Hoyd-Gigg Jack

Marc Ng

Aongus Desmulliez

Himanshu Mccarthy

Kevin Suyal

Duncan P Prior

Hand

Ng

Marc P Y Desmulliez

Aongus Mccarthy

Himanshu Suyal

Kevin A Prior

Duncan P Hand

UV Direct-Writing of Metals on Polyimide

I. INTRODUCTION

Direct-writing techniques of metallic patterns, such as laser or ink-jet printing, are additive processes which eliminate the traditional resist moulding and mask-based photo-patterning steps. Such techniques reduce the amount of wasted materials, the turnaround time and manpower cost. Laser, electron-beam and X-ray direct-writing have been employed for a long time in an enclosed pressurized chamber [START_REF] Ito | Pushing the limits of lithography[END_REF]. In search of low cost direct-writing processes without using any vacuum facilities, new types of processes have emerged which can be carried out through the coating of liquid solution or solid over-layer. The chart in Fig. 1 shows a review of the recent approaches employed by researchers worldwide. To minimize the effect of diffraction of light through liquid, direct-writing through a thin solid over-layer coating on the substrate is favored in order to achieve high resolution of sub-micron features. We propose a novel technique of direct laser writing, at atmospheric pressure, of metallic patterns onto a polyimide. Polyimide is a high performance polymer widely used in microsystems technology owing to its many desirable properties such as low dielectric constant (low-k), high mechanical strength, high resistance to moisture and heat (Tg = 365 o C), chemical inactivity as well as bio-compatibility. The ability to direct-write metals on polyimide substrates in air without using the conventional photolithography method can thus motivate developments of innovative applications such as advanced interconnections, chip packaging in microelectronics, integration of electrical functions in Bio-MEMS implants and micro-fluidic devices.

II. DIRECT-WRITING METHOD

The direct-writing approach presented here is based on the simple wet chemical pre-treatment of the polyimide substrate in order to incorporate silver ions into a thin depth of the polyimide surface layer [START_REF] Akamatsu | Site-Selective Direct Silver Metallization on Surface-Modified Polyimide Layers[END_REF].

Subsequently, a thin photosensitive polymer coating is casted on top of the iondoped polyimide. UV irradiation exposure through this over-layer coating releases electrons which lead to reduction of the silver ions which are electrostatically immobilized within the polyimide surface. The UV irradiation can be carried out via a UV lamp with a photomask or using a UV continuous-wave (CW) laser to achieve micro-patterning. After UV exposure, washing steps using deionized water and dilute H2SO4 solution remove the polymer over-layer coating and the remaining unreacted silver ions on the polyimide substrate. Finally the polyimide sample with direct-written silver patterns is annealed at 250 o C to allow for reimidization of the modified polyimide surface and diffusion and growth of the photoreduced silver particles within the Conductive copper layer is then deposited selectively on the silver surface by electroless plating. Each of the above process steps are explained as follows:

A.

Wet chemical pre-treatment Thin Kapton ® polyimide sheets were tapped onto microscope glass slides. The ion-exchange reactions start from the KOH chemical treatment of the polyimide substrates. When a polyimide substrate is immersed into 5M KOH solution at 50 o C for 5 minutes, the surface becomes hydrolyzed to a depth of at least 1 µm, in the form of a polyamic acid potassium salt. After rinsing with deionized water, the substrate is immersed in a 0.1 M AgNO3 solution at room temperature for 30 minutes in order to exchange sufficient Ag + ions from the electrolyte solution with the K + ions in the substrate. KOH and AgNO3 are both low cost, common inorganic salts. They are non-hazardous materials therefore no high waste disposal cost is incurred. The immersion processes does not require inert gas atmosphere and although a clean room environment may be preferred, it is not critical to the doping of ions into the polyimide substrates.

B. Coating of photoreactive polymer reducing agent

The polymer methoxy poly(ethylene glycol) (MPEG) can act as a reducing agent under UV irradiation [START_REF] Mallick | Polymer stabilized silver nanoparticles: A photochemical synthesis route[END_REF][START_REF] Mallick | Self-assembly of silver nanoparticles in a polymer solvent-Formation of a nanochain through nanoscale soldering[END_REF][START_REF] Mallick | Silver nanoparticle catalysed redox reaction: An electron relay effect[END_REF][START_REF] Mallick | Self-assembly of silver nanoparticles: Formation of a thin silver film in a polymer matrix[END_REF][START_REF] Mallick | Self Assembly of the Metal Nanoparticles: Formation of the Highly Oriented, Core-Shell Type, Bimetallic Gold-Silver Film[END_REF]. This polymer is utilized as an over-layer coating to enable the direct-writing process presented here.

The MPEG is dissolved in absolute ethanol with a concentration of 100 g.l -1 . At this concentration and using ethanol as a solvent instead of methanol or water, a smooth, thin (~1 µm) transparent coating is casted by spin-coating at 4000 rpm for 30 seconds. The MPEG solution can also be used to dipcoat contour surfaces for direct-writing.

C. UV direct-writing

The patterning process using a mercury arc lamp or a continuous wave (CW) UV laser can be carried out in air or in a clean room since the MPEG coating layer is solid at atmospheric pressure and does not need to be treated in any inert gas environment. This makes the handling of the coated substrates under UV exposure convenient and advanced manipulation such as 3D direct-writing can be achievable (Fig. 6). In addition, low power CW UV directwriting has an important advantage over heat dominated laser writing processes because, in those processes, diffusion of heat can give rise to spreading, irregularity of the final deposit, damage and debris on the substrate [START_REF] Chen | Laser-induced site-selective silver seeding on polyimide for electroless copper plating[END_REF]. A HeCd laser (λ = 325 nm, P = 3 mW after focus) has demonstrated successful direct-writing.

