Leonie O Vogt

Cyril Marino

Claire Villevieille
email: claire.villevieille@psi.ch

Electrode engineering of conversion based negative electrodes for Na-ion batteries

Keywords: Na-ion batteries, negative electrode, Sn particle, electrodes engineering

Due to lower costs and higher abundance of sodium, Na-ion battery technology can offer a good alternative to Li-ion batteries. Many researches focus on developing new cathode and anode materials but the importance of the electrode engineering on the electrochemical performance is often neglected. The electrode composition is especially crucial for conversion reaction based-materials where the composite electrode (active material, conducting additive and binder) has to buffer the huge volume change occurring upon cycling. This work highlights the differences observed on Sn-CMC electrode performance by using different Sn particle size (micro and nano particles) and evaluating the role of the conductive additive in the electrode. Carbon fibers (VGCF) demonstrate a good ability to surround micrometer particles but not especially nanometer particles leading to improve the performance of microparticles but not of nanoparticles. For a high loading electrode suitable for full cell applications (> 3.5 mg/cm 2 of active material), nanometer particles show limited performance for long term cycling. The combination of VGCF with micrometer particles seems to be the most promising composition to obtain good performances for conversion reaction based-materials.

Li-ion batteries are starting to reach their limits in terms of energy density, cost, and abundance, and progress is slower than was expected. Thus, researchers are currently re-investigating other alkali metals as Li substitutes, mainly focusing on Na. This system has been considered to be purely academic, and no real applications or prototypes have been developed to investigate its viability or possible commercialization, the only exception being the high temperature Na-S system , which was commercialized in the 1960s 1 . Recently however, the amount of research and number of papers devoted to the development of active materials for Na-ion batteries has increased exponentially, leading the community to consider the commercialization of Na-ion batteries in the near future [2][3][4][5][6] . As known from the electrochemical facts, specifically voltage and theoretical specific charge, there is no advantage to replacing Li with Na. However, when considering additional parameters such as abundance and cost Na-ion batteries may be competitive with the currently available Li-ion batteries.

Another significant advantage in term of both cost and weight is that no alloy exists between sodium and aluminum. This means that Al can be used as a cheaper current collector for both the positive and negative electrodes, lowering the total weight and cost of the cell pack. Furthermore, the electrolyte in Na-ion batteries has a higher conductivity compared to that of Li-ion batteries. The larger ionic radius of Na compared with Li results in a low de-solvation energy, which strongly influences the kinetics and allows high power Na-ion batteries to be envisaged. To achieve commercialization, suitable anodes 2, 4, 7-11 and cathodes [12][13][14][15][16][17][18][19][20][21] must be developed and studied in depth as the Na-system is often not analogous to the Li-system. A couple of years ago, the observation that pure commercial elements, such as Sn or Sb can electrochemically react with Na, leading to sustainable reversible capacities as high as 500 mAh/g over more than 100 cycles (2 times more than hard carbon), surprised the scientific community 7,22 . These results were especially interesting considering the immense volume change upon cycling, which exceeds 400%. However, while the search for new materials is progressing rapidly, the engineering of the electrodes is quite often neglected. Many groups in the literature reported as an example that PVDF binder (polyvinylidene fluoride) is working poorly in Na-ion batteries and that other type of binders should be used instead [23][24][25] . We were one of the first groups to demonstrate that the known-how acquired over the past few decades on Li-ion battery technology cannot be transferred one-to-one to the emerging Na-ion batteries 26 . In fact, we demonstrated that the most commonly used binder in Li-ion batteries (PVDF, polyvinylidene fluoride) is not suitable for Na-ion batteries as it decomposes to create NaF during cycling and thus the electrode loses its integrity and cannot accommodate the volume change of the conversion-based electrode materials anymore 26 . Besides binder, carbon-conductive agents (carbon fibers, super P, acetylene black, carbon nanotubes, etc…) are another source of variation from group to group, where many different types are used but it is difficult to crystalize which of them lead to the best cycling performance with conversion-based electrode materials for Na-ion batteries [27][28][29] .

The influence of the particles size (of the active material) is another often neglected factor in literature, since articles either focus on nano-particles or they focus on solid-state bulk particles (tens of μm range). The comparison between both particle sizes (nano and micro range) is not often reported and it is difficult to assess the advantages and drawbacks of each since the loading of the electrodes (using either nano, or micro range) can differ by a factor of 5 to 10 9, 30-33 . Based on the common approach of using very low loading for nanoparticle active materials (unsuitable for battery application) it is not surprising that the nanoparticles are often presented to be better in term of cycling stability, long term cycling and power performance, as for example reported recently by Nam et al. 34 and others [35][36][37][38][39] .

In this article, we establish an engineering guideline for conversion-based negative electrodes used in Na-ion batteries. Sn was selected as a model conversion based-material which undergoes high volume changes during cycling. Due to the instability of the PVDF binder in Na-ion batteries, shown previously, it was decided to use the promising binder Na-CMC. Here we compare the impact of different particle sizes on the electrochemical performance in electrodes of comparable loading. Additionally, we investigate the impact of the carbon conductive agent in both the micro-particle and the nano-particle system, looking specifically at the use and impact of carbon Super C and carbon fiber. The best combination of binder/conductive additive/particle size for improved cycling performance of conversion based negative electrodes in Na-ion batteries is then elucidated.

Experimental section

Electrode preparation

For the electrode preparation we used two different tin powders: Sn-micro (ABCR, 325 mesh) (hereafter called m-Sn) and Sn-nano (ABCR, ca.100 nm) (hereafter called n-Sn). For the conductive agent, two different types were used: carbon black SuperC65 (CB, Imerys) and vapor-grown carbon fibers (VGCF, Showa Denko). Electrodes without VGCF were prepared by casting a suspension of 70 wt % Sn-micro (or Sn-nano), 18 wt % CB and 12 wt % carboxymethyl cellulose sodium salt (CMC, Alfa Aesar) in deionized water onto Aluminum foil, used as current collector (described in the article as composition 1). For electrodes containing VGCF, the ratio between the elements was modified to 70 wt % Sn-micro (or Sn-nano), 9 wt % CB, 9 wt % VCGF and 12 wt % CMC (described in the article as composition 2). Once dried in the air, 13 mm diameter electrodes were punched out and were dried under vacuum at 120˚C for a few hours. The loading of active material on the electrodes was between 2.8 mg/cm 2 and 3.5 mg/cm 2 .

Electrochemistry

Electrochemical cells were assembled in an Ar-filled glove box using the electrode, a glass fibers sheet as the separator and metallic sodium (Sigma-Aldrich, 75 μm) as counter electrode. A mixture of 1 M NaClO4 dissolved in propylene carbonate (PC) was used as electrolyte. The cell performance was measured in galvanostatic mode at 25˚C between 5 mV and 1.0 V at a C/30 rate and monitored by an ASTROL cycling device. An additional potentiostatic step of 2h was added at the end of each discharge and charge to ensure proper sodiation and desodiation. All the potentials mentioned in the manuscript are given versus Na + /Na.

Scanning electron microscope (SEM)

SEM measurements were performed in a Carl Zeiss Ultra55 scanning electron microscope using the secondary electron mode.

Results and Discussion

Influence of Particle Size

The SEM pictures of m-Sn and n-Sn electrodes are displayed in Figure 1 and show the impact of the particles size on the electrode engineering. The m-Sn electrode reveals a broad range of particle sizes from ca. 1 µm to 15 µm. Some Sn particles are close to each other and in other areas there are big gaps, indicating a certain inhomogeneity at the micrometer range in the dispersion of the particles in the electrode. The images also show some Sn particles that are only partially covered by the carbon super C and the CMC binder and some cracks around the particles. Those fractures occurring in the electrode can be problematic considering the huge volume change occurring during the sodiation of Sn particles (especially for the m-Sn electrode). Particles that are only partially connected to the carbon/binder/current collector are likely to be totally disconnected during the cycling due to the strong strain caused by the volume change upon sodiation. This phenomenon would lead to a decrease of the cell performance. The n-Sn electrode shows a narrow particle size distribution since the Sn particles are below 500 nm, with most of them around 100 nm diameter. At low magnification (micrometer range SEM picture, Figure 1 bottom), the Sn nanoparticles seem to be well dispersed and embedded in the carbon/binder matrix unlike the m-Sn electrode. Thus, upon cycling versus Na the nano-particle system is expected to accommodate the volume expansion better than the microparticle system. However, the picture taken at higher magnification (Figure 1, bottom) reveals some n-Sn agglomerates, indicating that the electrode engineering could be optimized further. This agglomeration of nanoparticles may be a result of the high nanoparticle loading used here, as compared to the low loading generally reported in literature. In both cases, the images also indicate a good porosity such that the wettability of the electrode should not be a cause of fading. with a value close to 800 mAh/g and drops down to reach only 300 mAh/g after 20 cycles. Additionally, it is noteworthy that the difference between the charge and the discharge is very low, indicating a good Coulombic efficiency. For the n-Sn, the results are different with a similar specific charge for the first cycle at around 750 mAh/g, and stabilization during 6 cycles at a specific charge of 650 mAh/g. After 6 cycles, the specific charge collapses dramatically to reach only 50 mAh/g after 20 cycles. The Coulombic efficiency in the case of n-Sn is relatively low compared to the m-Sn.

The difference observed between the two electrodes m-Sn and n-Sn can be explained by the difference in particles size. The nanoparticles have such high surface area that when cycling in a potential window where the electrolyte is reduced, more electrolyte decomposition occurs and consequently more solid electrolyte interphase (SEI) forms, leading to a low Coulombic efficiency. With every cycle the SEI layer is thickened, its isolating character leads to an increased resistance which finally hinders cyclability.

