min). Wafers C, D, and E are also aligned and fusion bonded into a three-layer stack (1700 N, 30 s, vacuum) and annealed (1000 C, N 2, 1 h). Then, the stacks are diced into 12 dies (15 mm 15 mm) for manual assembly. B. Fabrication Challenges Interdigitated Turbine Blades: When the blades of the turbine are interdigitated during assembly, a clearance is required at the blade tips to prevent contact with the opposing etched surface. However, the rate of etching by DRIE is slower in

3 328 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 Fig. 1. Exploded view of the five-wafer-stack device and 3-D view of assembled device with internal flow paths [A 2-D view can be found in Part I (Fig. 2)]. (a) Top view. (b) Bottom view. (c) Flow paths through the assembled device. TABLE I DEEP REACTIVE ION ETCH PROCESSES the areas near walls than in open areas, as shown in Fig. 3(a); thus, the tips of blades could interfere with the surface near the adjacent row of blades. Also, protrusions between the blade rows due to micro etch masking by contaminants can cause similar problems. Therefore, additional clearance was given compared with the original design (1.5 μm). To ensure proper operation of the rotor, the rotor blades were etched μm, adding μm from the design etch depth [Fig. 3(b)], which defines the flow-path height for both rotor and stator. This will increase the stator tip clearance ratio (gap between the stator tip and the wall to the stator-blade height) from a design value of 3% to about 40%. This will reduce the efficiency and power of the turbine significantly because the flow portion through the clearance (called tip leakage flow ) does not undergo turning

4 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 329 Fig. 2. (continued.) Fabrication process flow of microturbopump (prior to assembly). Fig. 3. Nonuniformity of DRIE and overetch of rotor blades. Nonuniformity depends on feature and recipe of DRIE. (a) Typical surface after DRIE. (b) Overetch of rotor blades. Fig. 2. Fabrication process flow of microturbopump (prior to assembly). (i.e., no power extracted), but leads instead to mixing (i.e., more losses). The convex shape formed from the DRIE process could hardly be avoided although it can be improved to a certain level. One way to reduce the loss from the large leakage flow will be to increase blade heights so that the clearance ratio can be reduced. The 3-D effect of loss from the large clearance would need further study to estimate the performance of the turbine accurately. High-Aspect-Ratio Bearing Features: In the original design, there was supposed to be orifice holes for the main TB on wafer C (Fig. 4). Etching these holes turned out to be challenging due to the extremely slow etch rate derived from the high aspect

5 330 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 Fig. 4. Modification of TBs due to a challenge in etching narrow and deep holes. (a) Original design. (b) After fabrication. The main TB nozzles are not implemented in the current device, and the auxiliary TB nozzles are shortened. ratio and small feature area. Even with a substantially thick etch mask (12 μm of oxide), it was not possible to etch the small orifices through the wafer (Fig. 4). As fabricated, the device must therefore solely rely on its auxiliary TB for axial stiffness. Furthermore, the orifice depth of the auxiliary TB was reduced from 150 μm to85μm due to the low etch rate of the 10-μm holes compared with other features in the same layer. The etch rate of the holes appears to be 1/3 that of large features, so the features with large areas in the same mask, therefore, may be etched through the 450-μm-thick wafer before the orifices reach 150 μm deep. Hence, this change was inevitable for the given situation. Lithography Over Sharp Topologies: Due to the existing structures from four shallow etches on the front side of wafer D, the thickness of the PR spun on the surface was not even. After postbaking the PR, the thickness on the sharp corner of the features appeared to be thinner than in the planar areas. During etching, features with sharp corners such as the rim for the main TB and ridges of the spiral grooves of the pump and the seal, whose heights are 12 and 6 μm, were etched away. This occurred even with a 12-μm-thick PR. In order to prevent this undesirable damage, a 1-μm-thick SiO 2 film was deposited with PECVD as the mask. Because the PECVD process creates a conformal thickness regardless of the shape of features (provided they are not deep with high aspect ratio), this helped keep the extruded features undamaged during the etching process. Rotor Damage From Improper Handling: Because the rotor is separated from the wafer, much attention should be paid to handling it because the blades are fragile. Fig. 5 shows the result of mishandling the rotor. Mostly, blades in the fourth row were damaged because the tweezers may slide when grasping the rotor on the side wall. For handling the rotors, soft Teflon tweezers are recommended although there could still be some blade damage. Fig. 5. Rotor with broken rotor blades due to