D. Post-exposure anneal

After washing off the MPEG coating with deionized water and exchanging the unreacted silver ions with H + ions from dilute H2SO4 acid solution, the silver patterned polyimide is annealed at 250 o C in an oven for 30 minutes. The polyimide surface layer modified by the KOH solution is re-imidized back into the original polyimide molecular structure, whilst the silver metal particles encapsulated in this polymer matrix layer migrate and grow into bigger clusters.

E. Electroless copper plating

The silver clusters formed on the substrate are isolated by the polyimide polymer matrix. However, these silver surfaces serve as suitable nucleation sites which can initiate electroless metal crystal growth.

Therefore the directwritten silver patterns act as a catalyst seed layer for selective electroless plating.

A conventional low pH formaldehyde based electroless copper solution is used which contained 16 g.l -1 CuSO4, 48 g.l -1 NaK-tartrate, 28 g.l -1 NaOH and 12 ml.l -1 aqueous solution (37.2 wt.%) of formaldehyde.

Well-defined conductive copper was electrolessly deposited exclusively onto the patterned silver seed.

III. CHARACTERISTICS OF COPPER DEPOSITS

Greek Cross test structures (Fig. 2345) were fabricated via an acetate photomask and a mercury arc lamp with power intensity ~ 48 mW cm -2 . A range of line widths down to 50 µm was achieved which were all in good agreement with the photomask. The deposited copper has a shinny appearance, well defined edge and good adhesion to the substrate. The sheet resistance Rs of the test structure with 250 µm line width was found to be 0.453 Ω.cm -2 as calculated by measuring the voltage drop V across the adjacent arms of the cross with an applied current I across the two other arms of the cross: Rs = π/(ln 2) . V/I.

(1) The 3-D helix silver coil fabricated by laser direct-writing on a cylindrical polyimide substrate (Fig. 6) has continuity over the whole length of the helix (~ 28 cm) and a line width of ~ 15 µm which is in good agreement with the laser spot size employed.

IV. CONCLUSION

A novel direct-writing fabrication process for micropatterning of metals on polyimide is presented. This process can be carried out in air, on contoured surfaces, and requires only low fluence UV light sources and low cost source materials. The line widths of the tracks fabricated by UVphotomask method were in good agreement with that of the mask. Using laser direct-writing, smaller track width can be achieved with accuracy. Implementation of this method for fabrication of MEMS packaging and other devices requires reliable control over the resistivity, adhesion, feature size and thickness of the deposit.

Fig. 1 .

 1 Fig. 1. Review of recent research in direct-writing of metals on nonconducting substrates. The work described in this article is highlighted.

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

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

Fig. 2 .

 2 Fig.2. Greek Cross test structure with 250 µm line width.

Fig. 3 .

 3 Fig.3. Magnified image of the structure shown in Fig. 1.

Fig. 4 .

 4 Fig.4. Test structure with 100 µm line width.

Fig. 5 .

 5 Fig.5. Test structure with 50 µm line width.

Fig. 6 .

 6 Fig. 6. Long silver helix track fabricated by rotating the cylindrical polyimide substrate (on a 1 cm diameter glass pipette) on a horizontal transition stage under the laser. Inset shows a magnified image of the track with a line width of ~ 15 µm.

ACKNOWLEDGMENT

The authors would like to thank for the financial support from the UK Engineering Physical Sciences Research Council (EPSRC) through its Basic Technology Program. The work was conducted under the project entitled "A thousand Micro-emitters per square millimeters" referenced GR/S85764. The authors also acknowledge the support of EPSRC-funded project 3D-Mintegration (www.3dmintegration.com). 9-11 April 2008 ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5


D Flynn

Marc P Y Desmulliez

Design Methodology and Manufacture of a Microinductor

Potential core materials to supersede ferrite in the 0.5-10 MHz frequency range are investigated. The performance of electrodeposited nickel-iron, cobalt-iron-copper alloys and the commercial alloy Vitrovac 6025 have been assessed through their inclusion within a custom-made solenoid microinductor. Although the present inductor, at 500 KHz, achieves 77% power efficiency for 24.7W/cm 3 power density, an optimized process predicts a power efficiency of 97% for 30.83W/cm 3 power density. The principle issues regarding microinductor design and performance are discussed.

I.

INTRODUCTION

While the physical size of digital and analogue electronic circuits has been drastically reduced over the past 20 years, the size of their associated power supplies has been reduced at a much slower rate. As a result, the power supply represents an increasing proportion of the size and cost of electronic equipment. One of the main difficulties in the miniaturization of power conversion circuits such as DC-DC converters is the construction of inductors and transformers [START_REF] Huljak | Applied Power Electronics Conference and Exposition[END_REF][START_REF]International Technology Roadmap for Semiconductors 2001 Edition System Drivers (ITRS)[END_REF][START_REF] Lotfi | Issues and advances in high-frequency magnetics for switching power supplies[END_REF].

Recent efforts to miniaturise the overall size of DC-DC power converters, has resulted in an increase in switching frequency of the power circuit from the 100-500 kHz range to the 1-10MHz range. This effort has led to a reduction of the size of the energy storage components that dominate the converter volume. Several problems arise when frequencies are pushed into the MHz region. Core materials commonly used in the 20-500 kHz region such as MnZn ferrites, have rapidly increasing hysteresis and eddy current losses as the frequency is increased. Furthermore, eddy current losses in windings can also become a severe problem, as the skin depth in copper becomes small in relation to the cross section of wire used. Even if these problems are adequately dealt with, the resulting transformer/inductor is still one of the physically largest and most expensive components in the circuit.

