C H Huang

C Tsou

T H Lai

High Density Out-of-Plane Microprobe Array 1

In this paper, the high density out-of-plane microprobe array is demonstrated. The fabrication processes of proposed device including bulk micromachining, thin film deposition and electroplating. By depositing the various thickness of thin film with residual tensile stress, the deflection of bending beam could be precisely controlled. The horizontal variation of bending beam array could also restrict in a few micrometer. The concept for the device fabrication is depositing the Ti thin film on a suspension SiO 2 cantilever to bring out-of-plane bending deformation, and then use electroplating process such as low stress Ni film to increase the stiffness of the probe structure without change the beam's bending profile. Through the different electroplating bath, the bending microprobe with different material, such as Ni and Cu, could be fabricated for IC testing and interconnection applications. According to the mask design, the high density probe array could easily achieve. The prototypes of the 40×40 microprobe array in pitch sizes of 50µm, 80µm and 100µm were fabricated and characterized. Measurement result by the optical interferometer shown the mean bending height of twenty samples are about 3µm, 9µm and 26µm in different beam length 25µm, 55µm and 85µm, besides the maximum horizontal variations are less than ±0.8µm, ±1.2µm, ±2µm that corresponded to the pitch size of 50µm, 80µm, and 100µm, respectively.

I.

INTRODUCTION

MEMS technology has been developed rapidly in the last few years. More and more special micro structures were discussed in several publications. However, all of the structures were produced by consisting of the three fundamental structures, which included bridge, cantilever and membrane structures. Even the more complex structures were no exception. The cantilever with the property of simple design and easy fabrication among three kinds of fundamental structure; therefore, it was popular used in the design of MEMS device. In this paper, the cantilever beam used as the probe microstructure has been achieved with high density and out-of-plane bending deformation. The proposed microprobe array could widely applied in IC testing [START_REF] Zhang | Thermally actuated microprobes for a new wafer probe card[END_REF][START_REF] Kataoka | Low contact-force and compliant MEMS probe card utilizing frittingcontact[END_REF][START_REF] Shingo | Design and fabrication of an electrostatically actuated MEMS probe card[END_REF][START_REF] Kim | Silicon micro-probe card using porous silicon micromachining technology[END_REF] and flip chip interconnect [START_REF] Smith | Flip-chip bonding on 6-µm pitch using thin-film microspring technology[END_REF][START_REF] Haemer | Flexible micro-spring interconnects for high performance probing[END_REF][START_REF] Lunyu | Compliant cantilevered spring interconnects for flip-chip packaging[END_REF][START_REF] Hantschel | Fabrication of highly conductive stressed-metal springs and their use as sliding-contact interconnects[END_REF][START_REF] Chow | SolderFree Pressure Contact MicroSprings in HighDensity FlipChip Packages[END_REF], as shown in Fig. 1(a) and Fig. 1(b), respectively.

Due to the develop roadmap of the IC chip, the scale of line width were decrease rapid in every generation. Therefore, the technology of the fabrication, package and test were changed in order to compatibility with the IC development. In the test technology, the probe card plays an important role in the IC test. The property of fine pitch and the high pin count probe array will the trend of the IC development. Hence the traditional probe card which assembled by the manual were facing the bottleneck. Utilizing the out-of-plane bending structure fabricated by the MEMS technology could easily achieve these restrictions. Because of the out-of-plane microprobes could easy arrange in the fine pitch and the demand shape by the mask design. Consequently, the out-of-plane microprobe had the great potential to replace the traditional probe card. Besides, these out-of-plane microprobes also could use as the flip chip interconnects. As a result of the scale down to the micrometer degree, the scale factor will be concern in many physical phenomena. Particularly, the thermal expansion was one kind of the physical phenomena which were produce thermal mismatch between the flip chip and package substrate at the micrometer scale degree. Hence, the traditional flip chip interconnect of the solder ball will face to the challenge. Utilizing the out-of-plane microprobe array as the flip chip interconnect could overcome this problem. The thermal stress was released through the deformation of the bending structure. Therefore, it will be the candidate for the next generation interconnects of the flip chip.

Utilizing the out-of-plane bending structure as the microprobe or interconnect, the electrical or mechanical property will be concern in the every report but no one discusses the horizontal variation in each probe. However, the horizontal variation in each probe will be the important property along with the increasing probe number. If the probe array with the minima horizontal variation in each probe, the requirement of the mechanical property will be not so strict. Therefore, this paper focus on the utilizing of the simply and cheap process to manufacture the out-of-plane microprobe array which have the minimal horizontal variation.

There are several MEMS approaches reported on the fabrication of out-of-plane bending structure. Most of all fabrications of the out-of-plane structure were by the two layers material with different stress [START_REF] Kataoka | Low contact-force and compliant MEMS probe card utilizing frittingcontact[END_REF][START_REF] Shingo | Design and fabrication of an electrostatically actuated MEMS probe card[END_REF][START_REF] Kim | Silicon micro-probe card using porous silicon micromachining technology[END_REF], but those methods usually have some disadvantage. First, the structure wasn't robust enough and couldn't load with larger force than single layer. Supposing the bi-layer structures load a force, it was easily peeling between two layer interfaces. Second, these bi-layer structures have much material restrict in the fabrication. In order to obtain bending up structure, the material must be concern seriously. Therefore, the fabricate process were restrict from the material choice. In this paper, the simple fabrication was provided to achieve high density out-of-plane probe array with single layer. By depositing a tensile stress metal film on a suspension beam, the out-of-plane bending structure was formed. As such, the bending deformation could be controlled by the thickness of tensile stress metal. Moreover, a minimal horizontal variation in each microprobe could obtain. Furthermore, the electroplating process was employed to form the main probe structure. Utilizing this fabrication process not only obtain better mechanical characteristic but also have more flexibility in material choose, such as Cu and Ni-Co probe array.

II. DESIGN

The height density out-of-plane microprobe array was designed and demonstrated. The design concept of the out-of-plane microprobe array was shown in Fig. 2. The fabrication technology of proposed device included bulk micromachining, thin film deposition and electroplating. The characteristic of the out-of-plane microprobe array can be controlled by the deposition thickness and materials of the thin film. For example, the metal film such as aluminum (Al) by evaporation has the compression residual stress but titanium (Ti) brings tensile residual stress and the stress magnitude varies with the thickness of metal film. However, if the material with large tensile stress, such as chromium (Cr), used as a stressed cover layer, the cantilever beam will roll up or break. Therefore, the choused material was very important in the thin film deposition for the bending microprobe. In this paper, the Ti is used as the stressed material which cause the suspension microprobe bend up. By controlling the deposition thickness, the out-of-plane deformation of each microprobe could restrict in a few micrometer. After the pre-bending structure was fabricated by deposit thin film process, the electroplating process could used to increase the stiffness of the probe microstructure. In this paper, the electroplated nickel film was used as the pbobing structure. The nickel fabricated by electroplating process usually has lowered residual stresses than others materials. Therefore, it affects the deformation of the pre-bending probe slightly. Besides, the single layer structure could avoid some problems which will occur on the bi-layer structure such as peeling between two layer interfaces. Hence, the life-time of the microprobe could be increase.

In the mask design, the probe tip could design in different shape to satisfy different applications. For instance, the probe tip design in triangle shape could use in IC testing that pads made from Al or Cu has native oxide layer. The tip with triangle shape will effectively break the oxide film. In the other hand, when the probe tip design in the flat tip, the probe could use as the interconnect in the IC packaging. The tip with flat shape will increase the contact area with the substrate. Therefore, base on those different applications, the probe will design according to the user's requirement. The prototype of the micro probe array is designed in the three specifications, the pitch between the microprobe arrays are about 50µm, 80µm and 100µm respectively. The detail geometry parameter correspond to different pitch size between etch other is shown in table1.

In this paper, we emphasize utilizing the simple and reliable process to manufacture the microprobe array with minimal horizontal variation. Therefore, the prototype is not concern in detail such as electrical isolation between each probe and external environment electrical connecting; we just focus on the microprobe structure itself. However, we already have some ideal to overcome these problems. Utilizing the through-hole electrode, the microprobe could connect to external environment. In the order hand, the laser cutting process could isolate the electrical between each microprobe. According to the above mention methods, it could completely realize the high density, fine pitch commercial microprobe array.

III. FABRICATION

The out-of-plane structure array is fabricated with bulk machine and thin film deposition techniques. First step was the fabrication of suspended cantilever beam by photolithography and bulk etching process. Then deposit a metal with tensile stress on the cantilever beam to form the pre-bending structure. Finally, the electroplating process is used to enhance the stiffness of the structure. The detail process flow as show in Fig. 3 and describe in the following statements. First, the thermal oxide with 1µm thickness was grown on the silicon substrate (Fig. 3a). Next, define the release region around out-of-plane structure by photolithography process and oxide was wet etched by buffer oxide etcher (BOE) (Fig. 3b). Utilize oxide layer as the hard mask to etch the silicon substrate. The silicon substrate was etched by consistency 25% KOH solution at 75°C (Fig. 3c).

Then evaporate the Ti with 4000Å thickness on the suspension cantilever beam (Fig. 3d). The cantilever beam was bending upward due to the tensile residual stress in titanium layer. Next, the nickel was electroplated to enhance the stiffness of the structure (Fig. 3e). Finally, remove the oxide layer and etch titanium metal layer under the nickel cantilever beam to obtain the single layer structure (Fig. 3f).

In the step (d), if we do not execute next process rapidly, the Ti expose in the air environment will grow a native oxide layer and affect the follow up electroplating process. Therefore, the nickel (Ni) could use to deposit on the Ti layer as the seedlayer for following electroplating process. The nickel layer depositing by evaporation have low residual stress characteristic. Consequently, when the nickel deposit on the pre-bending beam, it'll not be affected the pre-banding height so much.