mishandling. Fig. 6. Cross section of five-wafer-stack microturbopump device, showing the main components [speed bumps are for measurement of rotational speed (refer to Section III-A)]. The top and bottom pieces are slightly separated for visualization. C. Device Assembly The device consists of an A B stack (top piece), a rotor, and a C D E stack (bottom piece), as shown in Figs. 6 and 7. The three pieces are assembled as shown in Fig. 7(b) for testing. The device is manually assembled using a pin-in-hole alignment structure, as shown in Fig. 8. Thin wires (or dowel pins) are inserted into the through holes of the bottom half first. The rotor is then dropped into position (rotor chamber), and the top half is aligned and lowered. Gross mechanical alignment is provided by the thin wires inserted in the through holes. Once the chips are in contact, the protruding silicon cylinder on wafer Fig. 7. Optical photograph of the complete pieces. (a) Three parts of the whole device. (b) Device assembly. B enters the triangular hole in wafer C. The alignment accuracy, which depends on the design size of the alignment features, was approximately 5 μm. The infrared (IR) photograph in Fig. 8(b) shows good alignment postassembly without interlocking of rotor and stator blade rows. Fig. 9 shows the assembled chip set and the interdigitated turbine blade rows. Although the blades

6 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 331 Fig. 8. Chip assembly process. (a) Cross-sectional schematic of chip assembly. (b) IR image of the assembly. The triangle feature corresponds to the trench in the bottom half, and the donut-shape feature is the silicon cylinder in the top half. The bright circle is the through hole for gross mechanical alignment. Fig. 10. Broken stator blades by quick air blow into turbine inlet with rotorstator misalignment. (a) Broken stator blades on the rotor. (b) Damaged stator with missing blades. Fig. 9. Assembled chip set. (a) Chip-scale view of the 15 mm 15 mm device. (b) IR image of blade rows. In (a), the 12 circular holes are ports for the supply of working fluids and measuring operational conditions. In (b), darker blades are rotors and brighter ones are stators. were fragile, particularly those of the stator, damage to the blades was not observed during the assembly. The benefits of using the manual assembly approach are as follows. First, the risk from multiwafer-stack fusion bonding is reduced. Second, it is simple to check when the device has a problem by opening it. Third, the rotor can be easily replaced by any other rotor from the same wafer C when it is broken without discarding the whole device. The assembled five-wafer device has dimensions of 15 mm 15 mm 2.3 mm. It has a multistage turbine, a hydrostatic TB, a hydrostatic JB, seals, and a pump. The main TB is not included in the final device, as explained previously. D. Start-Up Challenges Once aligned well, the blade rows should be concentric, keeping some clearance between the rotor and stator rows. However, in most of the assembly trials, they were found to touch or be stuck with each other. It was hard to observe this misalignment with IR imaging prior to packaging, but it may happen once flow is applied or during the final packaging steps. Unfortunately, many of the blades were broken during start-up, as shown in Fig. 10. Most of the broken blades were from the stator, while the rotor blades remained in good condition, as shown in Fig. 10(a). This illustrates the fragility of the SOI interface. It was identified that the gap between blade rows (25 μm) was less than the total tolerance (35 μm), defined as the sum of misalignments (15 μm) between the A B stack, the C D E stack, and the JB clearance (20 μm). If this tolerance is too large, it can allow interlocking between stator and rotor blade Fig. 11. Thick oxide deposition process for tolerance reduction. (a) Original condition. (b) Thick oxide deposition. (c) RIE. (d) BOE. rows. This tolerance was reduced by depositing thick oxide on the side wall of the alignment holes and JB wall, as shown in Fig. 11. Oxide (12 μm) was deposited on the front side of the C D E stack by PECVD. The chip was then exposed to reactive ion etching (CHF 3 + O 2 gases), which is anisotropic, to remove 10 μm of oxide from the horizontal surfaces. Two micrometers were left unetched to prevent damaging the silicon substrate because the RIE recipe has low etch selectivity (SiO 2 : Si = 5:1). Then, the remaining 2 μm of oxide was removed using isotropic wet etching (BOE 6:1). A final oxide layer of 10 μm therefore remains on the vertical walls to reduce the tolerances. After the process, the total tolerance was reduced from 35 to 15 μm (alignment tolerance between the two stacks: 5 μm, JB tolerance: 10 μm), and the interlocking issue was resolved. E. Modification to the Original Design In order to accommodate the limitations in fabrication techniques and ensure proper rotor operation, the size and geometry of some features changed from the original design, mainly the axial TB gap, the JB gap, the blade height, the main TB orifices, and the auxiliary TB orifice length. These modifications during fabrication have been described mostly in the previous Section II-B and E and are summarized in Table II. The axial operating range of the rotor changed from 1 to 2 μm (referto Part I and wafer D in Fig. 2) to facilitate the device testing by making it less sensitive to axial disturbances as well as to allow the rotors to be interchanged between different static structures (total thickness variation across the wafer measured to be 1 μm). The JB was fabricated similarly to the original design. However, later, the gap was modified by depositing a