Various types of microinductors and transformers using an array of different thin-film metal alloys and winding configurations have been fabricated and reported in the literature over the past 25 years [START_REF] Ohnuma | Co-Zr-O Nano-Granular Thin Films with Improved High Frequency Soft Magnetic Properties[END_REF][START_REF] Dezuari | High inductance planar transformers[END_REF][START_REF] Park | Ferrite-based integrated planar inductors and transformers fabricated at low temperature[END_REF][START_REF] Ahn | Micromachined Planar Inductors on Silicon Wafers for MEMS Applications[END_REF][START_REF] Park | A Sacrficical layer Approach to Highly Laminated Magnetic Cores[END_REF][START_REF] Arnold | Vertically Laminated Magnetic Cores byElectroplating Ni-Fe Into Micromachined Si[END_REF]. The results of these devices have produced mixed results with varying degrees of success due to the prominent challenge in the design and fabrication of such devices. The key to successful microinductor design is to achieve the optimal balance of the various design parameters for a given application e.g. inductance requirement, low winding resistance, set efficiency at high frequency, high power density etc whilst considering fabrication limitations.

The focus of this paper is to design and fabricate a microinductor that balances the various design tradeoffs yet displays adequate performance for a variety of power related applications. This includes using low-temperature fabrication techniques that permit integration with supporting IC circuitry. The component footprint will be minimized with a high winding packing density limiting the area used while still targeting reasonable inductance values within the 0.5-1 MHz, so that it could be used in a highly miniaturized switched mode power supply for stepping up and down a wide range of voltages. The impact of core geometry on core material properties is highlighted as is the degree of enhancement of the microinductor efficiency and power density following a simple optimization process.

II. INDUCTOR DESIGN

A. Magnetic Core Geometry

A decision on which category of component is optimal requires a comparison across a variety of factors such as power handling, efficiency, practicality of fabrication, inductance requirement etc. Consider a core material for a one-turn micro-inductor. The saturation flux density, sat B , [START_REF] Huljak | Applied Power Electronics Conference and Exposition[END_REF] is a predetermined constant for a given material. In the same way, the current applied to the component would be known for a given application. Therefore, the only free parameter left will be either the relative permeability or the flux path length, since

The core, typically a rectangle, triangle or a circle, has a certain perimeter. Considering a rectangular core within a one turn pot-core component, shown in Fig. 1 with the main flux path indicated. As frequency increases the skin depths of the magnetic material and winding decrease. Therefore, to minimise skin depth losses and maintain the inductance value i.e. overall magnetic core area, the device becomes thinner and wider. The resulting increase in flux path length requires an increase in relative permeability for the inductance value to be maintained. An increase in relative permeability for the same resistivity value will make the material susceptible to skin depth losses at reduced frequency.

Magnetic Core Conductor

Flux Path If, for the same area and core material, the rectangular geometry was changed to a triangular geometry, then the perimeter is reduced by one third. The required permeability is then reduced by 1/3 and greater skin depth is achieved for the same resistivity.

A one turn micro-inductor with a triangular trench structure was fabricated by Sullivan et al [START_REF] Daniel | Design of microfabricated inductors[END_REF]. The fabrication of this V-groove micro-inductor required sputter deposition. Sputtering is a slow and costly process, and due to delamination can only be used for small core areas. Section will demonstrate how limiting the number of laminate layers can have an adverse effect on component performance.

A circular component would suffer from the same technical disadvantages as the triangular component. Moreover, a circular component fabricated by micromachining methods would require non-planar construction. Minimal footprint and low profile are prerequisites for present and future magnetic components. Therefore, a circular component was not investigated.

Due to the aforementioned complications of triangular and circular geometries. the investigated components herein are of rectangular planar form. Disadvantages of this geometry will be overcome via tailored core properties and optimal component design.

Another consideration is the tailoring of magnetic film properties with magnetic field annealing. Both pot-core and solenoid geometries can be used with magnetic field annealing; therefore, this is not a design restriction.

B. Winding Geometry

The winding geometries generally encountered for the inductor are of the spiral, meander or solenoid forms. A single layer solenoid component allows near ideal performance [START_REF] Seitz | Squeeze More Performance Out of Toroidal Inductors[END_REF]. The number of turns per single layer of traditionally wound solenoid inductors is limited by the core's inner diameter and how tightly you can pack the turns together. To increase inductance, a larger core or multiple layers of windings would normally be needed. This would normally lead to an increase in component size and parasitic capacitance. UV-photolithography discussed within the manufacturing of the solenoid allows a higher winding density avoiding such methods.

Sometimes an important issue in choosing a winding configuration is the magnitude and distribution of external magnetic fields generated by the transformer or inductor. The conventional pot-core is regarded as performing more favorably in minimizing external magnetic fields, as the core encloses and so shields the windings. The result is low external fields regardless of the distribution of primary and secondary windings inside the core. Solenoid components also have low external fields due to the winding distribution. In transformer applications the primary and secondary windings are customarily distributed evenly around the core, lying on top of each other with little space between. Thus the currents cancel each other, so little external field results. In micro-fabrication the ability to finely pattern windings makes it relatively easy to arrange the windings so that primary and secondary currents locally cancel. Hence, external fields are not a primary concern in component development within this work.

Perhaps the most important factor in comparing the potcore, spiral type winding, and solenoid designs is the ease of manufacture. This parameter is the most difficult to quantify and obviously dependent on fabrication capability. In the solenoid case, a difficulty often encountered is the interconnection of the top and bottom winding layers to encircle the core. This requires accurate alignment between the layers to insure proper connections, requires lowresistance contacts and connection over a vertical distance greater than the thickness of the core. A novel flip-chip bonding procedure overcomes the aforementioned difficulties [START_REF] Hua | An Analysis of a Microfabricated Solenoid Inductor[END_REF].