For the m-Sn the fading is mainly caused by the particle fracture which occurs during sodiation. In fact when fully sodiated the particles expand 420%, leading to cracking which can cause loss of electrical contact of some particles. Those parts become dead materials that cannot cycle anymore, leading to a gradual fading of specific charge. These results are in accordance with the SEM observations and predictions, in that m-Sn was seen to integrate less well in the carbon/binder matrix than n-Sn (Figure 1) and consequently it is more likely to suffer from pulverization and electrical disconnection of particles, leading to gradual specific charge fading as seen in Figure 2. However compared to the n-Sn electrode, the m-Sn electrode has a better specific charge after 20 cycles, coming to the fact that during cycling most probably after each fracture of the particle, a core-shell process is occurring and thus "new-fresh" Sn can be revealed and can be cycled, which helps to maintain the specific charge higher than the isolating n-Sn. For the rest of the first sodiation, we can see the three characteristic potential plateaus of Na-Sn alloys at 0.2 V, 0.07 V and 0.01 V. For the n-Sn electrode, those potential plateaus are slightly lower in potential, indicating a higher polarization. This higher polarization probably arises from the larger SEI formation which hinders the lowest sodiation plateau to be accessed since with the additional overpotential it lies below the 5 mV cut-off voltage. On desodiation, a first potential plateau appears in the m-Sn electrode at 0.16 V but no such potential plateau is seen for the n-Sn electrode. Thereafter the desodiation plateaus of the n-Sn electrode (0.27 V and 0. 54 V) are analogous to those of the m-Sn electrode (0.26 V and 0.53 V), with the slight shift in potential being attributable to the higher polarization in the n-Sn system (thicker SEI).

At the 10 th cycle (Figure 3, right), the polarization has grown in both electrodes such that the lowest potential plateau is no longer accessible for both systems. In general the different potential plateaus on sodiation are no longer distinguishable for m-Sn electrode and for the n-Sn electrode only two vague potential plateaus are visible at ca. 0.3 V and ca. 40 mV. In contrast, on desodiation potential plateaus are still distinguishable for both m-Sn and n-Sn allowing some comparison to the first cycle.

For m-Sn the two potential plateaus at 0.27 V and 0.54 V are the most prominent features but have shortened in comparison to the 1 st cycle. For n-Sn a new, small potential plateau is seen at 0.22 V followed by one at 0.28 V. Then the potential increases rapidly with a small bend at 0.55 V and another at 0.81 V. So far we demonstrated that when the loading of the electrodes is the same, nanoparticles do not outperform micro-particles after few cycles. Instead, n-Sn shows a poor Coulombic efficiency and a dramatic specific charge fading. Nanoparticles can only compete with micro-particles if the loading is low. However, due to the limited storage capacity such low loading is not commercially viable. A similar comparative approach was used to investigate the most suitable conductive agent and optimal electrode engineering conditions for the best electrochemical performances.

Influence of conductive additive

SEM pictures of the electrode designed with VGCF are shown Figure 4 and can be compared with the one having only super C as a conductive additive (Figure 1). The VGCF, long rod shape around 100 nm diameters, can easily be distinguished in composition 2. For both particle sizes, the observations made previously about the dispersion of Sn particles in Super C-based electrodes are also valid for VGCFbased electrodes: the micrometer particles are not homogeneously distributed in the electrode and the nanometer particles agglomerate slightly. The new electrode constituent, VGCF, shows a homogeneous dispersion throughout the electrode. For m-Sn VGCF electrodes, the Sn particles are covered by the VGCF (Figure 4), such that electrode fractures around the micro Sn particles are bridged by the fibers. This is expected to lead to better electrical contact in the electrode during cycling. For the n-Sn, the situation is slightly different since the diameter of the VGCF is roughly the same as the one of the primary particles. The interaction between the n-Sn particles and the VGCF is therefore limited and VGCF mainly serves to connect aggregates. Looking now at composition 2 (with VGCF, Figure 5), we can see that the trend between m-Sn and n-Sn is roughly the same as the one observed for composition 1. While the composition 2 does not seem to affect cycling for n-Sn, for m-Sn it increases and helps to maintain the specific charge. As already stated bebore for m-Sn, when particle fracture occurs upon sodiation, this can lead to some particles becoming disconnected leading to uncyclable material. Due to its length and high conductivity carbon fibers can combat this problem, allowing more Sn particles to stay connected to the carbon/binder matrix and current collector which leads to a higher specific charge retention (ca. 400 mAh/g after 20 cycles). In n-Sn electrodes where the particles are well integrated and disconnection is not a problem but instead a thick SEI layer causes trouble, the addition of VGCF does not improve cycling performance.

This demonstrates that it is crucial to understand the cause of capacity fading and use the appropriate electrode engineering to combat the drawbacks of the active material. The VGCF improves the long term cycling performance of m-Sn electrode due to better electrical contact. However, there is a small cost in Coulombic efficiency due to SEI formation on the surface of the carbon fibers. For n-Sn, where thick SEI formation leads to dramatic specific charge fading, VGCF cannot help but instead, a coating of the Sn particles may help to reduce electrolyte decomposition to improve long-term performance.

The galvanostatic curves of m-Sn for the 1 st and 10 th cycle in composition 1 and composition 2 are shown in Figure 6. In the first cycle, a new feature at ca. 0.42V for the composition 2 electrode can be seen. This extra "bump" is most probably due to the SEI formation at the surface of the VGCF leading to a slightly reduced Coulombic efficiency as previously mentioned. We notice a similar plateau in the n-Sn galvanostatic curve for the composition 1. The addition of VGCF to the electrode composition reduces the overpotential compared to the standard Super C electrode, which is seen by the different Na-Sn sodiation potential plateaus all lying at slightly higher potential. This means that the lowest of the three characteristic tin sodiation potential plateaus at 0.01 V is fully accessible which leads to a higher specific charge for composition 2. On desodiation the curves obtained with composition 2 are analogous to the ones of composition 1 except that the potential plateau at ca. 0.16 V which corresponds to the lowest potential plateau in sodiation is longer.

At the 10 th cycle, features observed on sodiation have merged into sloping curves for electrode composition 1 and electrode composition 2. As already observed with composition 1, on desodiation the different potential plateaus are still distinguishable. With composition 1, the lowest potential plateau at 0.16 V is no longer visible, however, with composition 2 an activity is still present at this potential. At ca. 0.28 V both compositions show a potential plateau however with composition 2, it is much longer. At higher potentials, the last potential plateau at 0.55 V matches for both compositions. These differences in "length" (i.e. specific charge) indicate that VGCF is of great help to keep the cohesion of m-Sn electrodes during the high volume changes occurring while cycling. The same analysis was conducted for the n-Sn electrodes, and the galvanostatic curves are plotted and presented in Figure 7. For the first sodiation the curves are matching almost perfectly between the two compositions. Only a tiny difference at ca. 0.45 V and at 10mV where the last potential plateau is slightly shorter can be seen. On desodiation we notice that the capacity retention is better for composition 1 than for composition 2 (though only 20-30 mAh/g). After 10 cycles, the specific charge fades dramatically compared to the first cycles. The curves turn smoother compared to the first cycle which makes it difficult to see the different potential plateaus characteristic of the Na-Sn reaction. In both sodiation and desodiation we can see an increase of the polarization for the electrode with composition 2, which can be attributed to a thicker SEI resulting in higher resistance. A thicker SEI can be attributed to the nanoparticles themselves, since their specific area is high and leads to more electrolyte decomposition and when VGCF is present additional SEI formation occurs on its surface.

On sodiation, we can notice a small feature at around 0.4 V attributed most probably to the SEI, and then a potential plateau at around 50 mV attributed to the Na-Sn reaction. On desodiation, the different potential plateaus are more visible and look the same for the two systems. Once again the specific charge of the electrode cycled with the composition 2 is lower than the one cycled with the composition 1 due to the observed overpotential.

Conclusion

Contrary to the frequent reports of high performance nanoparticle batteries, we find that nanoparticles offer no long-term advantage compared to micro sized particles once the electrode loading is the same. Regardless of any electrode engineering, the drawback of the high surface area of nanoparticles and the consequent electrolyte decomposition on this surface cannot be mitigated. For electrodes containing the m-Sn, adding carbon fibers leads to stabilization of the specific charge compared to electrodes not containing such fibers. This improvement can be traced to the conductive fiber maintaining electrical contact between the active material and the conductive carbon/binder matrix during the volume changes occurring upon cycling. A composition of 70% active material, 12% carboxymethyl cellulose binder, 9% Super C65 and 9% VGCF shows the best electrochemical performance. We thus stress the need for careful electrode engineering already at the stage of fundamental research, to be able to make conclusive statements about how materials and their performances compare.

Figure 1 .

 1 Figure 1. SEM images of m-Sn-micro (top) and n-Sn electrodes (bottom) in electrode

Figure 2

 2 Figure 2 depicts the cycling performance of m-Sn and n-Sn in electrode composition 1. The specific

Figure 2 .

 2 Figure 2. Cycling performance obtained for m-Sn and n-Sn in composition 1.

Figure 3 .

 3 Figure 3. Galvanostatic curves for the 1st (left) and 10th (right) cycle for the m-Sn and n-Sn

Figure 4 .Figure 5 .

 45 Figure 4. SEM images of electrode containing VGCF as conductive agent (composition 2) of

Figure 6 .

 6 Figure 6. Galvanostatic curves for the 1 st (left) and 10 th (right) cycle for m-Sn electrodes using

Figure 7 .