7 332 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 TABLE II MODIFICATIONS TO ORIGINAL DESIGN OF COMPONENTS DURING FABRICATION Fig. 12. Schematic of microturbopump test setup. (TB) Auxiliary TB supply. (TB supp) TB support apply extra thrust force on the rotor. (Turb in) Turbine inlet supply. (JPP1) JB pressurization plenum supply 1. (JPP2) JB pressurization plenum supply 2. (Turb out) Turbine outlet exhaust. thick oxide (10 μm) on the wall, as described in Section II-E. This modification also increased the aspect ratio of JB length to gap from 20 to 40. A theoretical study [11] suggests that the smaller gap would give higher JB stiffness for the rotor operation, which leads to stronger restoring force in case of imbalance. The rotor blade height has been increased by 40% from the original design. This change will cause higher loss due to tip leakage; hence, further improvement is necessary in the design and fabrication of the rotor. The orifices of the main TB could not be etched due to extremely low etch rate in high aspect features. A new design was made in order to avoid this problem, but not implemented in this fabrication process [10]. Fortunately, an auxiliary TB was designed along with the main TB and is used as a replacement to the main TB for the rotor operation. The orifice length of the auxiliary TB has been changed from 150 to 85 μm as explained. The short orifice depth will decrease the flow resistance and decrease the stiffness of the auxiliary TB. III. DEMONSTRATION OF THE MICROTURBOPUMP The test procedure starts by characterizing the TBs followed by the JB, turbine, pump, and the whole system. Throughout, the test results are compared with the models of components developed in Part I to validate them and establish a design basis for this type of device. This modeling was performed based on the fabricated dimensions of each component. A. Experimental Setup The test setup shown in Fig. 12 is composed of gas- and liquid-flow systems, a speed measurement sensor, and a digital data acquisition (DAQ) system. The gas-flow system is used for testing bearing and turbine performance with air and consists of six separate pressure sensors (OMEGA PX4202) and five massflow controllers (MKS 1179A). The liquid handling system consists of a filter, a syringe, two pressure sensors, and a mass flowmeter. Filtered deionized water is used to keep the device clean. In order to trap bubbles formed inside the tube or connections, a syringe was installed upstream of the package. Flow pressures at the pump inlet and outlet are measured right upstream and downstream of the package using liquid pressure sensors (OMEGA PX209). The metering valve at the outlet regulates the pump outlet pressure for pump performance characterization. A liquid flowmeter with high resolution (0.3 ml/min, SENSIRION ASL ) was installed right downstream of the metering valve for realtime measurement of pump flow rate. The rotation rate is measured from a fiber-optic speed sensor (Philtec Model 6D). While the rotor spins, the optical probe senses the displacement signal from the speed bumps on the rotor, which generates square waves. The raw data are transferred to the PC-based DAQ system and converted to frequency signal through a fast Fourier transform to get the rotation rate [12]. Signals from the sensors, which are of analog form, are collected by a PC-based DAQ system through an analog input card (Keithley KPCI 3110). The DAQ system is implemented using Labview software, and the real-time data are displayed on the computer terminal and recorded in the computer. The flows can either be controlled with metering valves or mass-flow controllers through an analog output card (Keithley KPCI 3130). A test package for the microturbopump was used to connect the ports on the device to tubing leading to the sensors and fluid sources. It consists of three layers of Plexiglass and aluminum