In the pot core design, the inter-layer connection requirements are less severe because the windings are planar and manufactured in a single deposition process. A complication arises if the inductor winding is spiral as access to the output pad is required. This typically requires a fourth deposition step, or the use of external connections such as bond wires. The thickness of the conductor layers is also typically larger than the thickness of the core layers, therefore, the vertical distance over which connections must be made is larger in the pot-core design. Furthermore, to form a closed magnetic core the upper and lower sections of the core have to be joined. The points where the two layers join is within the main flux path and can generate an unfavorable leakage flux leading to an increase in winding resistance, electromagnetic interference (EMI) and a reduction in component inductance.

C.

Analytical Design FlowChart Fig. 2 displays the design flowchart for the development of a microinductor. The solenoid microinductor developed within this work formed a closed magnetic path. The accuracy of the analytical equations will be reduced with components that include air-gaps within the main flux path due to fringing effects. The equations within Fig. 2 are used for the theoretical data within section III.

ISBN: 978-2-35500-006-5 COMPONENT CHARACTERISATION

J I

A sat w =       +             +       × = δ δ δ δ δ b b b b b L L DC AC 2 cos 2 cosh 2 sin 2 sinh 2 1       +             -       × = δ δ δ δ δ ω b b b b b L R DC AC
The novel microinductor fabricated via flip chip assembly is shown in Fig. 3. The manufacture process of the inductor and properties of the core materials utilized are described in detail within [START_REF] Flynn | Study of a solenoid microinductor operating in the MHz frequency range[END_REF]. The (*) symbol is used to denote magnetically anisotropic films, as displayed in Fig. 3. A critical point in the performance of an inductor is the measurement of the inductance value as a function of frequency and the additional core losses. A Hewlett Packard 4192A LF-impedance analyzer was used to record the inductance, resistance and Q-factor over a frequency range of 1 kHz-10MHz. The resistance value shown in Fig. 4 is for an aircore component. The reason for omitting the magnetic core was to remove core AC resistance from the recorded value. Therefore, the increase in resistance at high frequency from the R DC value is due to skin and proximity affects with increasing frequency. The winding is 90µm thick and at 1MHz has a skin depth value of approximately 66µm. Due to the winding consisting of a single layer and inter-winding spacing optimized, the proximity effect is minimized and the main contributor to the ac resistance is assumed to be the skin effect. The resistance of the components with the magnetic core included are shown in Fig. 5. The minor variation between experimental and analytical values may be due to the limitations of the 2D model representing the intrinsically 3D nature of the leakage and eddy current flux. The self resonant frequency (SRF) of an inductor is defined as the frequency at which the reactance becomes zero, (null Q-factor). The measured reactance increases linearly with frequency; hence the SRF induced by parasitic capacitance effects does not occur in the frequency range of interest. The Q-factor of the electrodeposited alloys is displayed in Fig. 7. The maximum Q-factors of the alloys occurred in the 500 kHz-1MHz frequency range indicating that the thin film dimensions are acceptable for operation within this range. Due to the properties of the CoFeCu core the 10µm thick core layer remains below one skin depth at 1MHz. The thicknesses of the components windings are under 2xskin depth therefore AC winding losses are maintained at acceptable levels. Therefore the ratio of the stored power to dissipative power of the component, the Qfactor, was optimal in this frequency range.

The Q-factor response of the commercial alloy demonstrates the susceptibility of this particular film to skin depth effects. The skin depth effects increase the dissipative losses within the component, therefore reduce the Q-factor. The high Q-factor in the low frequency range is greatly reduced at 1 MHz with respect to the CoFeCu film.

Using the theoretical DC saturation current formula within Fig. 2. the maximum applied current can be determined. The performance of the prototypes is summarized in Table 1. K is taken to be equal to 4.44, sinusoidal waveform, and The saturation current of the Vitrovac material is only 1.33mA. Hence, high efficiency is achieved at the expense of power density. An appropriate reduction in the number of windings and increase in core area via lamination will result in the same overall inductance. This would reduce the major Joule loss mechanism, winding loss, and allow windings of larger dimensions to be fabricated in order to apply greater current. The 90µm thick windings limit the maximum current to 180mA which is well beneath the saturation current of the CoFeCu components. An optimal component design process is therefore required.

IV. OPTIMAL DESIGN PROCESS

A simplified optimal design procedure for the solenoid component is outlined within this section. The optimization is performed in two circumstances in reference to the CoFeCu prototype in section III; [START_REF] Huljak | Applied Power Electronics Conference and Exposition[END_REF] the aim is to maintain the inductance value whilst increasing the power density significantly and improve efficiency, and (2) the primary aims are two maintain inductance and improve efficiency. With good relative permeability, small coercivity, reasonable resistivity and the largest saturation current (1.24A), the CoFeCu alloy is selected for optimization. Once a frequency is selected, in this case 0.5 MHz, skin depth values for the windings and core laminate are 93µm and 17.4µm, respectively. For comparison with the prototype, the inductance required will be 0.3µH (Fig. 12). The core area and numbers of turns are therefore: Design constraints are included at this stage. Each CoFeCu laminate will equal ½ skin depth and the number of laminations will be restricted to 10. The number of laminations will affect component manufacturability and cost such that N is only 11 turns in our case. The maximum applied current is 1.24A, PCB current density is 10A/mm² and winding height is restricted to 2xδ. The dimensions and performance of the optimized component can therefore be calculated and are given in Table 2. Optimized component (1) produces a 6% increase in efficiency and 7 fold increase in power density. Conventional inductors and transformers will normally operate with efficiency in the range of 90-95%. Hence, the application would have to greatly benefit from the increase in power density to tolerate the level of efficiency. Optimized component [START_REF]International Technology Roadmap for Semiconductors 2001 Edition System Drivers (ITRS)[END_REF] uses the data from the prototype to identify eddy current core loss as the main loss mechanism. The core is reduced to ten 1µm laminates and via this adjustment improves efficiency by 20%. Comparing the two optimized components, a trade off between efficiency and power density is evident.