 7 Figure 7. Galvanostatic curves for the 1st (left) and 10th (right) cycle for n-Sn electrodes using either the composition 1 or the composition 2.

Acknowledgement

The Swiss National Science Foundation is thanked for financial support (Project 200021_156597). This work was performed within the Swiss Competence Center of Energy Research Heat and Storage (SCCER) framework.


Michael Sheehy

Michael Reid

Jeff Punch

Suresh Goyal

Gerard Kelly

The Failure Mechanisms of Micro-scale Cantilevers in Shock And Vibration Stimuli

HAL is

INTRODUCTION

The reliability of microelectromechanical systems (MEMS) has become an increasing concern with the current development and applications of micro-scale devices in fields such as military and aerospace. Devices which were once considered indestructible to a transient mechanical event are now being employed in harsh environments which generate stresses that can induce failure [START_REF] Srikar | The Reliability of Microelectromechanical Systems (MEMS) in Shock Environments[END_REF].

Key developments in MEMS fatigue and fracture research have been carried out through methods varying from resonance excitation, static deflection loading and design specific on-chip approaches to implement techniques such as electrostatic actuation [START_REF] Ozdoganlar | Experimental modal analysis for microelectromechanical systems[END_REF][START_REF] Epp | A Base Excitation Test Facility for Dynamic Testing of Microsystems[END_REF][START_REF] Pierron | Notch Root Oxide Formation During Fatigue of Polycrystalline Silicon Structural Films[END_REF][START_REF] Muhlstein | High-cycle fatigue of single-crystal silicon thin films[END_REF][START_REF] Liu | Low cycle fatigue of single crystal silicon thin films[END_REF]. Test procedures and metrology techniques have been simultaneously developed and these approaches have contributed to the characterisation of material and geometry properties of micro-devices. When establishing a new test method, silicon is the fabrication material of choice due to its inherent use in MEMS fabrication and also due to the brittle nature of single crystal silicon (SCS) and polycrystalline silicon [START_REF] Gad-El-Hak | The MEMS Handbook[END_REF]. Once a test method is developed with established materials and geometries, it is then applied to a broader range of materials for application specific designs. Reference [START_REF] Yi | Measurement of mechanical properties for MEMS materials[END_REF] investigates several approaches in literature with respect to fatigue and fracture testing of silicon devices. Fabrication processes are proven to have an effect on the fracture strength of a device whereby wet etching techniques yield lower silicon strengths than reactive ion etching (RIE) techniques and crystal orientation is also known to influence silicon strength within RIE techniques. Further investigation also shows that different test methods on an identical fabrication process yield variations in Young's modulus and fracture strengths. References [START_REF] Muhlstein | High-cycle fatigue of single-crystal silicon thin films[END_REF] and [START_REF] Liu | Low cycle fatigue of single crystal silicon thin films[END_REF] investigate the effects of fatigue through cyclic loading and show that SCS fails at lower strains than its static failure strain. Fracture sites are observed through scanning electron microscopy (SEM) and confirm that variations in strain amplitude under cyclic loading affect the crack propagation and failure mode as in [START_REF] Liu | Low cycle fatigue of single crystal silicon thin films[END_REF]. Reference [START_REF] Hull | Fractography: Observing, Measuring and Interpreting Fracture Surface Topography[END_REF] conducted a comprehensive review on the mirror, mist and hackle fractography of single crystals at several scales to provide a tool for the study of brittle cleavage fracture in single crystals.

However, even with advances such as these to characterise material response at a micro-scale, the majority of mechanical testing that has been carried out in current literature is carried out in static environments where excitation of a given device is relatively deterministic. This trend has partly been due to the limitations of instrumentation and existing techniques but predominantly to enable predictable boundary conditions in static environments which enable higher accuracy when correlating results theoretically [START_REF] Gad-El-Hak | The MEMS Handbook[END_REF].

With the growing number of commercial applications for MEMS, research at micro-scale has shifted from passive device characterisation to active device reliability, as seen with the development of the Texas Instruments Digital Micromirror Device [START_REF] Douglass | Lifetime Estimates and Unique Failure Mechanisms of the Digital Micromirror Device[END_REF]. The need for reliability characterisation of devices coupled with harsher application environments has led to a requirement for highly transient dynamic testing of micro-scale devices. There are a select number of methods of producing high amplitude repeatable shock pulses that are capable of exciting micro-scale devices to failure; such as drop tables, ballistics, Hopkinson pressure bars (HPBs) and explosives [START_REF] Brown | Harsh Military Environments and Microelectromechanical (MEMS) Devices[END_REF][START_REF] Li | Drop test and analysis on micro-machined structures[END_REF][START_REF] Tanner | MEMS reliability in shock environments[END_REF][START_REF] Lu | Dynamic loading of LIGA structures[END_REF][START_REF] Kelly | Analysis of high acceleration shock pulses part I: Geometry of Incident Mass[END_REF]. Due to the complexity of the test environments for these methods, there are also only a limited range of measurement techniques to realise any quantitative information from a test program. Therefore when creating a reliability test program, it is necessary to develop measurement techniques that are capable of monitoring the in-situ response of a micro-scale device as well as post analyses techniques to ascertain modes of failure. This paper will investigate the effectiveness of high strain rate shock tests on SCS micro-cantilevers as a test method for characterising fatigue and crack growth in MEMS materials. It will also correlate the transition from low to high strain rate testing and serve as a supplement to current design rules in predicting the failure modes of devices employed in shock environments.

FABRICATION PROCESS Bulk SCS micro-cantilevers were fabricated from 100mm p-type <100> SCS wafer of 525µm thickness with a primary flat orientation of [110]. The beams had characteristic lengths of 1-7mm with cross sections 100µm × 50µm, 100µm × 100µm and 100µm × 200µm. Several of the beams also had proofmasses at the beam tips to influence their response under deflection. Finally each beam type was reproduced with notches of 10% and 40% of beam width to induce failure and/or degradation in the material through stress raisers. Fig. 1 shows a front-side image of a SCS micro-cantilever sample within a frame for mounting on the test apparatus. The sample shown has a 5mm, 6mm and 7mm beams with proofmasses at the beam tips.

The fabrication process was carried out in 54 run steps consisting of cleaning, inspection and physical steps. A thin film oxide layer was deposited onto the front-side and backside of the wafer and the front-side mask was used to transfer the beam geometry onto the oxide surface by means of photolithography. The patterned oxide layer was then stripped off using RIE and the remaining photoresist was stripped from the front-side. A nitride layer was then deposited both front and back-side by means of low pressure chemical vapour deposition (LPCVD) to act as a barrier for a following potassium hydroxide (KOH) solution wet etch. The back-side mask was then used to pattern photoresist onto the wafer and the oxide/nitride layers were removed using a RIE etch. The remaining photoresist was stripped and a timed anisotropic KOH etch was used to remove 425µm of the SCS to obtain the 100µm beam depth. The nitride layer was then removed and aluminium and polyimide layers were deposited on the back-side to act as an etch stop and support the micro-cantilevers while under vacuum respectively. Deep RIE (DRIE) was then used to etch out the beam geometries on the front-side of the wafer. All sacrificial layers were then stripped from the wafer before dicing and cleaning to finish the fabrication process. Table 1 summarises the fabrication procedure into the main physical run steps and Fig. 2 shows a wafer cross section during the process steps. From reference [START_REF] Yi | Measurement of mechanical properties for MEMS materials[END_REF] it has been proven that the choice of fabrication process influences the strength of device under test. A WYKO NT3300 was used to measure the DRIE and KOH etch profiles by means of white light interferometry (WLI). Measurement was carried out in Phase Shift Interferometry (PSI) mode yielding up to 1nm resolution accuracy in the z-axis. Fig. 3 shows interferograms of representative etch profiles on the micro-cantilever beam surfaces where the depth in the images is increased in scale to enhance the profile. Fig. 3(a) shows a depth profile of a KOH etched surface of the proofmass of a micro-cantilever beam. Due to the timed etch stop in the KOH etch there will be variations in depth across the (001) plane etch surface. The surface roughness, R a , of the KOH etch is in the region of 20nm. In Fig. 3 The micro-cantilevers were mounted in a breakaway fixture on a modified Hopkinson pressure bar, as in Fig. 4. The HPB was made of an aluminium 6061-T651 incident rod of 19mm diameter and 3.658m length with strain gauges midsection. This rod was struck at one end by an aluminium striker bar of 19mm diameter and 362mm length. A pressurised gas chamber propels the striker bar into one end of the incident bar. At this impact point 2.8mm to 4.4mm thick stationary card was used as a programming material to shape the pulse wave. This generated the desired compression wave that travelled through the incident bar ultimately delivering a shock pulse to the breakaway fixture. The stainless steel breakaway fixture was bonded to the HPB such that it separated from the incident bar after the initial shock to avoid any additional excitation impulses. The opposing end was threaded for attachment of a Dytran 3200B 80,000g accelerometer. The strain gauge signal was amplified through a Vishay 2110 strain gage conditioner and acquired by a FLUKE PM3394A oscilloscope at 2.