IV. RESULT AND DISCUSSION

The scanning electron microscope (SEM) image of the fabricated microprobe structure with pitch size 50µm, 80µm and 100µm are shown in Fig. 4(a), 4(b) and 4(c), respectively. According to the SEM image, the bending height with excellent horizontal uniformity was observed. The deflection of each out-of-plane microprobe is determined by the optical interferometer and the measurement result shown in Fig. 5 In order to investigate the uniformity in each array which consist of 1600 (40×40) microprobes. The average value is obtained through the sampling twenty points in each specification. Figure 6 shows the measurement result about the relationship between the beam length and bending height. The error bars in this figure represent the deflection variation.

According to the Fig. 6, we could find some phenomenon in the experiment result. The beam length is in direct proportion to the bending height and the horizontal variation is also increasing along with the beam length. Because of the stiffness of the microprobe affect the horizontal variation in each array with different specification. The microprobe has the longer beam length, the lower stiffness is achieved. Therefore, the more horizontal variation is caused by the residual stress of the thin film. The prototype in each specification show the deflection variation are ±0.8µm, ±1.2µm and ±2µm corresponding to the pitch of microprobe array 50µm, 80µm and 100µm, respectively. The microprobe array with different material could easy fabrication by the different electroplating bath in the same process that provided in this paper. Fig. 7 shows the microprobe array fabricated by the Cu. Although the Cu is used to fabricate the microprobe, the characteristic of the excellent uniformity also exist. When the microprobe used in the IC test, the material of Ni is chosen as the main structure. Because of the nickel have more stiffness and also could endure the higher contact force than others. In the other hand, if the material of copper is chosen as main structure it is suitably use in the interconnect due to the lower resistance. Therefore, the microprobe array can made from different materials according to the various applications. In this paper, the high density out-of-plane microprobe arrays with excellent uniformity are demonstrated. The variation of out-of-plane deflection is ±0.8µm, ±1.2µm, and ±2µm corresponding to the pitch 50µm, 80µm, and 100µm, respectively. According to the fabrication process of the microprobe, we can use the effect of residual stresses of thin films deposition to bent the flexible micromachined probe with a large out-of-plane predeformation and combine post-electroplating technique to further enhance its stiffness. Besides, it is clear that using the process reported here can fabricate a microprobe array with the characteristics of both lower residual stresses and minimum deflection variation under large out-of-plane deformation. This technique has the potential for achieving high production yield with fine pitch, low cost, high pin counts and area array testing as well as flip chip packaging.

  Fig.1The applications of the out-of-plane bending cantilever beam.

Fig. 2

 2 Fig.2 The design concept of out-of-plane microprobe array Table1. The different pitch of microprobe array correspond to geometry parameter pitch(µm) Beam-length(µm) Beam-width(µm) 50 25 6 80 55 10 100 86 10

  (a), 5(b) and 5(c). The result shows the bending height are 3µm, 9µm and 26µm corresponding to the pitch 50µm, 80µm and 100µm, respectively.

Fig. 3

 3 Fig.4 The SEM image of the microprobe array under the different pitch size

Figure 6 .

 6 Figure 6. The relationship between the beam length and bending height

Figure 7 .

 7 Figure 7. SEM image of the out-of-plane microprobe array with copper electroplated

ACKNOWLEDGMENT

The authors would appreciate the financial support by Taiwan Semiconductor Manufacturing Company (TSMC), and the fabrication process support by Precision Instrument Support Center of Feng Chia University and the National Nano Device Laboratory (NDL) in providing the fabrication facilities.


R Voicu

R Muller

L Eftime

DESIGN OPTIMIZATION FOR AN ELECTRO-THERMALLY ACTUATED POLYMERIC MICROGRIPPER

published or not. The documents may come

INTRODUCTION

Microgrippers are promising tools used as endeffectors for systems that can handle and manipulate micro and nano scaled objects with application in various field of science and industry [START_REF] Kim | Metallic microgripper with SU-8 adaptor as end-effectors for heterogeneous micro/nano assembly applications[END_REF][START_REF] Kim | Polysilicon microgripper[END_REF].

In the MEMS area, there are many applications where microgrippers have reached an increasing importance. The applications of microgripper structures range from micro manipulation of micro particles, micro components and even cells to assembling and medical applications, and already have a serious impact on present and future technologies [START_REF] Kim | Metallic microgripper with SU-8 adaptor as end-effectors for heterogeneous micro/nano assembly applications[END_REF][START_REF] Kim | Polysilicon microgripper[END_REF][START_REF] Butefisch | Novel micro-pneumatic actuator for MEMS[END_REF][START_REF] Nguyen | A polymeric microgripper with integrated thermal actuators[END_REF][START_REF] Roch | Fabrication and characterisation of an SU-8 gripper actuated by a shape memory alloy thin film[END_REF][START_REF] Lorenz | SU-8: a low-cost negative resist for MEMS[END_REF][START_REF] Dellmann | Fabrication process of high aspect ratio elastic structures for piezoelectric motor applications[END_REF][START_REF] Voskerician | Biocompatibility and biofouling of MEMS drug delivery devices[END_REF][START_REF] Chronis | Electrothermally activated SU-8 microgripper for single cell manipulation in solution[END_REF].

Several microgrippers have been developed using different kind of actuating methods, by means of different physical effects, as piezoelectric, electromagnetic, electrostatic, electro-thermal, etc. Thermal actuation microgripper generally use thermal expansion by Joule heating, generated by an applied voltage. From various actuation principles that can be used in the design of microgrippers, electro-thermal actuators have some advantages as large displacements and relative low response time. The electric current passing through the thin metallic layer generates heats that deform/expand the material of the hot arms, the movement of the arms occurring.

Microgrippers manufactured of polymeric materials offer the advantage of a large displacement and a gentle handling force that can be ideal particularly for specific bioparticle manipulation, as cells. Operating temperatures below 100 0 C allow the manipulation of living cells and tissues. Polymers offer a much lower Young's modulus, and thus much lower actuation and handling forces [START_REF] Nguyen | A polymeric microgripper with integrated thermal actuators[END_REF][START_REF] Roch | Fabrication and characterisation of an SU-8 gripper actuated by a shape memory alloy thin film[END_REF]. For example, SU-8 has a Young's modulus in the range of 4.02 GPa[6] -4.95 +/-0.42 GPa [START_REF] Dellmann | Fabrication process of high aspect ratio elastic structures for piezoelectric motor applications[END_REF] and it was intensively used as structural material for microgrippers in the last years. Because of its biocompatible properties, SU-8 polymer can be used in a great variety of bio-MEMS applications [START_REF] Voskerician | Biocompatibility and biofouling of MEMS drug delivery devices[END_REF][START_REF] Chronis | Electrothermally activated SU-8 microgripper for single cell manipulation in solution[END_REF]. This paper presents two models of polymeric microgripper electro-thermo-mechanical actuated, using low actuation voltages, designed for SU-8 polymer fabrication. The electro-thermal microgrippers consist of two arms configurations 'hot and cold' and were designed and optimized, using finite element simulations [START_REF] Voicu | Method for overcoming the unwanted displacements of an electro-thermally actuated microgripper[END_REF]. Coupled electro-thermo-mechanical simulations, using COVENTORWARE tools, have been performed in order to describe the microgrippers behaviours in air and liquid as function of the applied voltage.

A comparison between different designs of the microgripper structures and between simulation results was performed in order to optimize the model. We obtained results for different environments operations where the microgripper manipulate the micro particles, as air and liquids. The optimization of the designed models was performed using finite element simulations and numerical methods in order to obtain a stable structure without any out of plane displacements that can occur.

II. DESIGN AND SIMULATIONS

A. Design configurations

The designed microgripper structures consist of some major parts: the fixed part, hot and cold underarms, arms and heaters (that we will be calling up and down parts) (see Fig. 1).

9-11 April 2008

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

The main dimensions of the microgrippers are 200 micrometers X 460 micrometers X 20.6 micrometers. The length of the all free arms is 400 micrometers, 2 micrometers is the thickness of the thermal oxide layer (SiO2), 20 micrometers the thickness of the polymer layer (SU-8), 0.3 micrometers the thickness for each of the Chromium/Gold layers (which are deposited on both sides of the SU-8 layer for the first model) (see Fig. 1).

Fig. 1 3D-meshed model of the microgripper simulated with CoventorWare program

The second microgripper model has one metallic layer with the same shape and has similar design as the first model with the difference only regarding the position of the metallic layer. So, for the second microgripper model the Chromium/Gold metallic layer with a thickness of 0.3 micrometers is deposited between two SU-8 layers, each with 10 micrometers thickness (see Fig. 2).

In operation, an electrical current will be applied to the gold element. The current will travel through the metallic layers producing the following movement: the part of the polymeric arms, which contains the metallic layers, will be heated and will expand, causing the displacements of the microgripper arms, and occurring the microgripper tips close.

The microgripper structure is initially in the open position with an opening distance of 20 micrometers. The microgripper is designed to operate through an integrated thermal element which is controlled by applying a voltage. The application of a voltage to the metallic layers of the structure will produce the closing of the microgripper and an object can be gripped. When we switch off the voltage, the arms of the structure will open, releasing the micro object.

B. Simulation results

Coupled electro-thermo-mechanical simulations, using COVENTORWARE tools, have been performed in order to describe the microgrippers behaviours in air and liquid as function of the applied voltage.

The model was meshed using hexahedral elements with 27 nodes and the numbers of volume element was optimised choosing the proper size of the mesh element.

As thermal boundary conditions, necessary to be set for simulations, the initial temperature of the whole structure and the temperature of the environment were considered to be 27 o C and the air convection coefficient was set to 20 W/m 2 K.