V. CONCLUSION

Micromachined inductors with different magnetic cores have been fabricated on glass using micromachined techniques borrowed from the LIGA process. The analytical results using the two dimensional model agree well with the experimental values of inductance.

The challenge of developing inductors and transformers for MHz DC-DC converter operation requires the development of suitable core materials and manufacturability of laminated layers. This fact is demonstrated by the results within Tables 1 and2. The micro-inductor presented in this article has proved that an efficient component can be manufactured via an inexpensive UV LIGA and electroplating process.

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

Fig. 1 .

 1 Fig. 1. Simple rectangular one-turn micro-inductor.

Fig. 2

 2 Fig. 2 The Design Flowchart of the Microinductor

Fig. 3 .

 3 Fig. 3. Fabricated micro-inductor. The O-shape core is assembled between the windings layers prior to flip-chip bonding. The component is approximately 5mmx2mmx0.25mm (lxbxh), respectively. When Ni-Fe is anisotropic the orientation of the Easy and Hard axes are indicated.

Fig. 4 .

 4 Fig.4. Winding resistance vs. frequency for 90µm thick copper windings

Fig. 5 .

 5 Fig. 5. Micro-inductor resistances as a function of frequency. The resistance represents the winding and core contribution.

Fig. 6 . 5 Fig. 6

 656 Fig. 6. Inductance vs. Frequency of the electrodeposited alloys

satB

  is used to maximize the power density of the component.

Fig. 7 .

 7 Fig. 7. Q-factor vs. frequency for the electrodeposited alloys

TABLE 1 MICROINDUCTOR

 1

	PERFORMANCE FOR RESPECTIVE CORES
	AT 0.5MHZ

TABLE 2

 2

		OPTIMISED PERFORMANCE & PARAMETERS
	Dimensions	Optimized component (1)	Optimized component (2)	Prototype Component
	Number of turns	11	33	33
	Turn thickness	180 µm	90 µm	90 µm
	Turn width	550 µm	200 µm	200 µm
	Turn spacing	55 µm	20 µm	20 µm
	Laminations	2	10	1
	Lamination thickness	5 µm	1 µm	10 µm
	Lamination width	500 µm	500 µm	500 µm
	Lamination insulation	5 µm	5 µm	5 µm
	Performance	Optimized component	Prototype Component	Prototype Component
	in V	0.17	0.51	0.51
	out P	0.2108 (W)	0.0918 (W)	0.0918 (W)
	B	pk	1.4(T)	1.4(T)	1.4(T)
	eddy P	5.54(mW)	0.22(mW)	24.6(mW)
	hys P	0.57(mW)	0.11(mW)	0.11(mW)
	Cu P	184(mW)	1.62(mW)	1.62(mW)
	Efficiency (%)	83	97	77
	Power density (W/cm³)	174.8	30.83	24.7

ACKNOWLEDGMENT

Thanks are extended to Prof Cywinski for the VSM data performed at the University of Leeds, and to Raytheon Systems Limited for performing the resistivity measurements. This work was made possible through the funding of the Scottish Consortium in Integrated Micro Photonic Systems (SCIMPS) funded by the Scottish Funding Council under the Strategic Research Development Grant scheme.


Jens Georg Kaufmann

Marc P Y Desmulliez

Dennis Price

Megasonic Enhanced Electrodeposition

à la diffusion de documents scientifiques de niveau recherche, publiés ou non, émanant des établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

INTRODUCTION

The increasing consumer demand for faster, lighter and smarter electronic devices calls for enhanced system integration and packaging technologies. Key to the increasing density of electronic components is the introduction of the high density interconnection (HDI) technology in printed circuit boards (PCB), resulting in multilayer technology and increasing amounts of electrical connections that need to be handled. Microvias are elements that are particularly important as they allow the reduction of the footprint of electronic components through the redistribution of interconnects in the underlying layers. Microvias are formed by mechanical drilling or laser ablation of the PCB material and subsequent electroplating is used to fill the cavity to render them electrically conductive. The microvias need however a conductive seed layer on the side walls to allow the electrodeposition of metal. The technologies used in that respect are mainly autocatalytic plating processes or direct metallisation [START_REF] Lau | An overview of microvia technology[END_REF].

To allow further integration and miniaturization, it is necessary to form microvias with an aspect ratio (height over diameter hole ratio) greater than 1:1. This current limit arises because of the difficult hydrodynamic conditions and current crowding effect at the mouth of the hole; both conditions attenuate the convection in the electrolytic solution near the surface of the substrate. This lack of effective agitation reduces the ion concentration in the solution within the immediate proximity of the microvia, increases the Nernst diffusion layer and limits therefore the deposition rate of the metal [START_REF] Castaldi | Limits of copper plating in high aspect ratio microvias[END_REF]. To overcome this limitation, extensive research utilising a variety of additives and current waveforms was performed. The additives used are mainly large organic inhibitors, and small complex builders that accelerate the deposition at the bottom of the via cavity. This is a very complex method, which needs a carefully controlled solution. However, super conformal plating of deep trenches and holes can be achieved, as exampled by the damascene process metallisation. The use of this process in high volume consumer electronics would nevertheless be problematic because of the high parameter variants found in the usually large volumes of solution used in PCM manufacturing. This paper proposes therefore to use megasonic agitation to reduce the diffusion layer and achieve optimal electrodeposition for microvias with aspect ratio larger than 1:1.