RESULTS AND DISCUSSION

Acceleration levels between 10,000g and 40,000g were required to induce failure mechanisms in the microcantilevers depending on the beam characteristics. This was achieved by launching the striker bar at a velocity of approximately 11-24m/s. In order of significance; proofmass, cross section dimensions, notch size and beam length determined the shock level required to induce failure in the beams. Fig. 5 shows characteristic acceleration amplitudes and durations from the strain gauge and accelerometer signals over the 10,000-40,000g input range. Several failure modes were found upon inspection of the micro-cantilevers across all acceleration amplitudes. The high speed imaging (HIS) provided information on the cause of failure and determined if it was a result of the initial shock pulse or other phenomena such as fatigue failure after several oscillations and collision between beams. HSI also highlighted out of plane vibration of the beams in relation to the primary shock direction which can be caused by irregularities such as buckling of the beams under high strain or misalignments upon impact. Although, this may also be attributed to excitation of out of plane resonant frequencies as high frequency transients are unavoidable when inducing high amplitude low duration shocks. Fig. 6 is a schematic of a frame with 5mm, 6mm and 7mm beams of 100µm × 200µm cross section and 40% notch. Shock direction and notch geometry are highlighted, and a high speed image of the shock event shows the failure loci to be at the base of the 6mm and 7mm beams in the region of the notch. In this image, the primary mode of failure for the shock excitation of the beams is shown. Depending on the acceleration level, failure occurs approximately 150µs after the shock is induced at the base of the beam.

A JEOL JSM-840 scanning electron microscope (SEM) was used to obtain low and high magnification electron images of fracture surface using secondary electrons at 20kV. It was observed under SEM that, with or without a notch, failure occurred at the beam base as predetermined by ANSYS simulations of maximum stress location. In the presence of a notch, the region of smallest cross-sectionalarea acted as a stress raiser to instigate failure. However, it was found that even though the notch root yielded the highest stresses, once failure mechanisms initiated, the failure path followed the octahedral cleavage planes of the single crystal silicon i.e. the {111} planes.

The images depicted in Fig. 7(a-b) show secondary electron images of the fracture/cleavage surfaces for 6mm and 7mm beams respectively. These images both show three distinct surfaces; the notch surface can be seen in the right of the image (also clarified in Fig. 6), the first fracture surface showing the {111} cleavage planes can be seen in the centre of the images and the second fracture surface can be seen in the left of the images. These images show beams which have failed immediately under a 10,000g acceleration shock. The WLI of the etch profiles highlight that failure mechanisms are initiated at surface flaws from fabrication processes. The fracture origin is not seen in detail but lies along the edge of two faceted {111} cleavage surfaces in the region of the notch root caused by tension of the beam under flexure. With regard to the primary failure mode, Fig. 8 is a higher magnification of the lower (111) cleavage surface of the 7mm beam highlighted in Fig. 7(b). Yet again here the octahedral cleavage face is well defined, and in the mid to lower right hand side of the image, striations highlight that even under the influence of the [001] direction notch root the cleavage planes still recede to a common (111) slip plane. It was also seen that in the absence of a notch, cleavage remained on a single (111) plane and was not redirected onto a second (111) plane (as seen in Fig. 7). This indicates that the presence of a notch influenced cleavage planes to redirect.

Thus failure instigation occurred at the intersection of (001) and (111) planes in the region of maximum stress from beam flexure and/or near surface flaws from fabrication processes. The absence of mist and hackle in these surfaces indicates that the fracture did not reach terminal velocity along the middle part of the spreading fracture front. This was further attributed to the fracture moving into a region of lower tension.

With regard to the second fracture surface observed, and in agreement with [START_REF] Hull | Fractography: Observing, Measuring and Interpreting Fracture Surface Topography[END_REF], a brittle cleavage fracture surface normal to the tensile axis can be seen here with the presence of mirror, mist and hackle, as in Fig. 9. Here, fracture nucleates from the right and propagates to the left with the SEM image showing transition of the three stages. In this sample the mirror region dominates the fracture surface with the mist and hackle observed near the trailing edge covering approximately 2µm and 6µm respectively. From this it is evident that the crack tip had a much higher propagation velocity as it approached the edge of the beam in comparison with the first fracture surface. This higher velocity produces greater micro-mechanical activity at the crack tip where the increase in both stress intensity and rate of energy release results in the mist and hackle appearance.

The findings in this paper and their correlation with current literature prove the viability of both HPB as a test method and HSI as a metrology technique for testing micro-devices under high-g shock. Initial cleavage facets occurred predominantly on the {111} planes and immediate high acceleration failures highlighted rougher surface topography whereas lower acceleration fatigued failures yielded more consistent mirrored topography.

The secondary fracture surfaces indicated higher velocity fractures due to the presence of mirror, mist and hackle topography and fracture surface normal to tensile axis.

WLI of the etch profiles in conjunction with a study of the dominant fracture surfaces of SCS shows that while surface flaws produced from fabrication processes initiate failure mechanisms, crystal orientation plays a primary role in crack propagation direction.

Fig. 1 .

 1 Fig. 1. SEM Image of SCS micro-cantilever structure displaying the relative lattice planes and directions.

  (b) the typical scalloping and footing effects of the DRIE is shown across the 100µm sidewall depth. The scalloping caused by the passivation/etching cycle resulted in a corrugation of approximately 90-95 steps for the 100µm depth. The footing effect of the etch results in an undulating negative curve of approximately 500nm in the sidewall. The WLI results indicate that the surface roughnesses of the KOH and DRIE etches are insignificant relative to the overall scale of the cantilevers. It is not likely that surface profile will influence the strength of the beams, however the corrugated profile of the DRIE etch may develop failure instigation loci.

Fig. 3 Fig. 4 .

 34 Fig. 3 Interferograms of (a) KOH etch of back-side of proofmass (b) DRIE etch profile of beam sidewall.

5 a

 5 5MHz. The accelerometer signal was conditioned through a National Instruments (NI) SC-2345 and acquired by an NI-USB-6251 at 1.25MHz. An IDT MotionPro™ X high speed camera with 9-11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-Computar TEC-55 telecentric lens monitored the impact event on the cantilevers to verify the temporal point of failure.

Fig. 5 .

 5 Fig. 5. Characteristic shock acceleration pulses taken from accelerometer and strain gauge signals.

Fig. 6 (

 6 Fig. 6 (a) Beam schematic highlighting notch geometry and shock direction and (b) high speed image of test specimen at point of failure.

Fig. 7

 7 Fig. 7 Fracture surfaces of (a) 6mm beam and (b) 7mm beam.

Fig. 8

 8 Fig. 8 Lower (111) plane fracture surface on 7mm beam.

Fig. 9

 9 Fig. 9 SEM image of the trailing edge of fracture surface indicating the (I) mirror, (II) mist and (III) hackle region.

  Fig. 10

TABLE 1 FABRICATION

 1

	PROCESS RUN STEPS

6µm Al. sputter on back-side to act as etch stop in DRIE etch 12 Polyimide support layer Polyimide coat on back-side to support beams under vacuum in DRIE etch 13 DRIE etch Front-side etch of beam geometry using oxide mask -100µm depth -6-10µm/min 14 Finishing Removal of sacrificial layers -Al./nitride/oxide -cleaning Fig. 2. Fabrication process of micro-cantilevers.

ACKNOWLEDGEMENT

The author would like to acknowledge the Tyndall Institute in Cork, Ireland for the fabrication of the micro-cantilevers under the National Access Programme (NAP). The author would like to acknowledge Prof. Rick Hall, University of Delaware for his helpful support and the use of the HPB test equipment. The author would like to acknowledge the IRCSET (Irish Research Council for Science, Engineering and Technology) for their support of this work. The work is also supported by Science Foundation Ireland under grant number 03/CE3/I405.


G De Pasquale

A Soma'

A Ballestra

Mechanical Fatigue on Gold MEMS Devices: Experimental Results

The effect of mechanical fatigue on structural performances of gold devices is investigated. The pull-in voltage of special testing micro-systems is monitored during the cyclical load application. The mechanical collapse is identified as a dramatic loss of mechanical strength of the specimen. The fatigue limit is estimated through the stair-case method by means of the pull-in voltage detections. Measurements are performed by means of the optical interferometric technique.

INTRODUCTION

Reliability of MEMS increased its relevancy together with the spread of micro-devices applications in the every day life. The functioning of micro-fluidic bioMEMS, for example, can have a significant impact on the vitality of patients; aerospace applications impose more stringent conditions on the performance and reliability to limit repairs and replacements, etc.

About structural reliability, it is necessary to examine effects of process parameters, geometry, loads and working environmental conditions, which are all aspects whose knowledge is at present very scarce. MEMS reliability deals with materials but also with the reliability of their mechanical components. Moreover MEMS materials are generally used as thin films and this makes mechanical characterisation more complicated and at present no standard procedures are available for the extraction of fatigue properties and for the detection of all factors influencing fatigue behaviour as environmental humidity [START_REF] S B Brown | Materials Reliability in MEMS Devices[END_REF], building processes, etc. Also the utilization of multiple test structures for testing can introduce a device-dependency of results [START_REF] Soboyejo | Reliability assessment of polysilicon MEMS structures under mechanical fatigue loading[END_REF]. Utilization of materials in thin films generally enhances material properties of the material itself with respect to traditional applications, but still the determination of mechanical properties at microscopic scale is remarkably uncertain. This is due to the variation of material characteristics among different producers, among wafers of the same producer and even among structures on the same wafer. At present, most part of the researches about material reliability deal with polysilicon [START_REF] C L Muhlstein | Mechanism of Fatigue in Micron Scale Films of Polycrystalline Silicon for Micro Electro Mechanical Systems[END_REF], since it is the most used material both for mechanical and electric parts of MEMS; it can be considered as a material with a dominant brittle behaviour. About test methodologies, some different examples are reported [4 -9].

Present work is based on gold devices representing an example of ductile material to which essentially corresponds a lack of information in the literature. The mechanical strength of the structure is affected by the damage producing inside the material during the oscillation; the vibration test produces a cyclical load on the structure and cyclical concentrated stresses which are at the basis of the fatigue phenomenon. The consequence of this is the progressive lost of performances till the final failure that can be enhanced by the pull-in voltage decrease.