8 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 333 plates [10], [12]. The first and third layers are composed of piping for working fluids such as air and water, to connect the ports [refer to Fig. 9(a)] on the chip to sensors, and to measure pressure and flow rate. The second layer made of an aluminum plate holds the device inside the square area to align it with the other plates. Between the silicon device and the package, o-rings are inserted for mechanical support and prevent leakage. A hole in the middle of the first plate was made to allow an optical probe to measure dynamic parameters of the rotor. B. Device Operating Procedure The turbopump device was designed to use steam as the working fluid for the turbine, seals, and bearings. However, it can also run on compressed air. In the work presented in this paper, compressed air was used because steam would require a more complicated test apparatus and can condense depending on the operational and thermal conditions, which causes flow-passage blockage, significant drag on the bearings, and rotor imbalance. At this early stage of development, issues related to steam operation were not fully incorporated to reduce development risk, but further studies on steam operation are ongoing. Since the analytical models developed for the design of the current device can estimate air operation, the air tests were used for validation. The device is set in motion by supplying pressurized gas for the bearings and the turbine. At first, pressure is applied from the bottom side through the auxiliary TB ports. A separate high pressure is applied to both of the JPPs, causing gases to flow axially through the JB to support the rotor in the radial direction. Then, pressure is applied to the turbine inlet with the turbine outlet valve closed to uniformly distribute the pressure on the turbine side. The auxiliary TB pressure is adjusted to balance the turbine pressure force and, thus, to float the rotor in the axial midposition. To start spinning the rotor, the turbine outlet valve is opened slowly, allowing flow and a pressure difference between the inlet and outlet of the turbine. If the rotor is stuck or is touching the top or bottom half of the device, the rotor does not spin. By regulating the auxiliary TB pressure or turbine outlet pressure, the rotor was easily released and started spinning. Once the rotor starts spinning, the rotational speed can be increased by inducing a greater pressure difference across the turbine by either opening the turbine outlet valve or supplying a higher pressure to its inlet. During the operation, the auxiliary TB pressure is kept constant to check the bearing gap, which is indirectly measured from the flow rate for a constant TB pressure. C. Hydrostatic TB Characterization Unlike other silicon MEMS turbomachines demonstrated to date, the axial load in the current microturbine is only balanced by one TB on the aft side of the rotor. In the device as fabricated, the auxiliary TB is used, which consists of a circular array of orifices that supply pressurized gas over the back side of the rotor to compensate for the downward pressure force imparted by the turbine flow. Therefore, the technique to support the rotor in the axial direction is different from the devices with Fig. 13. Comparison of data and model at a TB inlet pressure of 172 kpa. The error range of the flow rate is +/ 2 sccm. both front- and aft-side TBs [13]. At first, the auxiliary TB is supplied with high pressure air and the rotor is lifted up until it touches the top half (or ceiling of the rotor chamber). Then, pressure from the turbine inlet is applied to the rotor so that it can be pushed down toward the aft side. The turbine outlet remains closed along with the JB port, leading to a uniform pressure distribution on the rotor. This makes the JPP pressure equal to the turbine exit pressure. The TB flow is discharged to two directions, one inwardly and the other outwardly. The inward exit is open to the ambient, and the other one is connected to the JB port. By using this technique, the auxiliary TB was characterized. To do so, the bearing pressure was kept constant, and the turbine inlet pressure was increased gradually. The test results are shown in Fig. 13. The bearing force is calculated by integrating pressure distribution across the turbine, which is the same as the TB force at the balanced data point. As the turbine pressure increases, the bearing gap decreases and so does the flow rate of the TB. The model could calculate only up to 172 kpa (gauge pressure) due to a premature choking prediction. A comparison between the data and the model (Fig. 13) shows similar behavior, but the slope of the flow rate data is stiffer than that of the model. In the modeling, both the exits were assumed to be open to the ambient. However, the outward exit of the TB is connected to the JB port, which has a pressure near that of the turbine exit. The stiffer slope of the data in Fig. 13 appears due to the increased pressure at the outward exit, which reduces the overall flow rate. Data for various applied TB pressures are shown in Fig. 14. This behavior is well captured by a third-order polynomial fit and shows linear increase of flow rate and force for increasing applied pressure. This chart is useful for dynamic operation of the device by predicting the TB pressure to support the turbineside load. By extrapolating from this chart, pressures as high as 2.4 MPa are expected to be necessary to support the turbine-side load at design conditions (1.6 N) using the auxiliary TB, which is unfortunately not achievable with the current test apparatus and could limit the achievable operating speed.