II. MEGASONIC AGITATION

This paper describes a new method to enhance the capability and compatibility of metallisation processes based on megasonic assisted copper electrodeposition. The primary focus of this article concerns the filling of high aspect ratio microvias. However, this technology is transferable to the manufacture of various high aspect ratio metal MEMS structures as well as the development of photoresist in deep trenches. A high frequency acoustic streaming of frequency f allows the modification of the near surface hydrodynamics. The electrical potential of the ox-red reaction is governed by the activity of the ion near the electrode surface, according to the Nernst equation:

d OZ e a a F z RT E E Re 0 ln + = (1)
Increasing the activity, a, and therefore the concentration, c, near the surface of the substrate, results in a decreased overpotential.

( )

+ ⋅ = = Z M M c a Z γ (2) δ 0 c dx dc = (3)
More precisely, the Nernst diffusion layer which, in normal conditions, is governed by the velocity of the medium stream over the solid surface, as in (4), see [START_REF] Yin | Duplex diffusion layer model for pulse with reverse plating[END_REF], depends on 1/f 1/2 in the presence of an acoustic field, as in (5). The high frequency (above 500 KHz) not only reduces the thickness of the Nernst diffusion layer but avoids the creation of cavitations as traditionally experienced in ultrasonic agitation. These highly volatile collapsing bubbles inhibit the usage of acoustic streaming in the plating industry, due to the likelihood of board delamination. The use of acoustic streaming on the diffusion layer can be found in cleaning processes for micro parts. Furthermore the modelling of the megasonic induced flow in trench like cavities is also described in [START_REF] Nilson | Enhanced Transport by Acoustic Streaming in Deep Trench-Like Cavities[END_REF] II.

EXPERIMENTAL WORK In order to achieve the filling of microvias of high aspect ratio, a full Design of Experiments (DOE) has been conducted to analyse the influence of megasonic agitation on the plating quality of the microvias. The considered parameters include the current waveform, the power of the transducer used and the included angle of radiation between the transducer and the PCB. These parameters were chosen due to previous work in microvia plating. This abstract focuses at this stage on the fundamental differences introduced by using direct current (DC), pulsed current (PP) and reverse pulsed current (RP) plating. The different power levels for the megasonic agitation range between 0W and 500Watts of electrical power to distinguish between the normal agitation within the tank and the acoustic streaming. To change the angle between the transducer and the test vehicle (cathode) a custom carrier for the transducer was designed and manufactured. This carrier allows changing the radiation angle without changing the distance between the transducer and the centre of propagation.

The transducer itself consists of a 4 by 4 inch wide PZT crystal plate contained in a laser welded container. The experiments were conducted in a 60 litres polypropylene tank. The electrolyte contains the standard components for As measure of the quality of the microvias, several criteria were specified. The throwing power and the amount of deposited metal per unit time were measured. Finally, the highest achievable aspect ratio was recorded. The filling of the microvia was possible till an aspect ratio of >2:1 was reached. The filling profile of the holes shows a largely over proportional deposition at the bottom of the blind vias. Larger aspect ratio could not be achieved because of the lack of seed conductive layer deep into the microvia cavity. This result was obtained for 125 and 250 W of radiated power, at any angle and with a DC waveform. Pulse electroplating is not necessary.

III. Conclusions

These results demonstrate the world's first application of megasonic agitation for the plating of microvias of aspect ratio greater than 1:1. The improvement in the deposition rate, as well as the quality of the deposits, make megasonic agitation a serious contender for the filling of deep aspect ratio microvias. The throwing power of the plating, the relation between the thicknesses on the mouth to the thickness on the bottom reached 5. This allows super conformal filling, as known from the damascene process in semiconductor manufacturing. This method does not require any special plating chemistry or optimization of waveform parameters. The effect of megasonic agitation on the crystalline structure needs however further investigation.

Fig. 1

 1 Fig. 1 Megasonic generator. Fig. 2 Megasonic transducer of 4 by 4 inch 500 W Electrical power at 1MHz active area

Fig. 3

 3 Fig. 3 Schematic of the transducer mounted Fig. 4 Insufficient (left) or no filling (right) on the carrier focussing the sound of the microvias towards the sample holder


R R Mansour

S Fouladi

M Bakeri-Kassem

Integrated RF MEMS/CMOS Devices

A maskless post-processing technique for CMOS chips is developed that enables the fabrication of RF MEMS parallel-plate capacitors with a high quality factor and a very compact size. Simulations and measured results are presented for several MEMS/CMOS capacitors. A 2-pole coupled line tunable bandpass filter with a center frequency of 9.5 GHz is designed, fabricated and tested. A tuning range of 17% is achieved using integrated variable MEMS/CMOS capacitors with a quality factor exceeding 20. The tunable filter occupies a chip area of 1.2 × 2.1 mm 2 . I.

INTRODUCTION

There is a need for intelligent reconfigurable radio Frequency (RF) front-ends that can achieve maximum hardware sharing for various standards. Such an RF front-end requires reconfigurable blocks rather than switched architectures. These reconfigurable blocks need to meet several RF specifications such as smaller size, lower loss, higher tuning range, linearity, and low power consumption. The continuous shrinkage of minimum feature size in current silicon-based technologies such as CMOS provides an opportunity to meet these demands. Moreover, using these technologies, it is possible to integrate the reconfigurable RF blocks with the core signal processing circuitry leading to the implementation of compact intelligent reconfigurable RF front-ends.