The material properties used in simulations are listed in Table 1. We have analysed different displacements that can occur in the behaviour of the microgripper models when it is actuated. Fig. 3 and Fig. 4 show the remaining distance between tips of the microgripper arms and how the arms of the microgripper were closing when we apply 0 -0.3 Volts on the metallic heaters. Fig. 5 illustrates the maximum values of the temperature that can occur in the models structures (in the hot arms) when 0-0.3 Volts were applied. For the second model the maximum values of the temperatures were jest a little smaller than the values for the first model. In Fig. 6 it is illustrated the distribution of the temperature in the 3D microgripper model simulated with CoventorWare tools when we apply 0.25 Volts. Fig. 7 shows the out of plane displacements-the displacement in the vertical direction -for the same voltage applied on the metallic layers. For the second microgripper model the vertical displacements values are smaller with respect to the same voltage values. This means that the second model structure in less instable; for the same voltage values the microgripper is bending down slower (for the 0.05-0.3 Volts the vertical displacement increase with less then one micrometer). Further optimised design of the model can be done varying the dimension of thickness of the polymeric layers and of the metallic layer.

Our aim was to obtain an optimised design of the microgripper model in order to reduce the out of plane displacements that can occur in the microgripper behaviour when is actuated. The Fig. 5 show that the temperatures are similar and the Fig. 7 show an optimised vertical displacement. In order to observe the gripping force that occurred between tips of the first microgripper arms and the micro object gripped we have performed new simulation. In Fig. 8 it can be observed the contact between microgripper tips and the micro object. When we apply 0.25 Volts on the heaters the microgripper model can grippe a micro object with a diameter of 5 micrometers (or 10 micrometers for the second model). For the same voltage value we have simulated the contact pressure (0.061-0.17 MegaPascals) that appear between the tips and micro object (see Fig. 8) and it is obvious that we obtain a small gripping force needed to not damage the sample. These demonstrate the gentle handling force give by the polymeric material used to fabricate the microgripper arms.

To observe the behaviour of the microgripper operating in different liquids we have simulate the first model changing the convection coefficient settings. The values of the convection coefficient for different liquids are higher and variable.

It can be observed that operation in liquid requires higher voltages than operation in air do to the higher convection coefficient encountered in liquids (see Fig. 9).

III. EXPERIMENTAL RESULTS

Preliminary experimental work was performed for optimize the model of microgripper, using SU-8 polymer as structural material, because of his high thermal expansion coefficient and the possibility to obtain high aspect ration structures.

The fabrication was based on a 4 masks process, based on surface micromachining, using 1.7 micrometers SiO 2 as sacrificial layer. The microgripper's arms were manufactured by SU-8 polymer, hard bake at 150 ºC on a hot plate, having a thickness of 20 micrometers; they can move laterally. The heating resistance and the pads were of thin Cr /Au layer. The first steps of the technological process can be seen in Fig. 10. The choice of biocompatible materials together with the low actuation voltages required and large deflection produced at relatively low temperatures makes this microgripper highly suitable for bio-manipulation experiments in air or in aqueous media.

Future work includes the realization of experiments for micro manipulation of micro objects and biological particles.

We have proposed a new model optimized for a micropgripper, electro-thermal actuated at low voltage, manufacturing using polymeric "hot and cold" arms.

Fig. 2

 2 Fig. 2 3D-detail of the second microgripper model simulated with CoventorWare program

Fig. 3 Fig. 4

 34 Fig. 3 The remaining distance between the ends of free arms of the microgripper models

Fig. 5 Fig. 6

 56 Fig. 5 Maximal values of the temperatures in the microgrippers when a voltage is applied

Fig. 7

 7 Fig. 7 Maximal out of plane displacements for the models of microgripper when a voltage is applied

Fig. 8

 8 Fig. 8 Microgripper tips in contact with a micro object grippedsimulation

Fig. 10

 10 Fig.10 SEM picture of the (a) SU-8 microgripper (b) Cr / Au resistance fabricated based on FEM model

TABLE 1 MATERIAL

 1 PROPERTIES USED IN SIMULATIONS

	Material	SU-8	SiO2	Gold
	Density
	(Kg/µm3) X 10 -15	1.2	2.15	19.3
	Young's
	Modulus (MPa)	4.95	70	57
	X 10 3
	Poisson's Ratio	0.22	0.17	0.35
	TCE (1/K) X 10 -5	5.2	0.05	1.41 (300K)
	Thermal
	conductivity	2	14	2970
	(pW/µmK) X 10 5
	Specific Heat (pJ/kgK) X 10 15	1.675	1	0.129

ACKNOWLEDGMENT This work was supported by National Romanian Program INFOSOC-Project Nr. 28/2005 and CALIST-Project Nr.6111/2005.


R F Shyu

H Yang

J.-H Lee

Micro-electroforming Metallic Bipolar Electrodes for Mini-DMFC Stacks

Micro-electroforming metallic bipolar electrodes for mini-DMFC stacks

INTRODUCTION

Global warming effect results in serious environmental problems. Fossil energy consumption generated lots of carbon dioxide and related compound is the major reason. Clean energy with high conversion efficiency and low pollution effect is human's common objective. Fuel cell is an electrical power generator, it converts chemical energy to electrical power. There are two major fuel cells including PEMFC (proton exchange membrane fuel cell) and DMFC. Both systems need a FC stack as the reaction site. The key components include proton exchange membrane, gas diffusion layer (electrode), catalyst, and bipolar plate. The combination of proton exchange membrane, catalyst and gas diffusion layers is called MEA (membrane electrode assembly). Hydrogen fuel with catalyst help reacting with oxygen generates electricity and water. Bipolar plates are responsible for hydrogen and water flow in the FC stack.

Many conductive materials are suitable for mini-DMFC bipolar plates, silicon and stainless steel ate two major items [1]. Jiang et al used (100) silicon as the substrate and etched microchannels as bipolar plates [2]. A reactive area of 1.3×1.1cm 2 MEA was applied and generated the maximum output power density 2.31 mW/cm 2 . Methanol solution (1 M volume concentration) was operated at room temperature. Though silicon bipolar plates can be fabricated for micro-DMFC stacks, the brittle problem is easy to make cracks in assembly and results in failure [1]. Stainless steel provides low cost and high strength advantages compare to silicon. The skillful metal work can realize stainless steel for commercial applications [3]. Yang et al used stainless steel (SS316) as bipolar plate material for micro-DMFC stacks. A MEA with 5 cm 2 was assembly to a single DMFC. The maximum output power was 13mW/cm 2 when using 6 vol. % methanol solution and operated at temperature 31°C [4].

Different microchannel flow patterns have direct influence on fuel cell stack performance. Serpentine flow pattern has the best result based on practical test compared to other patterns. Bipolar plates with serpentine flow pattern contributed a better CO 2 exhaust on anode side. The further study also indicated the generated output efficiency using serpentine flow pattern is suitable for micro-DMFC stacks [5]. Microchannel profile related to liquid flow pattern is the other concerned issue. Microchannel with 580μm wide for mini-DMFC has a better performance than other sizes [6]. Different depths were studied and concluded mcirochannel with 300 μm deep suitable for the optimum efficiency [7]. Based on previous studies, many researchers suggested that fluidic channel with 300 μm in depth and 500 μm in width provides an efficient flow.

The experiment was to design such pattern in the mask. UV-LIGA process was applied to fabricate such pattern for bipolar plates with low electrical resistance. Stainless steel (SS316) with 1.5 mm thick was used as the substrate. The negative photoresist (JSR THB-120N) with 300 μm thick was triple-spun coating on the substrate. UV exposure and development then generated the fluidic channel mold in photoresist. Copper electroforming then filled the mold onto the substrate. The dimensional measurement is required to confirm the geometry accuracy by using an optical 3D profile. Electrical performance measurement of mini-DMFC to verify the experimental bipolar plates is measured by a commercial MEA and test system.

ISBN: 978-2-35500-006-5 II.
EXPERIMENTAL PROCEDURE Metallic bipolar plates with different patterns for microchannel flow were studied. Previous studies showed that the optimum microchannel with width 500μm and depth 300μm provides the best efficiency. Pattern design on the mask is shown in Fig. 1. Serpentine (left diagram) and interdigital (right diagram) flow patterns were designed as shown. Desired patterns are transferred from the designed mask in the lithographic process. In this experiment, a plastic mask was fabricated using laser writing onto a PET (Polyethylene terephthalate) used for PCBs (print circuit boards). The stainless steel (SS316) substrate (Fig. 2) was then spun with a layer of negative photoresist (JSR THB-120N) 100μm thick. The spin condition was 150rpm for 15 seconds. Prebaking in a convection oven at 100°C for 3 minutes is a required procedure. This removes the excess solvent from the photoresist and produces a slightly hardened photoresist surface. Repeat the coating and prebaking process in three times resulted in multi-layer thickness 300 μm. The mask was not stuck onto the substrate.

The sample was exposed through the plastic mask using a UV mask aligner (EVG620). This aligner had soft, hard contact or proximity exposure modes with NUV (near ultra-violet) wavelength 350-450nm and lamp power range from 200-500 W. The exposure time was 180 seconds and developing time for 40 minutes. The hardbake temperature was 120°C for 3 minutes by the hot plate. Commercial cooper sulfite electrolyte was used to electroform metallic plates. The specimen with microchannel flow pattern on the stainless steel substrate was immersed into the electrolyte. Cooper atoms were deposited into resist mold from the substrate. The input current density was 5ASD and the growth rate was 1 μm/min in theory. The plating elapsed for 5 hours to obtain microchannel thickness 300μm in cooper. Sodium hydroxide solution 150 ml containing 3.8 g NaOH was heated to 60°C. Photoresist mold was detached from the substrate after 15 minutes immersion. After cleaning with DI water and spray nitrogen air, metallic cooper microchannel bipolar plates were completed. After stripping the photoresist and sputtering silver with 1.2 μm, the metallic bipolar plate with surface resistance 284 mΩ was completed. The completed metallic bipolar plates for mini-DMFC stacks are shown in Fig. 3.