II. SPECIMENS AND TESTING PROCEDURE

The shape of test structures used is represented in figure 1 and the relative geometrical nominal dimensions are reported in table I; the device contains both the actuation part (represented by the perforated plate) and the specimen (the beam); this peculiarity is quite common in micro-systems for fatigue tests representing an example of the embedding inside a complete micro fatigue-machine. Described devices are designed ad-hoc by means of FE models in order to generate a shear stress inside the specimen [START_REF] De | Design and finite element simulation of MEMS for fatigue test[END_REF][START_REF] De | Reliability in MEMS: design of gold devices for mechanical fatigue tests[END_REF]. It is possible to observe that the beam deflection is much smaller than the plate length; this produces a rotation of the un-clamped extremity of the specimen approximately equal to 0.4 degrees, allowing to assume a vertical movement of it without rotation, corresponding to a pure shear excitation. The ITC-IRST [12] building technology is used; the lower electrode is realized by polysilicon, suspended parts are obtained through a sacrificial layer upon which a single (for the specimen) or a double (for the plate) gold layer is deposited. Nominal material parameters are E = 98.5GPa, ν = 0.42, ρ = 19.32ּ10 -15 kg/µm 3 . The value of static yield stress is studied in [START_REF] Espinosa | Size effects on the mechanical behaviour of gold thin films[END_REF] for gold material specimens and different thickness and width. In [START_REF] Espinosa | Size effects on the mechanical behaviour of gold thin films[END_REF] it is shown the strong influence of the dimension of the specimen on the tests results. This effect can be adressed to the intensification factor of the mechanical stress and to local effects. In the present work the failure of the specimen has been monitored by using global parameters such as pull-in voltage instead of local parameters as the mechanical stress.

Further work on identification of mechanical stress distribution on FE models will give the possibility to compare the stress local effect with the pull-in voltage variation.

The pull-in voltage results to be the most significant parameter able to monitor the cumulative damage of the material for the testing devices used. The pull-in of the not damaged structure is measured by supplying a DC voltage which is increased step by step; the cyclical load is generated by means of an AC voltage at the frequency of 20kHz. The excitation at resonance (about 28kHz measured) is avoided because of two main reasons: 1) the resonant amplification determines additional difficulties evaluating internal stresses in the material, 2) the progressive damage of the material could shift the resonant peak or modify the quality factor considering eventual changes of the structural stiffness and damping. Repetitive successive pull-in detections are performed and stored during the fatigue loading at fixed time intervals. The pull-in positioning is verified directly by an optical control through the interferometric microscope exploiting the translation of the fringes. The ZoomSurf3D (Fogale Nanotech) microscope [START_REF]Fogale Nanotech[END_REF] is used.

II. CYCLICAL EXCITATION VOLTAGE

The main purpose of a fatigue analysis is that the results produced by tests on specimens can be used for more general cases; this requires to extract from the specific test some fundamental quantities, as the maximum stress (σ max ), the minimum stress (σ min ), the mean stress (σ m = (σ max +σ min )/2) and the alternate stress (σ a = (σ max -σ min )/2) of the loading cycle, the stress ratio (R = σ min /σ max ), the fatigue limit (σ D ), and the number of cycles to failure (N).

The number of cycles of excitation needs to be carefully considered in the conversion from the frequency of the supply voltage to the loading cycles. The specimen oscillation is generated by the potential difference between the plate and the lower electrode that produces an attractive force; as reported in the diagram of figure 3, two periods T L of loading (or two cycles of actuation) correspond to one period T V of the alternate voltage. This is due to the present mode of deflection that not allows an oscillation around the undeformed position of the structure; the described situation takes place for all suspended structures in MEMS actuated by an alternate voltage. The resulting number of cycles N L of loading (and cycles of deflection) is related to the number of cycles of the alternate voltage N V by the relation

V L N N 2 = .
(1) It is evident that some portions of the specimen are always subjected to a local compression loading (here R is constant and equal to ∞) while symmetrical portions are always subjected to a local traction loading (here R is constant and equal to 0). These last regions are expected to be responsible of cracks nucleation. After these considerations all the fatigue loading parameters can be easily calculated from the value of the alternate load. The static displacement of the specimen is measured by the interferometric microscope at different levels of DC voltages and the curve reported in figure 4 is obtained.

This static acquisition is made assuming the absence of amplification of the vibration in the dynamic condition; this is reasonable considering that the excitation frequency is quite distant to the resonance.

III. EXPERIMENTAL RESULTS

The fatigue limit is estimated by the stair case method [START_REF] Collins | Failure of Materials in Mechanical Design[END_REF]; the reference number of cycles is fixed to 2ּ10 6 , the load step is 1V a and four load levels ranging from 12 to 15V a are selected. The results of stair case tests on six specimens are reported in table II, where the failure is indicated as 1 and the non-failure as 0. The fatigue limit at 50% of failure probability results to be V D = 13V; similarly the fatigue limits at 10 and 90% of failure probability are calculated as V D-10% = 12.3V and V D-90% = 13.7V.

V 1 14 V 1 1 13 V 0 1 12 V 0
Next figure 6 reports the measured values of pull-in voltages for different specimens during the accumulation of loading cycles. Only collapsed specimens are represented here; the failure is assumed to happen at the moment when the pull-in voltage registers a strong reduction of its amount. The figure 7 instead represents specimens that collapsed instantaneously because of the high level of fatigue loading (21 or 22.5V) and specimens that do not collapse as excited at a loading (10, 12 or 13V) equal or lower than the fatigue limit.

The Wöhler diagram resulting from reported curves is represented in figure 5 where the limit value is reported as pull-in voltage. Presented fatigue tests use the pull-in voltage as the parameter to monitor the progressive material damage; resonance frequency and quality factor have been used in the literature at this purpose, but they are observed not to be effective on devices used here because their variation is not appreciable. The pull-in detection determines a stress level in the material quite high if compared with the stress generated during the fatigue loading; by the FE model is estimated that the material reaches the yield condition locally. However this one-cycle actuation is not responsible for the plastic failure, as demonstrated by the fatigue loading at 0V reported in figure 7; here the specimen is only subjected to a series of successive pull-in actuations which actually not affect its strength.

All reported curves refers to stress-imposed fatigue tests; it is observed that, for high values of V a (and σ a ), the pull-in voltage curve can decrease during the fatigue loading till to reach the value of the alternate voltage of actuation. This condition causes a completely different type of vibration and the specimen results to be excited to the pull-in at each cycle; so the fatigue test becomes displacement-imposed and no more results comparable to the others can be obtained.

The mechanical fatigue collapse can by divided into three separate phases for simplicity: the crack nucleation, the crack stable propagation and the final failure. The last phase is not intended here as the complete rupture of the specimen, but as the dramatic loss of its mechanical strength. The first and the second phases are expected in the portions of the material which are subjected to an alternate traction loading (where R = 0).

The motion of dislocations in ductile materials takes place at cyclical stress levels much smaller than yield stress, resulting in a sliding movement of crystal planes near the surface of the loaded component. This process, influenced by the grain size, the lattice geometry, the temperature, etc. determines the superficial nucleation of the crack and its successive propagation inside the material. The additional effect of dislocations is to increase the local material strength (work hardening) when they accumulates at the grain borders or near a plasticized zone of the material; this is the situation resulting from the experimental curves reported in figure 6, where a sensible strength increase is evident before the collapse. The regions of the material that reach the plasticity condition after the pull-in detection determine a sort of barrier at the motion of dislocations in that zone.

The temperature influence on fatigue collapse is also observed during the tests; this parameter in fact strongly influences the motion of dislocations. In particular, when the time of the cyclical loading is increased between two successive pull-in detections, the pull-in value results to be quite smaller than the expected one; this is due to the additional mechanical heat transfer towards the material during the oscillation. The frequency of excitation also influences the temperature of the specimen.

Other factors affecting the described fatigue behaviour are the geometrical discontinuities (like the right angles used to connect the specimen with the plate and the support) which acts as stress multipliers, the typology of the applied stress and the building processes (especially related to thermal treatments and superficial roughness).

V. CONCLUSIONS

Mechanical fatigue behaviour of gold material applied to MEMS structures is studied by a special micro fatiguemachine; the pull-in voltage is used to monitor the damage progress and the resulting Wöhler diagram is traced in terms of alternate voltage. The fatigue limit is estimated through the stair-case method.

A further work will produce the developement of FE models to obtain a quantitative conversion from the fatigue parameters expressed in volts to material stress. The superficial analysis of damaged specimens by a SEM will recognize characteristic degradation features. The number of tests needs to be increased in order to give a statistically treatment to the results and other strategies to investigate the material damage could be adopted, as the detection of the specimen electrical resistance variation. Additional improvements to the analysis can be obtained by studying different specimen shapes or by adopting different actuation strategies.

Fig. 1 .

 1 Fig. 1. Geometrical shape of the micro fatigue-machine (the gap thickness under the plate is enlarged).

9- 11

 11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5

Fig. 2 .

 2 Fig. 2. Microscope image of the micro fatigue-machine.

Fig. 3 .

 3 Fig. 3. Qualitative actuation voltage cycle and fatigue load cycle.

Fig. 4 .

 4 Fig. 4. Conversion curve from alternate voltage to alternate equivalent stress.

9- 11

 11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5

Fig. 5 .

 5 Fig. 5. Wöhler diagram of fatigue tests curves reported.

Fig. 6 .

 6 Fig. 6. Experimental variation of the pull-in voltage with respect to the loading cycles; the fatigue loading is provided by the excitation voltage at 20kHz at the amplitude indicated for each curve.