9 334 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 Fig. 14. TB flow rate for different fixed TB pressure. (Symbols) TB test data and (solid line) cubic polynomial curve fit. D. Hydrostatic JB Characterization Pressure was measured as a function of flow rate supplied through the JB ports both in static and dynamic (rotating) conditions. In the static test, there was no flow or pressure supply to the turbine and TB, and the turbine exit valve was fully open (turbine exit pressure is almost the same as the atmospheric pressure). During the dynamic test, 200 sccm of constant flow rate was supplied to the turbine and 687 kpa of constant pressure to the TB to spin the rotor, while the turbine exit valve is partly open for turbine operation (turbine exit pressure is higher than the atmospheric pressure due to pressure buildup). Fig. 15 shows the differential pressure (dp) across the JB as flow rate increases. The pressure-flow behavior was similar between static and dynamic conditions. The dynamic data were obtained while the rotor was spinning in the range of r/min. The rotor started to spin at as low as 0.69 kpa and operated up to 83 kpa, which corresponded to a flow of 1160 sccm. Above this range, the rotor stopped spinning. This was observed in other devices also, although the range varied. In the dynamic data, a deviation from the static behavior was repeatedly observed near 400 sccm. This appears to be caused by the wobbling of the rotor, which happens when it spins near the natural frequency of the JB system. While the rotor wobbles, it spins off center and the flow resistance across the JB becomes smaller, which leads to lower pressure drop. Based on the rotating speed at 400 sccm, the natural frequency of the JB systems can be estimated to be 9000 r/min. According to a study on the JB [14], the wobbling at natural frequency happens at relatively low speeds (< r/min) and does not cause catastrophic crash of the rotor. Above the critical speed, the rotor operates stably, if the rotor imbalance is not greater than the JB gap. Upon fabrication, the etch precision can lead to the center of mass being slightly offset from the geometric center of the rotor, which creates an imbalance. Previous work has shown that the etching approach can provide rotors with less than 2 μm of offset, which is sufficient for high-speed operation [3], [15], but can lead to vibrations of the rotor. In our experiments, we did not notice excessive vibrations up to Fig. 15. JB test result. The dp is between the pressures supplied to the JPP and turbine outlet exhaust r/min but, instead, experienced sudden failure, as will be discussed next. E. Turbine Operation 1) Operating Procedure: The device is normally operated by maintaining the TB flow rate constant with a fixed supply pressure in order to keep the gap between the rotor and bottom constant. The turbine flow rate is then increased by regulating the dp between the inlet and exit of the turbine using the metering valves upstream and downstream of the turbine. When the downward thrust force from the turbine flow is too high for the TB to support the rotor, additional force is supplied by pressurizing the bottom side of the rotor through the pressure tap (TB supp). 2) Turbine Characterization: Precise control of the axial and lateral balance of the rotor-enabled high-speed operation, as shown in Fig. 16. Data were obtained from a device without broken blades (Device 1) and from a device with the first three rows of stator blades broken (Device 2). The TB flows were kept at 56 and 80 sccm at feed pressures of 823 and 689 kpa for devices 1 and 2, respectively. Both of the TB flow rates correspond to 90% 92% of the maximum flow rate expected when the TB gap is fully open (61 and 80 sccm for devices 1 and 2, respectively, at the aforementioned pressures), so they are expected to have almost identical axial positions. The JB flow rate varied from 100 to 300 sccm during the tests. At higher speeds, more flow was supplied to the JB to keep the lateral balance. Device 1 operated up to r/min with 1500 sccm of turbine flow rate supplied. The model predicts 0.38 W of mechanical power production and 35% of adiabatic turbine efficiency at this condition [15]. The data also agree well with model predictions, validating the approach. At the maximum speed, the rotor crashed due to sudden pressure drop in the TB, which was caused by an internal fracture of the silicon bearing structures. Careful investigation revealed that the massflow controller of the turbine flow failed at that speed, which caused a high pressure difference across the bearing structure,