III. EXPERIMENTAL RESULTS

A. Dimensional measurement

Serpentine microchannel profiles of bipolar plates were measured using a 3D surface profiler to measure their widths and depths. As shown in Fig. 4(a), the measured width is 498μm and the depth is 350μm. The three dimensional profile is also shown in Fig. 4(b). Rough surface came from large grains of copper deposition. That can be further modified in electroforming conditions such decreasing growth rate and adding some additives for grain refining. Interdigital microchannel pattern of bipolar plates was measured as shown in Fig. 5. The dimensional measurement results are close to the design size. There is about 10% dimensional error existed.

B. Mini-DMFC stack assembly

The mini-DMFC stack includes a 1.5x1.5cm 2 MEA from Micropower Tech, metallic bipolar plates from this experiment, and other fixtures as shown in Fig. 6. To understand the commercial DMFC performance using Micropower Tech's MEA, Fig. 8 shows its performance characteristics. Test equipment was the Antig's DMFC tes platform. The maximum output power density 12 mW/cm 2 with the current density 50 mA/cm 2 was measured. That's a reference to verify the experimental metallic bipolar plates since the commercial bipolar plate is made of graphite. Serpentine and interdigital microchannel patterns of metallic bipolar plates in mini-DMFC were assembly and measured their performance characteristics. Fig. 9 shows the performance characteristics of serpentine microchannel bipolar plates, its output performance is similar to the commercial product. Interdigital microchannel bipolar plates have the poor performance as shown in Fig. 10. Once again showing serpentine flow pattern is more suitable for DMFC stacks. Fig. 11 shows the functional test of mini-DMFC by using a mini-fan as the load. The output voltage 0.37 V is shown to drive the mini-fan running.

IV. CONCLUSION

Commercial MEAs with the fabricated bipolar plates were assembly into a 2x2 cm 2 stack. As found, the output power mainly dominated by MEA. MEA is the most important factor effect on the output power, though fluidic channel types on bipolar plates have a little effect on the stack performance. The completed mini-DMFC stack has the maximum power 9.33mW/cm 2 was measured. The power density is similar to macro-FC stacks. It proves the proposed micro-electroforming bipolar plates feasible to mini-DMFC stacks. The further applications trend to apply the mini-DMFC stacks for portable electronic products.

Fig. 1 .

 1 Fig. 1. Schematic diagrams of microchannel flow pattern on the mask.

Fig. 2 .Fig. 3 .

 23 Fig. 2. Stainless steel (SS316) substrates as the plating base material, each piece is 2x2 cm 2 .

Fig. 4 .Fig. 5 .

 45 Fig. 5. Measurement results of interdigital microchannel pattern bipolar plate; (a) cross-sectional view and (b) 3D morphology.

Fig. 6 .

 6 Fig. 6. Illustration of mini-DMFC stack parts.

Fig. 7 .

 7 Fig. 7. Comparison photos of conventional DMFC stack and mini-DMFC stack.

Fig. 9 .

 9 Fig. 9. Mini DMFC performance characteristics using serpentine microchannel pattern bipolar plates; (a) I-V curve and (b) output power.

Fig. 11 .

 11 Fig. 11. Functional test of mini DMFC stack with output voltage 0.37 V.

ACKNOWLEDGMENT This research is supported by the National Science Council of Taiwan under the grand number NSC96-2221-E-005-069-MY3.


P Martins
email: paolo.martins@insa-lyon.fr

C Malhaire

S Brida
email: sbrida@auxitrol.com

Damien Barbier

On the determination of Poisson's ratio of stressed monolayer and bilayer submicron thick films

In this paper, the bulge test is used to determine the mechanical properties of very thin dielectric membranes. Commonly, this experimental method permits to determine the residual stress (σ 0 ) and biaxial Young's modulus (E/(1υ)). Associating square and rectangular membranes with different length to width ratios, the Poisson's ratio (υ) can also be determined. LPCVD Si 3 N 4 monolayer and Si 3 N 4 /SiO 2 bilayer membranes, with thicknesses down to 100 nm, have been characterized giving results in agreement with literature for Si 3 N 4 , E = 212 ± 14 GPa, σ 0 = 420 ± 8 and υ = 0.29.

INTRODUCTION

The development of Micro Electro Mechanical Systems (MEMS) has become an economic stake since 80's and more recently the Nano Electro Mechanical Systems (NEMS) began to be developed with the downscaling trend. However, with the downscaling, accurate measurement of mechanical properties becomes a hot challenge especially since these properties may depend on the fabrication process. This may have consequences on MEMS performances and reliability [START_REF] Martins | Design of bossed silicon membranes for high sensitivity microphone applications[END_REF][START_REF] Zhu | Reliability of capacitive switches at high and low temperatures[END_REF][START_REF] Lixian | Numerical analysis of theoriacl model of the RF MEMS switches[END_REF]. Furthermore, architectures are more and more complex such as multilayers which make the determination of mechanical properties more difficult for each constituent material.

So far, no mechanical method exists for the simultaneous determination of the three main mechanical parameters: Young's modulus E, residual stress σ 0 and Poisson's ratio υ, except the well-known "Bulge test" method. Indeed, the bulge test is commonly used to determine the residual stress σ 0 and the biaxial Young's modulus E/(1-υ) on square or circular membranes. Some authors have also shown that the association of membranes with different shapes permits the determination of the Poisson's ratio [START_REF] Tabata | Mechanical property measurements of thin films using load-deflection of composite rectangular membranes[END_REF][START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF]. However, few studies deal with the determination of the Poisson's ratio of very thin film membranes with significant results [START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF]. Indeed, J. S. Mitchell et al. [START_REF] Mitchell | Examination of bulge test for determining residual stress, Young's modulus, and Poisson's ratio of 3C-SiC thin films[END_REF] relate the difficulties in determining this ratio because the bulge test method is very sensitive to the geometrical errors.

In this study, an attempt was made to determine E, σ 0 and υ by means of bulge test on very thin (~ 100 nm) Si 3 N 4 square and rectangular membranes with different length to width ratio (1 < b/a < 12). Few similar studies have been made on dielectric membranes with thicknesses down to 100 nm.

In this work, we have also assessed the mixture law as a general rule to extract the Young's modulus, residual stress and Poisson's ratio of each film of submicron thick Si 3 N 4 /SiO 2 bilayers.

II. BACKGROUND

The bulge test consists in applying a pressure P on a membrane and in measuring its maximal deflection h at its center (Fig. 1). ( )

3 3 0 1 2 4 4 2 t E t t E P C (a,b) h h f( ,a,b) h a a a1 12 1 σ = + + υ -υ α -υ (1)
t represents the membrane thickness, 2a and 2b are the membrane width and length, respectively. C 1 (b/a) and α are coefficients that depend on the membrane shape and f (υ, b/a) depends also on the membrane shape and on the Poisson's ratio. We can note that in the case of large deflections (h/t >> 1), the second term in Eq. 1 (depending on α) can be neglected.

Lots of studies have been made to optimize the coefficient values as a function of the membrane shape and in order to take into account the particular clamping conditions of micromachined membranes [START_REF] Tabata | Mechanical property measurements of thin films using load-deflection of composite rectangular membranes[END_REF][START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF] This experimental method is efficient to determine the residual stress σ 0 and the biaxial Young's modulus E/(1-υ) on thin film. Eq. 1 shows that E and υ are highly correlated and to find one of these parameters, the other must be assumed. To determine υ independently of E, experiments performed on square and rectangular membranes must be associated [START_REF] Tabata | Mechanical property measurements of thin films using load-deflection of composite rectangular membranes[END_REF][START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF][START_REF] Mitchell | Examination of bulge test for determining residual stress, Young's modulus, and Poisson's ratio of 3C-SiC thin films[END_REF]. Indeed, it is possible to compare the ratio of the cubic coefficients in Eq. 1 to the ratio of the function f (υ, b/a) for square and rectangular membranes (Eq. 2).

  υ =   υ   (2)
Moreover, the experimental results of J. J. Vlassak et al. [START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF], for rectangular membranes with b/a > 4, showed that membranes can be considered as infinite along the length and the deflection is independent to the aspect ratio.

The other interest of this study is to apply this method in the case of very thin multilayer films by using the simple formula of the mixture law (Eq. 3) [ [START_REF] Shojaei | Comparison of mechanical properties of TiN thin films using nanoindentation and bulge test[END_REF][START_REF] Lau | Processing and characterisation of micromachined actuators based on proton-irradiated P(VDF-TrFE) copolymer[END_REF] which can be applied for E, σ 0 and υ.

1 2 n composite 1 2 n total total total t t t M M M M t t t = + + ⋅⋅⋅ + (3)
M composite represents either the biaxial modulus or the residual stress of the composite membrane with n layers. t 1 , t 2 … t n are the thicknesses of each component layer, t total is the multilayer thickness and M 1 , M 2 … M n represent either E/(1-υ) or σ 0 of each layer.

III. SAMPLE PREPARATION AND EXPERIMENTAL SETUP

A. Sample preparation

Dielectric membranes have been fabricated on <100> ptype, double-side polished, 100 mm silicon substrates using a standard micromachining process. Silicon nitride films have been deposited at 835 °C by LPCVD on thermally oxidized silicon substrates (Fig. 2). Two wafers issued from the same fabrication process have been processed: Si 3 N 4 film (first wafer) with a thickness of 104 nm and bilayer Si 3 N 4 /SiO 2 film (second wafer) with a thickness of 188 nm (t(Si 3 N 4 ) = 90 nm and t(SiO 2 ) = 98 nm). Free standing membranes have been obtained through silicon anisotropic etching in a KOH solution (Fig. 2). Several samples have been obtained with different shapes (square and rectangular). The resulting sample characteristics are summarized in Tab. II. Pressure-deflection measurements (P (h)) have been performed using a WYCO NT1100 white-light interferometer microscope (Fig. 3). Pressures ranging from 10 mbar to 1 bar (depending on geometry) have been applied. Wax has been used to fix our samples on a sample holder. In this study, according to the simulations, the analytical model proposed by J. J. Vlassak et al. is still valid for our 200 nm thick or less membranes and it was used to determine the mechanical properties of the Si 3 N 4 monolayer and the Si 3 N 4 /SiO 2 bilayer self-standing films.