Fig. 7 .

 7 Fig. 7. Experimental variation of the pull-in voltage with respect to the loading cycles; specimens collapsed immediately or non-collapsed are reported.

TABLE I NOMINAL

 I DIMENSIONS (IN MICRONS) OF THE MICRO FATIGUE-MACHINE

	specimen length	50	holes side	20
	specimen width	10	holes interspace	20
	specimen thickness	1.8	number of holes	10x4
	plate length	420	lower electrode length	420
	plate width	180	lower electrode width	460
	plate thickness	4.8	gap thickness	3


K I Sinclair

T Tilford

M Y P Desmulliez

G Goussetis

C Bailey

K Parrott

A J Sangster

Open Ended Microwave Oven for Packaging

HAL is

I. INTRODUCTION

Microwave radiation fundamentally accelerates the cure kinetics of polymer adhesives. It provides a route to deposit heat energy primarily into the polymer materials [START_REF] Wang | Microwave Heating of Metal-Filled Electrically Conductive Adhesive Curing[END_REF]. Therefore, microwave radiation can be used to minimise the temperature increase in the surrounding materials such as the substrate and die during the cure process. This is especially important for devices incorporating either low thermal budget materials or interfaces with large thermal coefficient mismatch. The concentration of heat into the polymer during the cure process promotes its adhesion properties, as the magnitude of residual stress will be low between the polymer and materials, to which it is being bonded.

Multi-mode waveguide cavities have been used for the curing thermosetting resins [START_REF] Wilson | Microwave Curing of Epoxy Resins[END_REF], [START_REF] Strand | Fast Microwave Curing of Thermoset Part[END_REF]. The ability to couple high power into a cavity oven enables fast polymerisation and large temperature gradients within the epoxy resin [START_REF] Gourdenne | Possible Use of the Microwave in Polymer Science[END_REF]. A material within a waveguide cavity oven with lower loss characteristics will remain at a lower temperature compared to the higher loss epoxy materials due to the selective nature of microwave energy absorption. Fundamentally, this characteristic reduces the thermal expansion mismatch between carrier and die allowing curing to be conducted in the presence of thermally sensitive components. As a consequence, it is possible to increase [or a least not decrease] the lifetime of the product [START_REF] Noh | Wafer Bonding using Microwave Heating of Parylene Intermediate Layers[END_REF]. Convention thermal curing heats the substrate first before transferring heat into the polymer while microwave curing process heats the polymer directly due to the volumetric nature of the absorption mechanism [START_REF] Jow | Dielectric and Temperature Measurements during Microwave Curing of Epoxy in a Sweeping Resonant Cavity[END_REF].

A maturing microwave curing technology is variable frequency microwave (VFM). Current VFM curing methods in electronic packaging bulk 'heats' the entire package or substrate [START_REF] Laugh | 2 to 18 Broadband microwave Heating Systems[END_REF]. For VFM, the source frequency is swept through a wide bandwidth, resulting in a scintillation of the model fields within the volume of the cavity resulting in a uniform field pattern. As such, a constant heating distribution can be obtained within homogenous materials [START_REF] Geisler | Variable Frequency Microwave Curing[END_REF]. In addition, sweeping the source frequency eliminates the likelihood of arcing thus allowing microwave heating of metals and semiconductor without damage due to electrostatic discharge (ESD) [START_REF] Ramanoorthy | Investigation of Variable Frequency Microwaves for Cure of Flip Chip Underfills[END_REF]. Single frequency microwave (SFM) heating has been reported [START_REF] Karmazsin | Use of Continuous and Pulsed Microwave for Quick Polymerization of Epoxy Resins[END_REF]. The source power can be pulased in order to avoid arcing. Microwave heating has been successfully applied to heating glob-top and flip-chip underfill encapsulants [START_REF] Wei | Use of Microwave Technology for Rapid Cure of Chip-On-Board Glob Top Encapsulants[END_REF]- [START_REF] Fontana | Variable Frequency Microwave (VFM) Curing of Die Attach and Underfill Materials[END_REF], conductive and non-conductive epoxy adhesives [START_REF] Li | Development and Characterisation of Single Mode Microwave Cured Adhesives for Electronics Packaging Application[END_REF], [START_REF] Wang | Microwave-Transmission, Heat and Temperature Properties of Electrically Conductive Adhesive[END_REF]- [START_REF] Wang | Microwave Cure of Metal-Filled Electrically Conductive Adhesive[END_REF], polymer dielectrics [START_REF] Farnsworth | Variable Frequency Microwave Curing of Photosensitive Polyimides[END_REF]- [START_REF] Tanikella | Rapid Curing of Positive Tone Photosensitive Polybenzoxazole Based Dielectric Resin by Variable Frequency Microwave Processing[END_REF], lead-free solder interconnects [START_REF] Moon | A Novel Technique for Lead-Free Soldering Process Using Variable Frequency Microwave[END_REF] and wafer bonding [START_REF] Noh | Wafer Bonding using Microwave Heating of Parylene Intermediate Layers[END_REF]. However, microwave heating is yet to be performed simultaneously with fine placement or alignment of the die or wafer onto the board. This lack of capability reduces the packaging throughput, as suggested by Fig. 1. The simultaneous fine placement and curing of the device into a larger assembly creates productivity gains by combining assembly, placement and bonding into a single processing step. The open-ended oven achieves curing by exposing the target package to the evanescent fields close to the 'open' exterior surface of a waveguide cavity resonator partially filled with a low loss dielectric [START_REF] Sinclair | A novel RFcuring technology for microelectronics and optoelectronics packaging[END_REF]. While it is well known that evanescent fields have been used to generate electromagnetic coupling in a variety of microwave devices, the use of such fields for microwave heating within packaging assemblies is novel to the best of the authors' knowledge.

This paper presents a curing analysis of a polymer-silicon target sample using a novel multi-physics numerical solver and results obtained through experimental analysis of the system. Numerical results show VFM heating and curing of a idealised polymer material using the system to be feasible. Experimental results demonstrate heating within a lossy target sample placed at the boundary between the dielectric filling and the open region. The source power is adjusted using a closed feedback routine allowing the temperature to be controlled within the target sample. The thermal distribution within the sample is recorded using a thermal imaging camera and forms the basis of the feedback control.

II. OPEN ENDED MICROWAVE OVEN

The cavity oven takes the form of a rectangular waveguide partially filled with a low loss dielectric material. The target sample is placed at a location within the air filled section of the cavity. Fig. 1 shows a schematic representation of the cavity. Resonant modes can be trapped within the dielectric even if one end of the cavity is open due to the quasi-open circuit formed by the dielectric-air interface and the shortcircuit at the copper wall. The fields within the air filled section are evanescent and attenuate exponentially with distance from the dielectric-air interface since the cut-off frequency of the air filled section of waveguide is less than the resonant frequency of the mode within the dielectric. Radiation losses can be minimised if the permittivity of the dielectric and the length of evanescent waveguide are chosen such that the evanescent fields are minimal at the open face. Power from the source is coupled into the cavity using either a probe (as in Fig. 1 and Fig. 3) located on the copper end wall or by slot aperture from waveguide. In order to take advantage of the continuity of the normal electric flux at the dielectric-air interface [START_REF] Sangster | Open-ended microwave oven for flip-chip assembly[END_REF], the oven is operated using a transverse magnetic (TM) resonance.

A sample of lossy material, placed in the evanescent field, will experience ohmic heating if sufficient power is supplied.

The absorbed power is primarily dependent on the magnitude of the average electric field within the sample volume and the dielectric loss tangent of the material itself [START_REF] Kraus | Electromagnetics[END_REF]. In order to maximise the dielectric power loss within the sample, i.e. the heating potential, the target sample should have a large loss tangent relative to the dielectric filling. An efficient heating system would be one that has minimal loss within the dielectric filling and heavy dielectric losses within the target sample.

The prototype oven shown in Fig. 1 can operate with a single low order mode (for SFM) or, in principle, with multiple higher orders modes for VFM. Sweeping the source will couple a large number of modes (or combination of degenerate modes) producing scintillation of the field pattern within the transverse plane of the evanescent section. A discussion of the electromagnetic properties of the openended oven can be found in [START_REF] Sangster | Open-ended microwave oven for flip-chip assembly[END_REF].

III. NUMERICAL MODELLING

In order to accurately model the process of microwave curing a holistic approach must be taken. The process cannot be considered to be a sequence of discrete steps, but must be considered as a complex coupled system. Microwave energy is deposited into the dielectric materials, inducing heating. This in turn progresses a cure process and induces thermal stresses within the material. The system must be considered as a whole because each of these processes fundamentally influences the others. For example, the EM fields effect temperatures but are, in turn, affected by temperature and cure state through a change in the dielectric properties of the material. Thus, despite the very different thermal and electromagnetic timescales, there is a requirement to closely couple the computational electromagnetic, thermal, curing and stress analyses.

The numerical model used in this contribution is an extension of the approach developed for food processing applications [START_REF] Tilford | Microwave Modelling and Validation in Food Thawing Applications[END_REF]. The model comprises a FDTD electromagnetic solver coupled with an UFVM multiphysics package. Separate meshes are developed for electromagnetic and thermophysical solutions, with coupling implemented through an inter-domain cross mapping process. Overlapping solution domains are defined; the thermophysical analysis mesh is confined to the load material while the electromagnetic mesh occupies the entire oven domain (encompassing the thermophysical domain).