10 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 335 Fig. 16. Rotation rate as a function of air flow rate through the turbine. Data are presented for two different devices: Device 1 has all stator blade rows undamaged, and Device 2 has only the fourth stator blade row (other blades were broken due to bad alignment during the device assembly). The modeling is based on the data from the device without broken blades. The error range of flow rate is +/ 200 sccm. leading to its fracture. With device 2, a maximum speed of r/min was achieved at a turbine flow rate of 3300 sccm. It appears that the rotor began rubbing with the top portion of the device at this maximum speed, stopping its rotation. Device 1 ran about two and half times faster than device 2 for the same flow supply. This is because there is no flow turning in the three damaged stators of device 2, thus leading to less turbine power production and slower rotation rates. Further details of turbine characterization are presented elsewhere [15]. F. Pump Performance 1) Operating Procedure: The spiral-groove viscous micropump is operated, first, by achieving stable bearing and turbine operation with air as in the previous sections. Once the rotor operates at a desired speed, plastic tubes are connected to the pump inlet and outlet ports of the package to supply and drain water. Dynamic operation is required for the hydrodynamic seals to meet their function and prevent liquid from reaching the high drag areas under the disk. The speed usually drops by 30% 70% depending on the conditions due to the increase in viscous drag when the pump gets wet. Monitoring the rotor speed confirms that liquid entered the pump. 2) Pump Characterization: For pump characterization, an inward pumping device was used with flow exiting from the pump side of the device. The pump flow rate is regulated from zero to maximum flow with the metering valve. The inward pumping device enabled full pump characterization without leakage. On the contrary, it is difficult to perform this test with the outward pumping device due to the limitation in seal performance, particularly in high-pressure conditions. The pump performance chart (Fig. 17) was completely characterized for speeds up to r/min. The maximum achieved flow rate and maximum pressure rise in this test are 9 mg/s and 240 kpa, respectively. The pump model based on lubrication theory is in close agreement with the data. The model predicts 7.2% of maximum efficiency over the range Fig. 17. Spiral-groove viscous micropump performance chart. Shaded areas represent uncertainty of the gap between the rotor and the pump in modeling, which is 1 ± 0.25 μm. The modeling does not include the centrifugal effect due to negligible contribution. Contours of pump efficiency (in percent) are also shown, emanating from the origin. This device has inward-pumping spiral grooves. The uncertainty of pressure is ±7 kpa, and that of flow rate is ±3% of each measurement. of operating speeds, which is at least one order of magnitude superior to other MEMS pumps, with efficiencies typically less than 0.5% [16]. Here, the pump efficiency is defined as the ratio of pumping work (which is the multiplication of pressure rise and volume flow rate) to the shaft mechanical power required to drive the pump. Because of the moderate speed operation, the centrifugal effect could not be identified, and the modeling analysis did not show a clear difference between inward and outward pumping at this speed range. High-speed operation is necessary to see the effect, as mentioned in Part I. In the performance chart of conventional centrifugal pumps of high Reynolds number flows (Re > 10 4 ), the performance curves are rounded at the ends due to various losses such as flow recirculation, flow blockage, and cavitation. However, a micro viscous pump has linear performance curves matching well with the model prediction, in which we did not consider such losses. This suggests that the viscous pump is negligibly affected by the losses that large-scale pumps suffer since the viscous drag dominates over other loss mechanisms not considered here. Typically, the head rise of centrifugal pumps is proportional to the square of rotational speed. In contrast, that of the viscous pump is linearly proportional, as can be seen in Fig. 17 and (4) of Part I. This implies that the efficiency can remain constant, even at maximum efficiency, for varying operational speeds, as shown in Fig. 19. The results of measurement and model prediction of the pump performance reiterate the adequacy of using the viscous pump for a microscale Rankinecycle power-generation system, i.e., it can deliver the desired mass flow with enough pressure rise at a reasonable efficiency. G. Microturbopump System Performance The overall turbopump performance was characterized by keeping the pump outlet valve opening fixed and measuring the inlet and outlet conditions of the turbine and pump (Fig. 18).

11 336 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 Fig. 18. Microturbopump speed as a function of turbine pressure difference while operating with liquid in the pump. The uncertainty of pressure is ±7 kpa. Fig. 20. Relative power consumption of each component predicted by the model for the geometry as fabricated (turbine: generation, other bars: consumption). The TB was intentionally overdesigned for high drag in order to mimic the load of a generator that would be integrated on the rotor for a Rankine Microturbine power plant on a chip. The seal drag is higher than the design value since it is assumed here to be wet. After the experiment, the device was opened to check if the seal was wet during the operation, but it was not clear from the inspection by an optical microscope. The assumption of a liquid in the seal increases the drag (higher viscosity than gas) and leads to a conservative estimate of mechanical efficiency. Under this assumption, the seal dissipates most