B. Experimental results

In the case of 2M, 3M and 4M samples, with the same geometrical parameters (Tab. II), experimental results have been obtained with a good reproducibility. Moreover, for all membranes, no hysteresis phenomenon has been observed during load and unload cycles showing a linear behavior of the membranes despite of the large induced deflections (> 90 µm).

A P/h as a function of h² normalized representation of the pressure-displacement results can be made in order to extract the y-intercept and the slope of the curves (see fig. 5 and6). A Poisson's ratio of 0.3 for LPCVD Si 3 N 4 was assumed according to the literature values [16] to calculate the Young's modulus E and the residual stress σ 0 for both Si 3 N 4 and Si 3 N 4 /SiO 2 membranes. These values are summarized in Tab. II for each sample.

For the monolayer Si 3 N 4 membranes, a mean Young's modulus value of 212 ± 14 GPa and a mean residual stress of 420 ± 8 MPa have been found. For the composite Si 3 N 4 /SiO 2 bilayer membranes, the results were 147 ± 8 GPa and 107 ± 2 MPa for the Young's modulus and the residual stress, respectively. In the case of Si 3 N 4 membranes, a Poisson's ratio scattering is observed between 0.22 and 0.29 for an expected value between 0.25 and 0.3 (LPCVD Si 3 N 4 ). As regards the Si 3 N 4 /SiO 2 bilayer membranes, the composite Poisson's ratio results are more scattered and higher than for the Si 3 N 4 monolayer (0.23 < υ < 0.41) whereas values lower than those of Si 3 N 4 monolayer membranes were expected (except for 5B/1B samples). Even if the Poisson's ratio obtained for Si 3 N 4 monolayers are close to the expected value, it is obvious that the determination of an accurate Poisson's ratio is very difficult, especially for bilayer membranes. Moreover, the high uncertainties in Tab. III are calculated from the lateral dimensions uncertainties showing the importance to know accurately these geometrical parameters.

Membranes that were issued from the same wafer should have the same mechanical properties but differences lower than 10 % were observed on Young's modulus results (Tab. II). These differences could come from the presence of a film thickness gradient or a stress gradient across each wafer. In this study, a mean thickness value was assumed for each wafer to calculate the mechanical properties. This may also explain the scattering on the Poisson's ratio values calculated for Si 3 N 4 . Moreover, sometimes we have observed underetching profiles on some Si 3 N 4 /SiO 2 membranes when lateral dimensions were lower than 1 mm (see Fig. 7). Even for large length/width ratio, this slightly changes the clamping conditions of the membranes and may etching influence the experimental results.

The model used to determine the Poisson's ratio is also critical. For example, for mono or multilayer membranes, the model of E. Bonnotte However, when two membranes, square and rectangular, lead to very close values for the Young's modulus (we can assume that the film thickness is similar for the two samples and that the clamping conditions are good), then a precise determination of the Poisson's ratio can be made. Indeed, the 1M (square) and 5M (rectangular) Si 3 N 4 membranes, lead to very close Young's modulus values and the calculated Poisson's ratio (Tab. II and III), in close agreement with literature values for LPCVD Si 3 N 4 [START_REF] Vlassak | A new bulge test technique for the determination of the Young's modulus and the Poisson's ratio of the thin film[END_REF][START_REF] Huang | Mechanical properties of polymer thin film measured by the bulge test[END_REF]16]. The same observation can be made for the 1B and 5B Si 3 N 4 /SiO 2 membranes giving a composite Poisson's ratio of 0.23.

D. Application of mixture law

Using the mixture law (Eq. 3), an attempt was made in order to calculate the mechanical properties of thermal SiO 2 with our experimental results.

With E (Si 3 N 4 ) ≈ 212 GPa and E (Si 3 N 4 /SiO 2 ) ≈ 147 GPa, we obtained E (SiO 2 ) ≈ 87 GPa. With σ 0 (Si 3 N 4 ) ≈ 420 MPa and σ 0 (Si 3 N 4 /SiO 2 ) ≈ 107 MPa, we obtained a compressive stress σ 0 (SiO 2 ) ≈ -180 MPa. Finally, with υ (Si 3 N 4 ) ≈ 0.29 and υ (Si 3 N 4 /SiO 2 ) ≈ 0.23, we obtained υ (SiO 2 ) ≈ 0.17.

The mechanical properties calculated on thermal SiO 2 are in close agreement with the literature values [START_REF] Józwik | Optomechanical characterization of compressively prestressed silicon oxynitride films deposited by plasma-enhanced Chemical vapor deposition on silicon membranes[END_REF][START_REF] Thévenet | 17ième Congrès Français de Mécanique[END_REF][START_REF] Zohni | Investigation thin film stresses in stacked silicon dioxide / silicon nitride structures and quantifying their effects on frequency response[END_REF].

V. CONCLUSION

These results show that the determination of E, σ 0 and υ by means of the bulge test method remains possible even for deep submicron monolayer or multilayer thin films. Large deflections can be imposed to the membranes without any plastic deformation, which simplifies the associated mechanical model. Finite Element simulations show that the coefficient values found by J. J. Vlassak et al. were well suited for our studied samples. But the accuracy of the results depends strongly on the geometrical parameters especially the thickness of the membranes. Young's modulus values and residual stress have been determined with accuracy better than 10 %. But the accuracy on the Poisson's ratio is about 20% in the best case. This highlights the difference between theory and experience because achieving well-controled free-standing submicron thick films is not trivial. Finally, a simple mixture law has given promising results on standard materials.

Fig. 1 .

 1 Fig. 1. Bulge test principle Mechanical properties like Young's modulus E, the residual stress σ 0 and the Poisson's ratio υ can be determine from [8-10]:

  [10-13]. Tab. I presents different values of C 1 (b/a), f (υ, b/a)) and α found in the literature as a function of the membrane shape. In this study, the coefficients C 1 (b/a) and f (υ, b/a) have been recalculated with Finite Element simulations (FE) using ANSYS software in order to verify their validity for very thin films membranes (~ 100 nm) and to compare with the literature values (Sees paragraph IV).

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

Fig. 2 .

 2 Fig. 2. Membranes fabrication: process steps

Fig. 3 .Fig. 4 .

 34 Fig. 3. Optical interferometer setup

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

Fig. 5 .Fig. 6 .

 56 Fig. 5. Normalized pressure-displacement (P = f (h²)) curves of Si3N4 membranes

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

  et al. leads to extreme Poisson's ratio values (>0.45) compared to that of J. J. Vlassak et al. This last model is the most appropriate to calculate υ.

Fig. 7 .

 7 Fig. 7. Si3N4/SiO2 Back-side membranes after anisotropic wet etching (pictures obtained using interferometer microscope)

TABLE II RESULTS

 II OBTAINED FOR EACH MEMBRANE USING υ = 0.3 Tab. III shows the different Poisson's ratio values obtained for different pairs of samples and from the analytical model proposed by Vlassak et al. and from Eq. 2.

			n°	2a (mm) 2b (mm)	b/a	σ0 (MPa) E (GPa)
			1M	3.104	3.104	1	439±27 210±16
	Si3N4	(t = 104 nm)	2M 3M 4M	2.131 2.131 2.14	2.131 2.131 2.14	1 1 1	400±27 217±19 409±25 214±16 429±29 211±18
			5M	1.138	2.131	1.9	414±34 219±26
			1B	1.89	1.89	1	104±8 150±14
	Si3N4/SiO2	(t = 188 nm)	2B 3B 4B	0.662 0.750 1.39	0.662 0.750 7.80	1 1 5.6	113±9 153±17 100±8 156±17 103±8 139±15
			5B	0.27	3.28	12.1	115±10 145±16
	C. Determination of the Poisson's ratio


Kyu Jeung

Hyun Chang

Jingli Lim

Seung Wook Yuan

Park

Wafer Level Package for Image Sensor Module

A new ISM (image sensor module) WLP (wafer level package) for reflow process is designed, fabricated and tested. The ISM WLP is composed of polymer bonding layer, glass cap wafer for particle free process and CIS (CMOS Image Sensor) chip wafer which has micro via hole interconnection. During the last decades, WLP is highlighted as the next generation ISM Package method for many advantages like high yield (particle free process), small form factor (3D interconnection), low assembly cost and so on. Nevertheless these benefits, there are some problems like micro via hole fabrication, low temperature insulation process (inside hole), bottom side oxide etching, warpage control according to wafer level bonding using different material, and whole process temperature limitation for micro lens damage. Among various fabrication methods for ISM package, COB (Chip on board), COF (Chip on film), and L, T contact WLP from ShellCase are generally used. In case of COB and COF package, it has difficulty in particle control during assembly process. In case of ShellCase type WLP has very complicated fabrication process. Additionally, most of above package has disadvantage in size point of view. Through suggested ISM WLP using through interconnection via, wafer level fabrication & packaging technology is realized. It can not only solve problems of conventional packaging structures but also tremendously reduce the manufacturing & assembly cost (include time) of ISM package and realize real chip scale package. Based on sensor size, 3.67 X 3.42 X 0.39 (H) mm3 WLP is designed. During the parametric study using commercial 3-D simulation programs, silicon thickness, polymer bonding layer thickness, and glass thickness were chose the effective factor. And considering the optical and electrical analysis, we decide the parameter: silicon thickness is 0.1mm, polymer bonding layer thickness is 0.04mm, and glass thickness is 0.25mm. The fabrication process is composed bonding layer patterning, wafer bonding, thinning, via etching, passivation layer deposition, bottom oxide opening, metal plating, bottom electrode patterning, solder ball formation, and dicing. A new concept of ISM WLP has been founded to be suitable structure for low cost, small form factor application. I.