Electromagnetic analysis is carried in out in a rectangular domain encompassing the extent of the proposed cavity oven. A classical Yee scheme [START_REF] Yee | Numerical Solution of Initial Boundary Value Problems Involving Maxwell's Equations in Isotropic Media[END_REF] has been implemented solving Maxwell's equations in the time domain with harmonic excitation on a tensor product computational mesh. In order to maintain numerical accuracy despite wavelength variation an automatic mesh generation algorithm has been employed.

Electromagnetic and thermophysical domains are interlinked through a cross mapping algorithm, based on a spatial sampling approach and capable of creating rapid conservative mappings between the meshes. Two sets of mappings are generated; firstly dielectric data is mapped from the thermophysical analysis domain into the FDTD ISBN: 978-2-35500-006-5 domain prior to EM solution. Subsequently, power densities are mapped from the FDTD domain into the thermophysical analysis domain prior to solution of the thermophysical PDE's.

IV. EXPERIMENTAL ANALYSIS

Experimental studies of the system have been conducted using a prototype system based on coupling of the TM 33 modes between 8Ghz and 12Ghz. Optimised coupling of the cavity over a wide bandwidth (required more VFM heating) has yet to be implemented within the design; As such, only SFM experimental results are presented within this paper.

A 25W travelling wave tube with protection circuitry and regulated power supply is connected to the cavity with a coaxial transmission line. A thermal imaging camera (X) is used to record the temperature distribution within the test sample and provides the basis for the control routine implemented using Labview [27]. A temperature profile can be specified allowing the temperature within the test sample to be controlled through automatic regulation of the source power using the feedback routine.

A cavity with dimensions 25.5mm x 25.5mm x 110mm was fabricated encompassing a low loss, ceramic dielectric material and is shown in Fig. 3. The dielectric filling has dimensions equal to 25.5mm x 25.5mm x 100mm with a relative permittivity at room temperature of 6 and tan δ of 0.0005 at 10GHz. The dielectric is recessed within the cavity at the open end, allowing for an air region of 10mm in length.

The test sample consists of a lead free solder paste placed between two circular coverslips of 15mm diameter with thickness of 0.15mm. The overall thickness of the sample can varied between 0.4mm and 0.6mm. This variation in test sample volume has insignificant impact on the electromagnetic performance of the cavity as the perturbation of the resoance frequency of the coulped mode is negligable. The complex permittivity of the lead-free solder was measured at room temperature using a dielectric probe (Agilent 85070E) and was found to have a relative permittivity of 4.6 and a tan δ of 0.6 at 10GHz. The cavity was excited by a coupling probe located centrally within the transverse plane on the copper end wall. The combination of parameters described above results in several TM 33 modes between 10GHz and 10.8GHz (a frequency range for which microwave measurement equipment is readily available). Fig. 4 displays the frequency response of the cavity over a bandwidth of 10GHz to 10.8GHz. The length of the probe is optimised to produce maximum coupling of the TM 335 mode resonant at 10.424GHz. |S 11 | = -13.2dB at this frequency indicating good impedance matching.

A measurement of the cavity incorporating temperature feedback was conducted. The average temperature over the area of the coverslip is recorded using a thermal imaging camera. A target temperature versus time is plotted (red) over a period of 170s. The source power is automatically adjusted such that the temperature of the sample is controlled. Fig. 5 shows agreement between target and measured temperature within the lossy test sample thus demonstrating the validity of the control routine. The measurement was repeated with no ISBN: 978-2-35500-006-5

sample present within the cavity. The temperature of the dielectric surface is plotted over the same period. Similar coupling is achieved for the TM 335 mode within this geometry allowing for comparative power levels to be delivered to the cavity. The surface temperature of the dielectric increases by 7 o C after 170s. Therefore, the increase in the temperature within the test sample is primarily due to emersion within the evanescent field as opposed to thermal conduction from the dielectric filling material which is, in this example, negligible. The temperature distribution within the test sample after 75s is plotted in Fig. 6. A uniform temperature distribution is displayed within the test sample. Although the cavity is optimised for the TM 335 mode, it is clear that the thermal distribution does not correlate with a 'pure' TM 3,3,5 mode pattern as anticipated. This is due to quasi-degeneracy and is attributable to the fact that the available Q-levels are insufficiently high, causing the modal resonances of neighboring modes to overlap. As such, unwanted resonant modes are be coupled into the cavity disturbing the idealised TM 335 field pattern. The effect of the quasi-degenracy can be reduced by lowering the dielectric loss of the dielectric material forming the resoanting region or by increasing the conductivity of the metal cavity walls. Quasi-degeneracy can also be suppressed by employing mode selective coupling involving the use of more than one excitation probe.

V. CONCLUSION

This paper outlines the design and implementation of the open-ended waveguide cavity oven. A coupled numerical solver has been developed that allows the electromagnetic fields, temperature distribution, degree of cure and stress within the target package to be characterised. Numerical results demonstrate that, not only does the design of the cavity support heating using the evanescent fringing fields within a cutoff region of the cavity, but also that sufficient heating is obtained to facilitate curing of a polymer load. Experimental outputs demonstrate that control over the temperature within the load can be achieved through utilising a closed loop feedback routine. Finally, the experimental output supports the conclusion that the heating mechanism within the load occurs through ohmic losses due to interaction with the evanescent fringing field.

Further work includes the optimisation of the evanescent field through inclusion of a frequency selective surface (FSS) or layered dielectric configuration at the dielectric-air boundary and the characterisation of properties of microwave cured epoxies and encapsulant using the open-oven geometry.

Fig. 1

 1 Fig. 1 Schematic showing layout of the cavity oven and the comparison between a 'conventional' assembly process and the proposed open-ended oven.

Fig. 3 .Fig. 4 .

 34 Fig. 3. Experimental cavity with dielectric shown on the right hand side.

Fig. 6 .

 6 Fig. 6. Temperature distribution within target sample.

Fig. 5 .

 5 Fig. 5. Temperature versus time relationship showing the temperature observed within the test sample and at the surface of the dielectric when no sample present.

ACKNOWLEDGEMENTS

This work was supported by the Innovative Electronics Manufacturing Centre (IeMRC), co-ordinated by Loughborough University, United Kingdom, through the project FAMOBS (FE/05/01/07). This work was also made possible through the funding of the project 3D-MINTEGRATION (EP/C534212/1) sponsored by the Engineering and Physical Sciences Research Council (EPSRC). 9-11 April 2008 © EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5


Tenghsien Lai

Changhan Huang

Chingfu Tsou

Design and Fabrication of Acoustic Wave Actuated Microgenerator for Portable Electronic Devices 1

HAL is

INTRODUCTION

The past few years have seen an increasing focus on energy harvesting issue, including power supply for portable electric devices. Utilize scavenging ambient energy from the environment could eliminate the need for batteries and increase portable device lifetimes indefinitely. Several different ambient sources, including solar, vibration and temperature effect, have already exploited [START_REF] Raffaelle | integrated solar power system[END_REF][START_REF] Roundy | Micro-electrostatic vibration to electricity converters[END_REF][START_REF] Strasser | Micromachined CMOS Thermoelectric Generators as On-Chip Power Supply[END_REF]. In particular, scientists are currently trying to find more efficient ways of utilizing solar cells for electricity production [START_REF] Elgamel | High efficiency polycrystalline silicon solar cells using low temperature PECVD process[END_REF][START_REF] Rotich | Micromachined thin solar cells with a novel light trapping scheme[END_REF][START_REF] Strobl | European roadmap of multijunction solar cells and qualification ststus[END_REF]. Each energy source should be used in suitable environment, therefore to produce maximum efficiency.

Ambient energy generators have the potential to replace battery power as an energy source in a variety of practical applications. For example, kinetic (or vibration) energy generators harvest electrical energy from movement present in the application environment. Various kinetic generators use piezoelectric materials [7~10], electrostatic [11~13] and electromagnetic

[14~20] transduction mechanisms. Piezoelectric materials are perfect candidates for vibrational energy scavenging as they can efficiently convert mechanical strain to an electrical charge without any additional power, but it is difficult to implement at micro-scale. Electrostatic is a good solutions for the construction of microgenerators because of its relative ease of integration with microelectronics. The disadvantage of this approach is that a separate voltage source is required to charge the capacitor. Electromagnetic conversion uses vibration to move a conductor in a magnetic field. Existing prototypes generate low voltage output to be usable. However, the amount of energy generated by those approaches depend the application environment and the efficiencies of the generators and the power conversion electronics. The goal is to develop miniature, self-contained and renewable power supplies that convert energy from an existing source in the environment into electrical energy.

In this study, we present an acoustic wave actuated microgenerator with high output voltage suitable for commercial portable electronic devices. We use the energy of acoustic waves, such as the sound from human voices or speakerphone, to actuate a MEMS-type electromagnetic transducer. This provides a longer device lifetime and greater power system convenience. It is convenient to integrate MEMS-based microgenerators with small or portable devices. Our aim is to run a small electronic device, like a cell phone, and may soon be able to power a laptop.

II. DESIGN AND ANALYSIS

The proposed acoustic wave actuated microgenerator, as shown the Fig. 1, consists of a top wafer with planar coil and a bottom wafer with a suspended flat and support beams as well as a high magnetic energy permanent magnet. The top wafer is etched through its thickness by anisotropic wet etching (KOH). The wafer-to-wafer bonding process is aligned to ensure correct placement of the coil relative to the magnet. The top surface of the magnet is in line with the planar coil. The thickness of the magnet is same as that of the top wafer. Thus, the planar coil has tight gap spacing with the permanent magnet; this enhances the induced voltage. Larger output power can be obtained by link an array of identical microgenerators.

ISBN: 978-2-35500-006-5 Cross-section view.