of the turbinegenerated power and the mechanical-drive efficiency (40%) and fluid-to-electricity-projected efficiency (2.5%) are lower than the design values. These results suggest that keeping the seal dry in the device operation is crucial for high-power-conversion efficiency. Fig. 19. Measured and predicted pumping performance (power and efficiency) of the microturbopump. In this test, the rotor was spun up to r/min, which corresponds to 25 m/s in tip speed, with 41 kpa of turbine dp and 24 mg/s of turbine flow rate, while the pump pressurized water by 88 kpa with a flow rate of 4 mg/s. At the maximum speed achieved, the turbine model predicts W of mechanical power production and an isentropic efficiency of 16% with Re = 186. The pump model suggests that the pump efficiency remains constant near its maximum (7.2%) over the operating range (Fig. 19) unlike big-scale machines, which run at lower efficiencies at low speeds due to higher loss [17]. Based on the system model calculation, out of the total power produced by the turbine, 10% was consumed by the viscous pump, while the rest was dissipated by other components through viscous drag (Fig. 20). The breakdown of power consumption in the microturbopump is shown in Fig. 20, based on the model calculation at the maximum operating speed (also similar for other speeds). The turbine bar indicates power production, while the others represent consumption or dissipation. IV. PROJECTED PERFORMANCE OF A MICRO-RANKINE POWER GENERATOR The measured performance, particularly the rotor speed, is still far lower than designed. It is due to the sudden fracture of the TB, as mentioned earlier. During the operation, there was no serious evidence of rotor instability such as wobbling up to the failure speed, and it will be possible to reach higher speeds when bearing (particularly TB) and turbine flows are carefully controlled. However, these and other results [15] validate our component and system models. Based on the model predictions at the design speed, the turbine would have 70% of isentropic efficiency and the pump would have 7%. This estimation of the efficiencies is close to the assumptions of Mueller and Fréchette [18], which suggest that a whole Rankine Microturbine power plant on a chip could produce several watts of electric power with 1% 12% of thermal efficiency, depending on the heat addition and discharge conditions, for a similar-size device as the current one. As expected, the pump efficiency is relatively low compared with the MIT microcompressor (50% 70% in optimum conditions [3]), which is part of a Brayton cycle micro-gas-turbine engine. However, the power consumed by the pump is almost

12 LEE et al.: SILICON MICROTURBOPUMP FOR RANKINE-CYCLE POWER-GENERATION MICROSYSTEM II 337 trivial compared with the turbine power available. In contrast, the microcompressor inherently consumes a large portion of the turbine-generated power, which results in poor thermal efficiency for a micro gas turbine, predicted to be less than 3% [18]. V. C ONCLUSION A microturbopump device with a multistage turbine and viscous pump supported on gas-lubricated bearings has been fabricated and demonstrated. The device was made of five bonded wafers, which went through 14 photolithography and etching processes. The 4-mm-diameter rotor was spun up to r/min with compressed air only (dry pump) and up to r/min while pumping water at the rate of 4 mg/s with 88 kpa of pressurization. At the maximum speed reached while pumping, the turbine produced W of mechanical power, of which 10% was consumed by the viscous pump. The pump performance chart was also completely characterized for speeds up to r/min and suggests 7.2% of maximum efficiency over the range of operating speeds. The rotational speed was limited by axial balance limitations, which will be addressed in future work to increase the operating range. Based on the test results and extrapolations to higher speeds using our model predictions, the approach studied here is expected to lead to reasonable thermal efficiency and very high power density for a complete Rankine Microturbine power plant on a chip. The use of a viscous spiral-groove micropump combined with a microturbine in a planar MEMS configuration has been shown to be adequate and promises to meet the pressure and flow requirements for a Rankine power cycle. However, the choice of silicon as a structural material would lead to excessive heat loss through the device, so alternative insulating materials must be considered to achieve the device s potential. This paper therefore proves the concept of the rotating subsystem for a Rankine Microturbine and provides a validated design basis for future development. The technology demonstrated herein will also contribute to the development of other types of power MEMS, such as micropumps for chip cooling or high-pressure fuel delivery for a microrocket engine. High-speed microfabricated silicon turbomachinery and fluid film bearings, J. Microelectromech. Syst., vol. 14, no. 1, pp , Feb [4] P. Kang, S. Tanaka, and M. Esashi, Demonstration of a MEMS-based turbocharger on a sing rotor, J. Micromech. Microeng., vol. 15, no. 5, pp , May [5] N. Savoulides, S. A. Jacobson, H. Li, L. Ho, R. Khanna, C.