INTRODUCTION

Many researcher forecast mobile phone market will be increase 10 % annual average growth between 2007 to 2010. These predictions are based on two fundamental trends. First, the growth of BRICs (BRICs means Brazil, Russia, Indonesia, and China) market has been most visible, in worldwide phone sales growing. This trend requests cost effective mobile phone, including ISM (low-end market). Second, the move to 3G will be make new high end market. This trend requests additional low grade ISM. According to these reasons, there were needs which is low cost, small size, low grade (like VGA) ISM. To satisfy these needs, there were many packages are developed. These include COB (Chip on board), COF (chip on film), and WLCSP (like shallcase type). [START_REF] Sengupta | Electronics Manufacturing Technology Symposium[END_REF][START_REF] Viswanadam | Electronic Packaging Technology Conference[END_REF][START_REF] Darveaux | Thermal and Thermomechanical Phenomena in Electronic Systems[END_REF] In spite of these efforts, there are many limitations in cost and size point of view. In this paper we have developed a fabrication technology of WLP for ISM, using the polymer based wafer bonding, through via hole interconnections from the backside of the Si substrate to the electrical pad. This package consists of a glass cap wafer for protecting an image sensing area from particle during fabrication process, adhesive bonding layer for low temperature wafer bonding, 3-D interconnections for real chip scale package, metal re-routing and solder bumps on the backside for reflow process. All processes of the packaging were done at wafer level. This paper describes the optimal package design, fabrication process, and evaluation results for ISM. Using this package, we can satisfy most of market requirements like user assembly cost reduction (reflowable package), chip scale package (through via interconnection method), increasing manufacturing efficiency (particle free process). II.

DESIGN

A.

Basic Concept of Package.

Fig. 1 shows the basic concept of WLP for ISM. The package is consisted of glass cap wafer, polymer bonding layer, and CIS wafer which has through interconnection via.

The glass cap wafer has a function which is not only window for image processing, but also protection layer contamination during fabrication process. Adhesive polymer bonding layer is chosen for low temperature bonding. Because micro lens on active area are maintained under 200 in all fabrication steps. WLP must be cost effective and ultra small package. Under these constructions, we chose proper methods and optimized each process steps to achieve a low-cost and high-density WLP. Through via interconnection is chosen for chip scale package, ultra small package, and solder ball is attached for reflowable package. This reflowable package can eliminate additional interconnection cost between chips to board.

B.

Simulation.

We predict warpage, stress distribution and shear stress on the proposed package structure to figure out critical area on the package where can cause failure. Warpage of package closes zero due to the similar CTE values between top wafer and bottom wafer. Bonding material is more critical factor due to the high CTE. CTE mismatch can be source of failure as delamination on the bonded area. Device wafer thickness slightly affects warpage and stresses on WLP ISM. Shear stress increase by bits due to the increase of device wafer thickness. We can fine these results in figure 3

A. Process Overview

The schematic fabrication process of the Reflowable ISM WLP is shown in Fig. 2. A 4-inch sized CMOS Image Sensor wafer was prepared with thickness of 525um. And a 4-inch pyrex glass wafer with thickness of 250um was used as the protective cap wafer for the package. The cap glass wafer was bonded with the image sensor wafer by adhesive bonding process. This adhesive bonding process is chosen because of low bonding temperature. A DFR (Dry Film Resist) layer was used as bonding material because of low cost and convenience of use. After the bonded wafer was thinned to the required thickness by conventional CMP process, backside vias were formed through the silicon wafer towards the Al pads of the image sensor. And an insulation film of SiO2 was deposited as the passivation layer on the wafer surface, as well as inside the vias. The passivation layer on surface of the Al pads was removed by etching process, and then electrical connection was established by forming conductive leads from Al pads to the backside of the wafer through the backside vias. Finally solder balls were formed on the backside of the wafer for reflow process.

IV. PROCESS CONDITION AND EXPERIMENT RESULTS

For the packaging technology of image sensor, low-temperature processes are preferred for the protection of the sensor device and micro-lens. In this paper, all the process steps were performed at temperature lower than 150 except the solder ball reflow process. Key processes and experiment results are explained below in detail.

A. Bonding Process

For the application of ISM WLP, choosing the cover layer whose role is protection of the image sensing area may be restricted to the glass wafer because of several requirements like transparency and durability. In this fabrication, polymer bonding method is adopted as the suitable technique to bond the glass wafer to the top side of the CIS wafer. The Vacrel, dry film resister (DFR) of DuPont, photosensitive adhesive material, is used as polymer bonding material.

A brief description of bonding process is followed. First of all, the DFR is laminated and patterned on the glass wafer, and then glass wafer with DFR pattern is bonded to CIS wafer through the high pressure and temperature. Due to the weakness in high temperature of CIS wafer which contain temperature sensitive material, the maximum bonding temperature is decided inherently as shown in figure 3 about 165 . Figure 4 represents the result of polymer bonding. In case of bonding process, the strength of bonded area can be the indicator whether the quality of bonding is good or not. In this project, the shear strength is adopted as the index for deciding the quality of bonding results and 'Dage series 4000' equipment is used. As shown in Table 1, the shear strength of

B.

Backside via etching.

DRIE process was used to form vertical vias from the wafer backside surface to the insulation layer at the bottom of Al pads. Via diameter was 50um and aspect ratio was optimized as ~2.0 for an optimum balance of process capability requirement between via etch process and passivation layer deposition process. Process result is shown in figure 5.

In the DRIE process for the backside via, small scallop and notch were needed for sequent passivation process. So The DRIE process was optimized to get good sidewall surface smoothness, i.e. scallop, small notch effect on via bottom corner, and acceptable etch rate as well. Both HF(high frequency) and LF(low frequency) processes were used in the via etch step to get benefit of both good etch rate and notch-free process. Optimized scallop (<200nm) and notch (<1um) were achieved in this step to meet the process requirement of the passivation layer deposition process.

C. Backside via passivation.

A SiO2 passivation layer was deposited on wafer backside and inside via holes by low-temperature PEVCD process at 100 . The process was optimized for good step coverage, sufficient passivation thickness and acceptable deposition rate. Process result is shown in figure 6. A 1um-thick passivation layer at via bottom and sidewall was achieved to provide sufficient passivation effect. Bottom electrical connection process Among various interconnection methods, via interconnection technology is selected to get diverse advantages like reducing the chip size. Generally in case of via interconnection method, metal filling by copper electroplating and copper polishing for planarization are used. It is very difficult to control the quantity of overburden of copper electroplating. Therefore for the subsequent fabrication process, removing the risen part of copper column is widely accepted method. However, during the copper polishing, it is possible to attack the passivation layer around the via hole and surface and then leakage current may be happened through the damaged passivation area. In this project, that possibility is basically eliminated by deleting the polishing process. The peculiarities of fabrication compared to the previous fabrication are first, deposited metal layer can not be used only seed layer but also redistributed line layer. Second, by using dry film resister for redistributed line patterning, copper polishing process can be eliminated. There are several advantages by using this fabrication process like the reducing the second metal deposition process, removing the possibility of happening of leakage current and so on. Figure 7 shows the schematic view of the patterning process of redistributed line with the overburden of copper filling. Figure 8 represents the SEM image of the result of copper filling process. . SEM image of metal filling using copper electroplating method.

V. CONCLUSION

A new concept, wafer level package for image sensor module which is composed Pyrex glass, polymer bonding layer and CIS wafer have been found to be suitable structure for the low cost reflowable package. We obtained optimal process parameter under constraint, like temperature. This package is realized with simple wafer level process technology. It is fabricated with small size (chip scale, 3.67 X 3.42 mm 2 ), thickness (400um), and low cost. It is suitable for mass production.

The presented package can find applications where a compact overall device size is needed while keeping low temperature process, such as a image

Fig. 1 .

 1 Fig. 1. Schematic Diagram of WLP for ISM.

Fig. 2 .

 2 Fig. 2. Geometry structures of the Reflowable ISM WLP.

Fig. 3 .

 3 Fig. 3. Simulation results at device wafer thickness 100um case.

Fig. 4 .

 4 Fig. 4. Schematic fabrication process of the Reflowable ISM WLP. 1. DFR patterning. 2. Adhesive bonding. 3. Wafer thinning.4. via etching. 5. passivation layer deposition. 6. DFR lamination. 7. passivation layer etch. 8. Seed layer deposition. 9. Mold for electroplating. 10. via filling by electroplating. 11. Seed layer etch. 12. Solder ball formation

  edge portion are different because of the non uniformity of bonding equipment. Even though there is differentiation in shear strength between center portion and edge portion, the shear strength of the weakest area, center portion, is around 26 MPa.

Fig. 5

 5 Fig. 5 the pressure and temperature profiles for polymer bonding in the ISM WLP process.

  Fig.7. SEM inspection of the cross-sectional view of via.

9- 11

 11 Fig. 8. SEM inspection of passivation layer deposition

Fig. 9

 9 Fig.9 Schematic view of the patterning process of redistributed line with the overburden of copper filling.

Fig. 10

 10 Fig.10. SEM image of metal filling using copper electroplating method.

TABLE I

 I

		Material property
	Material Name	Young's Modulus (GPa)	CTE (ppm/K)	Poisson's ratio	Tg (ºC)
	Pyrex.7740	63	3.25	0.2
	Adhesive film	3.3 (25 ºC)	69 88	0.34	74
	Silicon	130	2.6	0.27

TABLE II Parametric

 II study results.