The commercial software ANSYS was used to simulate the resonant frequency and vibration mode of suspension microstructure for the actuation frequencies produced by human vocal cords. The purpose is to achieve the micro-generator with higher charge collection efficiency in a small area under a specific driving frequency. Noting the frequency range of 200 Hz -1.5 kHz in sound pressure level characteristics is the most important range in human vocal and electrical acoustic instruments, as shown in Fig. 2. The schematic diagram of the mechanical structure is shown in Fig. 3(a). The middle point of each suspension beam is fixed to the silicon substrate. The material parameters and device dimensions of the microgenerator are shown in table 1. The typical simulated result for the suspended flat with a permanent magnet is shown in Fig. 3(b). The figure demonstrates the first resonance mode of the microstructure is an out-of-plane piston-like motion and the natural frequency is 1012 Hz. The electromagnetic software Ansoft Maxwell EM3D was employed to characterize the variation magnetic flux due to the mechanical and electrical energy conversion is coupled through the magnetic field across the conductor coil. In this simulation process, we design fifteen-turn square spiral coil made from Cu with 20 µm gap spacing on a substrate. Each coil has a width of 20 µm and a thickness of 10 µm. The first-turn coil has one winding over the edges of the magnet. The size of magnet is same as in table 1. Figure 4 show the typical simulation result for the magnet above one-turn planar coil at a distance of 10 µm.

The figure shows that the bottom edge of magnet closer to the coil, the more magnetic flux is induced. The magnetic flux drops very rapidly as the magnet moves away from the center of the coil. By the post process function of the Maxwell EM3D, the magnetic flux of each planar coil was calculated. For a 3 mm × 3 mm microgenerator, the estimated output voltage is approximately 0.58 mV at 1 KHz driving frequency with ± 50 µm amplitude range.

III. FABRICATION

To meet the requirements of high-voltage output under a specific range of driving frequency, we used electroplating and bulk micromachining to fabricate the planar coil and suspension microstructure, and then integrating an Nd-Fe-B based permanent magnet on the suspended plate. The fabrication processes of the planar coil suspension microstructure are shown in Fig. 5 and Fig. 6, respectively. Figure 5 shows the fabrication processes of the planar coil that starts with a 500 µm thick 4-inch <100> silicon wafer (Fig. 5a). First, the (100) single crystal silicon substrate was placed in a furnace at 1050 °C for 150 minutes to grow a 1 µm thick wet thermal oxide layer (Fig. 5b). Then a 0.1 µm-thick Cr film and a 0.2 µm-thick Ni film were deposited by thermal evaporation as adhesion and seed layers respectively (Fig. 5c). Next, a 25 µm-thick of photoresist AZ4620 was spun and patterned by photolithography. After that, a Ni layer was deposited by electroplating to fabricate the planar coil, and then the residue photoresist was removed (Fig. 5d). It is noted that the Cu with the excellent conductivity is suitable as the coil material than the Ni, however the process of electroplating Ni is easy to verify the possibility of our design and to improve the yield. To isolate the electrical properties, the exposed region of Cr adhesion and Ni seed layers were etched by a wet etch solution (Fig. 5e). The backside oxide layer was patterned by photolithography and then etched by BOE solution (Fig. 5f). After that, the exposed regions of silicon were wet etched by KOH solution (25%) at 80 °C for 4 hour to through the wafer thickness, and finally SiO2 layers were etched by HF to obtain a single layer of Ni microstructure (Fig. 5g). The fabrication process of the suspension microstructure in Fig. 6(a) ~ (g) is same as that of the planar coil except step (d) and the mask layout. The Ni suspension microstructure is deposited by electroplating technology and is 20 µm thick. The electroplated Ni film has the advantages of low-residual stress, high density, and superior mechanical strength and can be formed as a high-aspect-ratio microstructure by photolithography [START_REF] Sharpe | Mechanical properties of LIGA-deposited nickel for MEMS transducers[END_REF][START_REF] Mazza | Experimental determination of mechanical properties of Ni and Ni-Femicrobars[END_REF][START_REF] Greekand | Young's Modulus, yield strength and fracture strength of microelements determined by tensile testing[END_REF]. As to the mechanical properties of electroplated Ni film from the previously works [START_REF] Sharpe | Mechanical properties of LIGA-deposited nickel for MEMS transducers[END_REF][START_REF] Mazza | Experimental determination of mechanical properties of Ni and Ni-Femicrobars[END_REF][START_REF] Greekand | Young's Modulus, yield strength and fracture strength of microelements determined by tensile testing[END_REF][START_REF] Stephens | Mechanical property evaluation and failure analysis of cantilevered LIGA nickel microposts[END_REF] referred and the quasi-static bending beam test using nanoindentation system in our research [START_REF] Tsou | Bending Characterization of Electroplated Nickel Microbeams[END_REF] presented are summarized in Table Our research presented [START_REF] Mizuno | Investigation of a resonance microgenerator[END_REF] 179~225 660~1120

IV. EXPERIMENTAL RESULTS

This research has successfully designed and fabricated an acoustic wave actuated microgenerator using the presented processes. Typical fabrication results for the planar coil and suspended plate are shown in Fig. 7 and Fig. 8, respectively. Figure 7 show the square spiral coil has fifteen-turn with 20 µm gap and each planar coil has a width of 20 µm and a thickness of 10 µm. The measured resistance from pads is approximate 58 Ω. The size of etching through-hole is 2 mm × 2 mm. Figure 8 shows the suspension microstructure with slight bending deformation from the fix end is due to the effect of residual stresses. This result would not affect the latter assembly and performance measurement of microgenerator significantly. Fig. 9 shows the top view of a 2 mm × 2 mm × 0.5 mm microgenerator with a permanent magnet. The pad setting is for the wire bond to output the induced electrical energy. The residual magnetic flux density (Br) of the magnet is 1.2 Tesla, according to the manufacturer's data.

To understand the dynamic characteristic and the performance of the microgenerator, Laser Doppler Vibrometer (LDV) and an acoustic speaker with a power of 3W were employed to measure the resonance frequency and induced-voltage of the microgenerator, as shown in Fig. 10. The functions of the power supply and oscilloscope are to provide the initial bias-voltage to detect the output voltage, respectively. The gap spacing between speaker and microgenerator was 5 mm. After a driving voltage of 1V (peak to peak) with sin wave signal has been input the speaker, the dynamic responses measured at atmospheric pressure by LDV is show in Fig. 11(a Also, if the driving voltage is increased to 10 V, a maximum out-of-plane displacement of 11.5 µm can be achieved, as shown in Fig. 11(b). In this case, the measured induce-voltage at the first resonance frequency 470 Hz is 0.24 mV (peak to peak) when the driving voltage is 1 V. Therefore, using the proposed fabrication methodology and operation principle, we are able to manufacture a microgenerator with high-power output. The generated energy can apply to an electric cell or as a secondly battery, although the measured amount of electrical energy could not supply enough power to run a cell phone. In addition, high induced-voltage can be achieved by using the coil with high conductivity such as Cu or under a specific driving frequency by increasing the thickness and the number of coils, as well as by reducing the gap spacing between coil and permanent magnet. Comparing experimental results to analytic predictions, we note many differences in first natural frequency and induced-voltage. We suggest many differences arise from the error in material parameters and in manufacturing geometry precision, the effect of the thickness of suspension structure may have been particular disruptive. The measured thickness of electroplated Ni is about 14 µm, which is less than the 20 µm in the proposed design. In addition, the measured amplitude is only ± 3 µm, also less than the intended ± 50 µm significantly. However, we can predict the performance of the microgenerator accurately by using available data from past experiments and eliminating all errors associated with critical device dimensions to meet some specific acoustic driving frequency.

V. CONCLUSION

A novel configuration for an electromagnetic microgenerator has been modeled and confirmed by experiments at a small macroscopic scale. The performance of the planar coil and suspension structure has been characterized by software analysis and experimental results. Fabrication result show a 3 mm × 3 mm× 1 mm microgenerator had maximum induce-voltage 0.24 mV obtained at a driving frequency of 470 Hz. Based on a MEMS type electromagnetic transducer, our design provides several important features such as the planar coil has the same level and a small gap spacing with the permanent magnet in in-plane to enhance the induced voltage. The suspension microstructure can be modified appropriately to meet a specific acoustic driving frequency to increase the efficiency of energy conservation. The small size, high efficiency, and potentially low cost make microgenerator desirable for numerous portable electronic devices. In the future, microgenerators can be combined with batteries and integrated circuit to make energy harvesting system for small devices.

Figure 1 .

 1 Figure 1. Design concept of the microgenerator: (a) 3-D and (b)Cross-section view.

Figure 2 .Figure 3 .

 23 Figure 2. Frequency responses of human vocal formant and commercial acoustic instruments

Figure 4 .Figure 5 .Figure 6 .

 456 Figure 4. The magnetic flux in the space simulated by Ansoft MaxwellEM3D

2

 2

  .

  ) and 11(b). Fig.11(a) shown the magnet had a maximum displacement 2.8 µm at the first resonance frequency (piston-like motion) of 470 Hz.

Figure 7 .Figure 9 .Figure 10 .Figure 11 .

 791011 Figure 7. SEM image of the top wafer with planar coil

Table 2

 2 Mechanical properties of electroplated Ni film.

	References	Young's modulus	Yield strength
		E (GPa)	σy (MPa)
	[11-14]	125~231	323~1550

ACKNOWLEDGMENT

This project was supported by the National Science Council of Taiwan under grant of NSC-95-2221-E-035-016. The authors appreciate the Precision Instrument Support Center of Feng Chia University, and the NSC National Nano Device Laboratory (NDL) in providing the fabrication facilities.