-J. Teo, J. M. Protz, L. Wang, D. Ward, M. A. Schmidt, and A. H. Epstein, Fabrication and testing of a high-speed microscale turbocharger, J. Microelectromech. Syst., vol. 17, no. 5, pp , Oct [6] A. Deux, Design of a silicon microfabricated rocket engine turbopump, M.S. thesis, MIT, Cambridge, MA, [7] L. Jamonet, Testing of a microrocket engine turbopump, M.S. thesis, MIT, Cambridge, MA, [8] S. Diez, Preliminary performance characteristics of a microfabricated turbopump, M.S. thesis, MIT, Cambridge, MA, [9] C. Lee, M. Liamini, and L. G. Fréchette, Design, fabrication, and characterization of a microturbopump for a Rankine cycle micro power generator, in Proc. Solid State Sens., Actuator, Microsyst. Workshop, Hilton Head Island, SC, Jun. 4 8, 2006, pp [10] C. Lee, Development of a microfabricated turbopump for a Rankine vapor power cycle, Ph.D. dissertation, Columbia Univ., New York, NY, Feb [11] Z. S. Spakovszky and L. X. Liu, Scaling laws for ultra-short hydrostatic gas journal bearings, J. Vib. Acoust., vol. 127, no. 3, pp , Jun [12] L. G. Fréchette, Development of a microfabricated silicon motordriven compression system, Ph.D. dissertation, MIT, Cambridge, MA, Sep [13] C. C. Lin, Development of a microfabricated turbine-drive air bearing rig, Ph.D. dissertation, MIT, Cambridge, MA, Jun [14] L. Liu, Theory for hydrostatic gas journal bearings for microelectro-mechanical systems, Ph.D. dissertation, MIT, Cambridge, MA, Sep [15] C. Lee, S. Arslan, and L. G. Fréchette, Design principles and measured performance of multi-stage radial flow microturbomachinery at low Reynolds numbers, J. Fluids Eng., vol. 130, no. 11, p , [16] D. J. Laser and J. G. Santiago, A review of micropumps, J. Micromech. Microeng., vol. 14, no. 6, pp. R35 R64, Jun [17] E. Logan, Turbomachinery Basic Theory and Applications, 2nd ed. New York: Marcel Dekker, [18] N. Mueller and L. G. Fréchette, Performance analysis of Brayton and Rankine cycle microsystems for portable power generation, in Proc. ASME Int. Mech. Eng. Congr. Expo., New Orleans, LA, Nov , 2002, pp ACKNOWLEDGMENT The fabrication work was performed in part at the Cornell NanoScale Science and Technology Facility (CNF), a member of the National Nanotechnology Infrastructure Network, which is supported by the National Science Foundation under Grant ECS The authors would like to thank F. Gauthier for providing the 3-D device drawings. REFERENCES [1] C. Lee and L. G. Fréchette, A silicon microturbopump for a Rankinecycle power generation microsystem Part I: Component and system design, J. Microelectromech. Syst., vol.20, no.1, pp ,Feb [2] K. E. Peterson, Silicon as a mechanical material, Proc. IEEE, vol. 70, no. 5, pp , May [3] L. G. Fréchette, S. A. Jacobson, K. S. Breuer, F. F. Ehrich, R. Ghodssi, R. Khanna, C. W. Wong, X. Zhang, M. A. Schmidt, and A. H. Epstein, Changgu Lee received the B.S. and M.S. degrees in mechanical engineering from Hanyang University, Seoul, Korea, in 1995 and 1997, respectively, and the Ph.D. degree in mechanical engineering from Columbia University, New York, NY, in He was a Postdoctoral Researcher with Columbia University between 2006 and 2010 in the Department of Mechanical Engineering under the supervision of Prof. James Hone. Since September 2010, he has been an Assistant Professor with the Department of Mechanical Engineering, Sungkyunkwan University, Suwon, Korea. His current research interest is in characterization of mechanical and physical properties of 2-D nanomaterials such as graphene and hexagonal boron nitride and their applications. He has measured mechanical properties of graphene and characterized tribological properties of 2-D materials and investigated optical and electronic properties of hexagonal boron nitride and molybdenum disulfide.

13 338 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 20, NO. 1, FEBRUARY 2011 Mokhtar Liamini received the B.S. degree in mechanical engineering from the National Polytechnic School, Algiers, Algeria, in 2002, and the M.S. degree in mechanical engineering with specialization in thermofluidics from the Université de Sherbrooke, Sherbrooke, QC, Canada, in 2005, where he is currently working toward the Ph.D. degree and where he designs and fabricates Rankine microturbopumps operating at high temperature with a special focus on the thermal insulation of the devices for operation at high temperature. Luc G. Fréchette (M 04) received the B.Ing. degree from the École Polytechnique de Montréal, Montreal, QC, Canada, in 1994, and the S.M. and Ph.D. degrees from the Massachusetts Institute of Technology, Cambridge, in 1997 and 2000, respectively. He is currently the Canada Research Chair in Microfluidics and Power MEMS and a Professor of Mechanical Engineering at the Université de Sherbrooke, Sherbrooke, QC. From 2000 to 2004, he was a Faculty Member with Columbia University, New York, NY. His expertise is in microengineering of miniature systems for energy conversion, such as heat engines (microturbines), fuel cells, cooling microsystems, and micro energy-harvesting devices. His activities range from integrated device development to more fundamental fluidic, heat, and mass transfer aspects at small scale. He also enjoys developing MEMS sensors and actuators for aerospace and other harsh environments. Dr. Fréchette is a member of the American Society of Mechanical Engineers.