	Device Wafer Thickness	Warpage	Max. Mises Stress	Max. Shear Stress
	75um	2.49 um	111.8 Mpa	111.8 Mpa
	100um	2.84 um	112.2 Mpa	112.2 Mpa
	125um	2.58 um	112.2 Mpa	112.2 Mpa
	150um	2.3 um	112.1Mpa	112.1Mpa

Glass Cap Wafer ; Particle free process Adhesive ; Wafer Level Bonding ; Low Temp. Bonding 3D Interconnection ; Chip Scale Package Solder Ball ; Reflowable Package

	9-11 April 2008
	III. FABRICATION
	Cap Wafer
	Bonding Material
	Bottom Wafer
	©EDA Publishing/DTIP 2008	ISBN: 978-2-35500-006-5

  1. DFR patterning. 2. Adhesive bonding. 3. Wafer thinning. 4. via etching. 5. passivation layer deposition. 6. DFR lamination. 7. passivation layer etch. 8. Seed layer deposition. 9. Mold for electroplating. 10. via filling by electroplating. 11. Seed layer etch. 12. Solder ball formation

9-11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5 sensor module, bio chips, micro valve, optical switch and display.


Chun-Ping Jen

Numerical Investigation of Laser-Assisted Nanoimprinting on a Copper Substrate from a Perspective of Heat Transfer Analysis

INTRODUCTION

Nano-imprinting lithography (NIL) has been developed over a decade [START_REF] Chou | Imprint Lithography with 25-nanometer Resolution[END_REF][START_REF] Guo | Nanoscale Silicon Field Effect Transistors Fabricated Using Imprint Lithography[END_REF][START_REF] Chou | Imprint Lithography with Sub-10 nm Feature Size and High Throughput[END_REF][START_REF] Wu | Room-temperature Si Single-electron Memory Fabricated by Nanoimprint Lithography[END_REF][START_REF] Zhang | Fabrication of 60-nm Transistors on 4-in. Wafer Using Nanoimprint at All Lithography Levels[END_REF] and is now a promising method for nano-patterning and nano-fabrication. The basic concept of conventional nano-imprinting includes a mold, an etching resist layer and a sample substrate. The mold has some nano-scale features on the surface fabricated by either E-beam lithography or focus ion bean techniques. The thermo-plastic polymer such as poly-methylmethacrylate (PMMA) is usually used as the resist layer. By heating until above its glass transition temperature (Tg), the mold can impinge into the resist layer and form a pattern. Following by reaction ion etching, the nano-pattern is transformed to the resist layer and the substrate. It is sequent followed by standard lithography processes to achieve nano-structures on the substrate surface. However, the heating mechanism of conventional nanoimprinting is slow and usually causes misalignment due to the different thermal expansion between the mold and the substrate.

Recently, laser-assisted nanoimprinting lithography (LAN) method illustrated in Figure 1a was proposed [START_REF] Xia | Ultrafast Patterning of Nanostructures in Polymers Using Laser Assisted Nanoimprint Lithography[END_REF]. This new imprinting method shares the same concepts discussed above and combines the advantages of the traditional nano-imprinting and laser-assisted direct imprint [START_REF] Wu | Room-temperature Si Single-electron Memory Fabricated by Nanoimprint Lithography[END_REF]. The LAN process depicted schematically in Figure 1a utilizes a laser pulse to irradiate through a quartz mold and melt a thin polymer film (~200 nm) on the substrate. Upon the thin polymer film is melted, the pre-loaded mold is imprinted into the polymer resist layer. This LAN method has several obvious advantages over the thermal-based nano-imprinting technique in terms of shortening the processing time and reducing the heating of the substrate and in this way the misalignment due to the thermal expansion mismatch of the mold and substrate can be avoided [START_REF] Xia | Ultrafast Patterning of Nanostructures in Polymers Using Laser Assisted Nanoimprint Lithography[END_REF]. This method shows a great potential for future nano-patterning and fabrication of nano-structures. In the present study, the novel concept of that copper was adopted as the substrate instead of silicon, which is conventionally used, was proposed. The micro/nano structures on the copper substrate could be fabricated using chemical/electrochemical etching or electroforming; following by the patterns have been transferred onto the substrate using LAN process. Alternatives of the substrate materials could lead versatile applications in micro/nano-fabrication. To demonstrate the feasibility of this concept numerically, this study introduced optical multiple reflection theory to perform both analytical and numerical modeling during the process and to predict the thermal response theoretically.

II. THEORETICAL MODELING

During the heating process of the substrate by the pulsed laser, one-dimensional transient heat-diffusion equation can be employed herein.

The physical domain for the heat-transfer system considered here is illustrated in Figure 1b. Assuming that the laser spot is larger than the mold and all materials involved are homogeneous; therefore this system can be simplified as one-dimensional. The source term of heat-generation in the substrate originating from the energy absorption of a laser pulse can be described as:

) ( ) exp( ) , ( t I x A t x S S β β - = (1)
where x and t are the spatial variable and time, respectively. A S is the absorption of the incident light, β is the optical absorption coefficient which is related to the imaginary part of refractive index, κ , and the laser wavelength, λ, by

λ πκ β 4 = [8]. I(t)
is the power density function of the incident laser pulse (power per unit area, W/m 2 ). A Gaussian-shaped laser pulse is assumed in the present study and the power density function can be expressed as: τ is chosen as the triple of the standard deviation (σ) and the irradiating time is the double of max τ . For the LAN process, the incident laser propagates through four materials: air (vacuum); quartz mold (fused silica); thin film (polymer) and copper substrate. The optical absorbance in the polymer thin film should be calculated considering optical multiple interferences. Hence, the heat generation in the thin film and the intensity of the excimer laser passing into the substrate can be obtained. The calculation matrix of electric vector based on the fundamental equations of optical characteristic matrix method [START_REF] Macleod | Thin-film Optical Filters[END_REF] can be expressed as followings:

⎟ ⎟ ⎠ ⎞ ⎜ ⎜ ⎝ ⎛ - - = 2 2 max max 2 ) ( exp 2 ) ( σ τ σ π t I t I ( 2
⎥ ⎦ ⎤ ⎢ ⎣ ⎡ ⎥ ⎦ ⎤ ⎢ ⎣ ⎡ = ⎥ ⎥ ⎦ ⎤ ⎢ ⎢ ⎣ ⎡ - + - - - - + - - - - - m m i i m i m i m m m E E e e r e r e t E E m m m m 1 1 1 1 1 1 1 δ δ δ δ (3) m m m m n n n t + = - - 1 1 2 (4) m m m m m n n n n r + - = - - 1 1 (5) λ π δ m m m d n 2 = (6)
where and are the forward and backward electric waves at the interface of the m th layer, respectively. and are the complex refractive index and the thickness of the m th layer, respectively. The calculation is for normal incident radiation and assumes the materials are optically homogeneous. After solving the above matrix, the electric vectors in each layer are obtained. Therefore, the energy absorbance in the m th layer can be calculated by the following equation:

+ m E A - m E m n m d m ⎟ ⎠ ⎞ ⎜ ⎝ ⎛ ⋅ - ⋅ + ⎟ ⎠ ⎞ ⎜ ⎝ ⎛ ⋅ - ⋅ = -- - - + + - - 2 1 1 2 2 2 1 1 ) Re( ) Re( ) Re( ) Re( m m m m m m m m m E n E n E n E n A (7)
In the present investigation, the transmission at air(vacuum)-quartz mold interface is calculated using simple definition of the optical transmissivity between two materials, i.e. The incident laser is then considered as propagating through quartz mold, polymer resist layer and substrate. The energy absorbed in the polymer resist layer and the substrate is calculated by the model of the optical multiple interference discussed above. The energy absorbance is assumed as uniform distributed in the polymer resist layer. Furthermore, the incident energy entering the substrate is absorbed and distributed as the heat source formulated in Eq. ( 1). First of all, convective and radiative heat losses at the surface of the substrate and heat absorption by the quartz mold are neglected and the properties of all materials are assumed to be constant (temperature-independent). The fused quartz mold and the polymer resist layer are treated as highly transparent to laser wavelength without energy absorption. Furthermore, the heat conduction of the quartz and polymer are so small that the surface of substrate can be treated as a thermal-insulated boundary. As a result of these assumptions, the mold and polymer resist layer could be reasonably excluded from the heat transfer system and the problem could be simplified as a semi-infinite slab ( ∞ < < x 0 ) as shown in Figure 1c. The superposing effects of a large number of instantaneous sources (one-shot) are adopted to obtain the time-and spatial-dependent temperature profile analytically. First, consider the x-direction through an infinite domain with constant properties. The equation for time-dependent heat conduction in the x-direction is expressed as following:

2 2 x t ∂ ∂ = ∂ ∂ θ α θ (8) where i T t x T t x - = ) , ( ) , (
θ is the excess temperature relative to the initial (far-field) temperature of the domain

( ). i T C k ρ α = " q
is the thermal diffusivity of the material. ρ denotes the density, C represents the specific heat, and k is the material's thermal conductivity. The excess temperature distribution of the infinite domain in which an instantaneous plane source, (J/m 2 ), is released once at t=0 is [START_REF] Bejan | Heat Transfer[END_REF] :

⎟ ⎟ ⎠ ⎞ ⎜ ⎜ ⎝ ⎛ - = t x t C q t x 4 exp ) ( 2 ) , ( 2 2 1 " α πα ρ θ (9)
The initial and boundary conditions in the modeling domain can be expressed as:

i T t x T = = ) 0 , ( ; i T t x T = ∞ = ) , ( ;
0 0 = ∂ ∂ = x x T ( 10
)
where is the initial temperature of the substrate. Since the gradient of temperature at x=0 equals to zero, the calculated domain can be extend to an infinite medium