Tunable filters and amplifier tunable matching networks are among the reconfigurable blocks that if integrated with RF CMOS electronics can reduce the overall system size, weight and cost. A number of integrated tunable filters have been demonstrated using solid-state tuning elements. Despite the small area and high quality factor of these active tuning elements, the major disadvantage is their susceptibility to noise and lower power handling due to nonlinearities in the active semiconductor devices.

This paper demonstrates the feasibility of integrating MEMS with CMOS to realize reconfigurable front end with a superior RF performance in terms of linearity and insertion loss. It also presents a novel post-processing technique to create MEMS varactors in CMOS chips. The paper will also address other applications that can benefit from the successful integration of MEMS with CMOS.

II. FABRICATION PROCESS

A CMOS micromachining process was reported in [START_REF] Fedder | Tunable RF and Analog Circuits Using On-Chip MEMS Passive Components[END_REF] to fabricate CMOS-MEMS tunable capacitors. The capacitors were based on interdigitated beam structure, and electrothermal actuators were used for tuning. We propose a postprocessing technique that enables the fabrication of parallelplate tunable capacitors with a higher capacitance value, smaller size and zero dc power consumption using electrostatic actuation mechanism. The CMOS-MEMS postprocessing of tunable capacitors starts with a chip fabricated using the TSMC 0.35μm CMOS technology. There are two polysilicon and four metal layers available through this process. The process offers the possibility for the fabrication of MEMS structures using metal interconnect layers. The layout of the tunable parallel-plate capacitor and its crosssectional view are presented in Fig. 1. The top and bottom plates are made of M3 and M1 layers, respectively and the second metal layer (M2) is used as a sacrificial layer which will be removed after post processing in order to create an air gap between the plates. By using this approach after release, the total gap between the plates consists of two dielectric layers each having a thickness of 1μm and a 0.6μm air gap. The dielectric layer prevents the occurrence of a short circuit when the dc tuning voltage exceeds the pull-in voltage, and the top plate snaps down on the bottom plate.

Three maskless post-processing steps are required to construct the tunable capacitors. They also include improving RF performance by etching away the silicon under the device and eliminating substrate loss. A schematic view of the postprocessing steps is presented in Fig. 2. The first processing stage involves the removal of silicon dioxide around the MEMS structure as presented in Fig. 2(b). This can be done using the last metal layer (M4) as a mask and reactive ion etching (RIE) of the oxide layer [START_REF] Baker-Kassem | Novel High-Q MEMS Curled-plate Variable Capacitor Fabricated in 0.35 μm CMOS Technology[END_REF]. The main purpose of this step is twofold. First, to expose the sacrificial M2 layer that will be etched during the next post-processing step. Secondly, to create windows through the oxide down to the silicon substrate, which are used to form a trench under the capacitors improving their quality factor and self-resonance frequency. It is important to keep an oxide layer around the structural metal layers (M1 and M3) to protect them during the removal of the M2 sacrificial layer. This is accomplished by extending the masking metal layer (M4) on top of the structural layers. However, the second metal layer which is required to be exposed after the RIE step, should be extended further than M4.

As shown in the cross-sectional view of the proposed capacitor in Fig. 1, the extension of M4 over M1 and M3 is 2µm. assuming that the RIE etch recipe is anisotropic, this ©EDA Publishing/DTIP 2008 ISBN: 978-2-35500-006-5 9-11 April 2008 leaves a 2µm of side-wall oxide to protect the structural metal layers. The M4 layer also covers the CMOS electronics circuitry. Fig. 3(a) presents an SEM image of the capacitor after the first post-processing step. As shown in this figure, M2 and the silicon substrate are exposed and ready to be etched during the next step. The second post-processing step as shown in Fig. 2(b) involves the etching of the sacrificial layer and substrate using wet etching techniques. The sacrificial layer is removed using PAN etch followed by TiW etchant. In addition to the sacrificial layer removal, the exposed M4 layer used as a mask during the RIE step will be removed. For wet etching of silicon, diluted electronic grade TMAH is used. For an acceptable RF performance, a minimum trench depth of 75µm is required. An SEM image of the CMOS-MEMS capacitor after the wet etching step is shown in Fig. 3(b). As shown in this figure, in order to facilitate the etching of the sacrificial metal layer sufficient number of release holes are included on the top plate of the capacitor. During the last post-processing step, devices are rinsed in IPA and released in a critical point dryer (CPD) system to avoid any stiction problems. This is followed by a final RIE step that is required to remove the protecting oxide layer on top of M3 and RF pads as shown in Fig. 2(c). We have been also successful in implementing the process in 0.18 μm CMOS. Figure 4 shows an integrated 2-pole tunable filter implemented using four MEMS/CMOS varactors. The measured performance of the tunable filter is shown in the same figure [START_REF] Fouladi | An Integrated Tunable Band-Pass Filter Using MEMS Parallel-Plate Variable Capacitors Implemented with 0.35 μm CMOS Technology[END_REF]. Figure 5 illustrates SEM pictures of MEMS varactors implemented in CMOS chips using the above proposed post -processing technique. The varactor exhibits a Q value of better than 300 at 1.5 GHz [START_REF] Baker-Kassem | Novel High-Q MEMS Curled-plate Variable Capacitor Fabricated in 0.35 μm CMOS Technology[END_REF].

Figure 1

 1 Figure 1 Layout and cross-sectional view of the parallel-plate MEMS tuning elements fabricated using the 0.35µm CMOS process.