( i T ∞ < < -∞ x
) which is symmetric at x=0. According to the concept of the superposition procedure, the medium which is heated by the heat-generation sources, , the time-dependent temperature profile is solved analytically and expressed as:

) , ( t x S [ ] ∫ ∫ ∞ ∞ ⎥ ⎥ ⎦ ⎤ ⎢ ⎢ ⎣ ⎡ - - - - = p t d d t x t C S t x 0 - 2 2 1 ) ( 4
) ( exp ) ( 2 ) , ( ) , ( τ ξ τ α ξ τ πα ρ τ ξ θ (11)
where is the irradiating time of a laser pulse. Thus, the temperature inside the substrate is rewritten as following:

p t [ ] i t T d d t x t C S t x T p + ⎥ ⎥ ⎦ ⎤ ⎢ ⎢ ⎣ ⎡ - - - - = ∫ ∫ ∞ ∞ 0 - 2 2 1 ) ( 4 ) ( exp ) ( 2 ) , ( ) , ( τ ξ τ α ξ τ πα ρ τ ξ for , ( 12
) 0 ≥ x 0 ≥ t
The calculated domain has been extended to an infinite medium; hence, the virtual heat sources have to impose in the negative region of the medium to conserve the total energy absorbed in the substrate. The source term S(x,t) should be written as:

) ( ) exp( ) , ( τ ξ β β τ ξ I A S S - = for ∞ < < -∞ ξ (13)
The analytical solutions of the transient temperature distribution in Eq. ( 12) can be integrated using 104-point Gauss-Legendre quadrature.

To solve the transient temperature distribution of the entire system, i.e. quartz/polymer/substrate, the governing equation of energy conservation coupled with the initial and boundary conditions as well as the source term could be numerically solved by a fully implicit finite difference method [START_REF] Patankar | Numerical Heat Transfer and Fluid Flow[END_REF]. This method employs a two point backward difference approximation in time (Euler implicit scheme), a central difference approximation for the second derivative in the spatial coordinate, and an upwind difference approximation for the first derivative in the spatial coordinate. Since the thermal-affected region in the entire system is confined to the region nearby the surface of the substrate, the computational domains of quartz mold and substrate are considered as 200 μm.

III. RESULTS

Table 1 listed the thermal properties and optical parameters of materials used in this study. Excimer lasers with three different wavelengths (ArF, 193 nm; KrF, 248nm; XeCl, 308nm) are considered as pulsed laser sources herein. Figure 2 shows the surface temperature of copper and silicon substrate irradiated by XeCl excimer laser with 20 ns pulse duration. The temperature increases as the laser fluence increases. The temperature at the surface of copper substrate is much lower than that at silicon substrate due to the smaller absorption coefficient of copper. However, the literature suggests the laser fluence (XeCl, 20 ns, single pulse) irradiating on silicon substrate should be at least 0.35 J/cm 2 [START_REF] Xia | Ultrafast Patterning of Nanostructures in Polymers Using Laser Assisted Nanoimprint Lithography[END_REF]). The peak temperature at silicon surface should be around 700 ℃ when the laser fluence is 0.35 J/cm 2 , as shown in Fig. 2b. Therefore, the required laser fluence irradiates on the copper substrate could be adopted as 0.6 J/cm 2 according to the results in Fig. 2a, which could be the suitable condition for the LAN process. The surface temperature histories of copper substrate irradiated by XeCl excimer laser with fluence of 0.6 J/cm 2 for different pulse duration are depicted in Fig. 3. Both the results obtained from the analytical and the numerical approaches are shown in this figure. The numerical approach considers the existence of fused silica mold and polymer resist and the results exhibit the analytical results are almost the same with the numerical results, which means the existence of mold and polymer resist could be ignored, i.e., the assumption of the thermal-insulated boundary mentioned above is reasonable. The increase of the pulse duration decreases the maximum temperature at the surface of the substrate (several hundreds of Celsius degrees). The temperature profiles with time at the different locations inside the substrate with the irradiation of 0.6 J/cm 2 of a XeCl excimer laser (pulse duration is 20 ns) are depicted in Fig. 4. The temperature drops dramatically as shown in this figure, for instance, the maximum temperature is about 760 ℃ at the surface and drops less than 150 ℃ at 5 μm in depth. The temperature affected region is limited within the extremely shallow depth in the substrate in the same manner as employing silicon as the substrate conventionally. To investigate the temperature of the polymer resist layer, the numerical results at the center of the polymer resist layer and the substrate's surface with the irradiation of 0.6 J/cm 2 of a XeCl excimer laser for different pulse durations are illustrated in Fig. 5. The results indicate the maximum temperature at the center of the polymer resist layer decreases slightly and the time which the maximum temperature occurs delays tens of nanoseconds while the pulse duration increases. The temperature in the polymer resist layer exceeds the polymer glass transition temperature (about 100 ℃) for all cases. The process time of LAN will increase several tens of nanoseconds while increasing the pulse duration of the excimer laser; meanwhile, the maximum temperature at the silicon's surface could be lowered for several hundreds of degrees. The results successfully demonstrate the feasibility that copper could be adopted as the substrate instead of silicon, which is conventionally used. The heating processes under different laser sources, such as ArF (193 nm), KrF (248 nm) and XeCl (308 nm) are discussed in the following figure. Figure 6 illustrates the temperature profiles with time at the center of the polymer resist layer and the substrate's surface for different laser sources when the laser fluence is 0.6 J/cm 2 and the pulse duration is 30 ns. The results in Fig. 6 reveal the temperatures at copper surface and the polymer for these cases are almost the same because copper is insensitive to the three wavelengths.

Based on the optical simulation above-mentioned, the energy absorbances in the copper substrate are 71.42%, 73.09% and 73.13% for ArF, KrF and XeCl laser sources, respectively. As shown in the insert in Fig. 6, the temperature at the copper's surface irradiated by ArF is slightly lower than that by KrF and XeCl. However, the energy absorbances in the silicon substrate are 39.5%, 42.1% and 53.3% for ArF, KrF and XeCl laser sources, respectively. The copper substrate absorbed much higher energy from the excimer laser pulse (around 20% more) than the silicon substrate, nevertheless, the temperate at the copper substrate is much lower than that at silicon substrate. This is owing to larger thermal conductivity of copper, which is more than two-fold of that of silicon.

IV CONCLUSIONS

This work has considered optical multiple reflection theory to perform both analytical modeling and numerical simulation of laser induced heating during LAN process on a copper substrate and to predict the thermal response of the polymer resist and substrate theoretically. The analytical modeling ignored the existence of the fused silica mold and polymer resist; however, as shown in the results, the analytical solutions were well agreed with the numerical simulation, which considered the existence of the fused silica mold and polymer resist. The temperature in the polymer resist layer exceeds the polymer glass transition temperature (about 100 ℃) while the copper substrate irradiated by the excimer laser with the fluence of 0.6 J/cm 2 presented in the present study. The results successfully demonstrate the feasibility that copper could be adopted as the substrate instead of silicon, which is conventionally used. The results revealed the increase of the pulse duration could decrease the maximum temperature at the surface of the substrate dramatically and increase only several tens of nanoseconds sequentially. The temperatures at copper surface and the polymer for these cases are almost the same because copper is insensitive to the laser sources, which are ArF, KrF and XeCl. The copper substrate absorbed much higher energy from the excimer laser pulse (around 20% more) than the silicon substrate, nevertheless, the temperate at the copper substrate is much lower than that at silicon substrate owing to larger thermal conductivity of copper.

) 9 -

 9 11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5 where I max is the maximum power density of the laser pulse at max τ = t and σ the standard deviation of the function and can be determined by the full-width at half-maximum (FWHM, ) of the pulse,

9- 11

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

9- 11

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

Figure 1 :Figure 2 :Figure 3 :Figure 4 :

 1234 Figure 3: The surface temperature histories of copper substrate irradiated by XeCl excimer laser with fluence of 0.6 J/cm 2 for different pulse duration obtained from analytical approximation and numerical simulation which considered the existence of fused silica mold and polymer resist layer.

Figure 5 :

 5 Figure 5: The temperature profiles with time at the center of the polymer resist layer and the substrate's surface when the XeCl laser fluence is 0.6 J/cm 2 under different pulse durations.

Figure 6 :

 6 Figure 6: The temperature profiles with time at the center of the polymer resist layer and the substrate's surface for different laser sources when the laser fluence is 0.6 J/cm 2 and the pulse duration is 30 ns.

Table 1

 1 Thermal properties and optical parameters of materials used herein. http://www.io.tudelft.nl/research/dfs/idemat/Onl_db/Id123p.htm; f http://www.corning.com/.

		Substrate			Si	Copper b		PMMA	Fused Silica f
	Density (kg/m 3 )			2300 a	8940		1170 c		2201
	Thermal conductivity (W/mK)		160 a	352		0.16 d		1.30
	Heat capacity (J/kgK)			707.71 a	384.9		1380 c		787.52
	Refractive index, ArF		(0.872, 2.757) b	(0.972, 1.403)	1.492 e	(1.560841, 0.0)
	Absorption coefficient (nm -1 )		0.1795 b	0.0914				0.0
	Refractive index, KrF		(1.570, 3.565) b	(1.470, 1.780)	1.492 e	(1.508601, 0.0)
	Absorption coefficient (nm -1 )		0.1806 b	0.0902				0.0
	Refractive index, XeCl		(5.013, 3.689) b	(1.350, 1.710)	1.492 e	(1.485663, 0.0)
	Absorption coefficient (nm -1 )		0.1505 b	0.0698				0.0
	a Campbell	2001;	b Lide,	2003;	c Van	Krevelen	1976;	d Chu	et	al.	2001;
			Substrate

e

Polymer Mold Substrate Polymer Mold Laser Substrate RIE