Fig. 3 .

 3 Fig. 3. SEM image of the capacitor after first (a) and second (b) postprocessing steps.

Figure 4

 4 Figure 4. a) Optical photograph of the MEMC/CMOS tunable filter, b) the measured return loss at different bias voltages.


Zhen-Zhun Shu

Ming-Li Ke

Guan-Wei Chen

Ray-Hua Horng

Chao-Chih Chang

Jean-Yih Tsai

Chung-Ching Lai

Ji-Liang Chen

Design And Fabrication of Condenser Microphone Using Wafer Transfer And Micro-electroplating Technique

A novel fabrication process, which uses wafer transfer and micro-electroplating technique, has been proposed and tested. In this paper, the effects of the diaphragm thickness and stress, the air-gap thickness, and the area ratio of acoustic holes to backplate on the sensitivity of the condenser microphone have been demonstrated since the performance of the microphone depends on these parameters. The microphone diaphragm has been designed with a diameter and thickness of 1.9 mm and 0.6 μ m, respectively, an air-g a p t h i c k n e s s o f 1 0 μ m, a n d a 2 4 % a r e a r a t i o of acoustic holes to backplate. To obtain a lower initial stress, the material used for the diaphragm is polyimide. The measured sensitivities of the microphone at the bias voltages of 24 V and 12 V are -45.3 and -50.2 dB/Pa (at 1 kHz), respectively. The fabricated microphone shows a flat frequency response extending to 20 kHz.

I.

INTRODUCTION

Microphones are electro-acoustic transducers that convert acoustical energy into electric energy. At present, electro-acoustic transducers cover a wide field of applications in cellular phones, digital cameras, hearing aids and so on. There are few different designs of microphones. Most of the silicon microphones are condenser microphones because of their high sensitivity, low noise level, and flat frequency responses in a wide bandwidth [START_REF] Goto | Hi g h -performance condenser microphone with single-crystalline silicon diaphragm and b a c k p l a t e[END_REF][START_REF] Mu L L E N B O R N | T h e f i r s t low voltage, low noise differential silicon microphone, technology development and measu r e me n t r e s u l t s[END_REF]. Condenser microphones consist of two parallel plates, namely a flexible diaphragm and a rigid backplate. During the last decade, the miniature microphones based on silicon micromachining techniques were the subject of research and development [START_REF]i o n o f s i l i c o n c o n d e n s e r mi c r o p h o n e u s i n g c o r r u g a t e d d i a p h r a g m t e c h n i q u e[END_REF]. In general, for the micro-electro-mechanical system (MEMS) microphone applications, mechanical structures are fabricated on silicon wafers using thin film deposition, lithography and etching techniques [START_REF] Mi A O | An a l y t i c a l mo d e l i n g f o r b u l k -micromachined c o n d e n s e r mi c r o p h o n e[END_REF][START_REF] Al F O N S De H E | S i l i c o n mi c r o p h o n e d e v e l o p me n t a n d a p p l i c a t i o n[END_REF]. Coupled with matured silicon micromachining and MEMS techniques, the size scale of microphones can be decreased efficiently. In fact, lowering of the sensitivity is obtained by condenser microphone miniaturization. To improve its mechanical sensitivity, a corrugated diaphragm and several complex processes for making silicon condenser microphones have been reported [START_REF] Mi A O | De s i g n c o n s i d e r a t i o n s i n mi c r o ma c h i n e d s i l i c o n mi c r o p h o n e[END_REF][START_REF] Kr E S S Ma N N | S i l i c o n condenser microphones wi t h c o r r u g a t e d s i l i c o n o x i d e / n i t r i d e e l e c t r e t me mb r a n e s[END_REF][START_REF] Kwo N | Do u b l e -chip condenser microphone for rigid b a c k p l a t e u s i n g DR I E a n d wa f e r b o n d i n g t e c h n o l o g[END_REF][START_REF] Tajima | Hi g h -performance ultra-small single crystalline silicon micr o p h o n e o f a n i n t e g r a t e d s t r u c t u r e[END_REF]. Although these researches achieved high mechanical sensitivities, the extra processes will increase the production cost. To overcome these disadvantages, the condenser microphone fabricated by the wafer transfer and micro-electroplating technique wherein the substrate can also be recycled. This paper describes the new fabrication process and design as well as the acoustic characterization of the microphone.

II. Microphone Design

Condenser microphones consist of two parallel plates which form a variable gap capacitor. Figure 1 shows the basic structure of the condenser microphone. The upper plate is a thin and flexible diaphragm which can be made of polyimide, silicon nitride, or polysilicon. The lower plate is a thick and fixed backplate. The open-circuit sensitivity S o of a condenser microphone is defined by the product of the diaphragm mechanical sensitivity S m and the electrical sensitivity S e :

e m o S S S  (1)

A. Diaphragm

The mechanical sensitivity of the microphone can be expressed as

dP dw S m  ( 2
)
where w is the deflection of the diaphragm, and P is the pressure acting on the diaphragm. At low frequencies, the mechanical sensitivity is approximately given by [START_REF] S C H E E P E R | A r e v i e w o f s i l i c o n mi c r o p h o n e s[END_REF] where A ,  and d h a r e t h e d i a p h r a g m' s a r e a , t e n s i o n stress and thickness, respectively. Thus, choosing the diaphragm material to obtain a lower initial stress will result to a higher mechanical sensitivity [START_REF] Torkkeli | C a p a c i t i v e mi c r o p h o n e wi t h l o w-stress polysilicon membrane and high-s t r e s s p o l y s i l i c o n b a c k p l a t e[END_REF], on the other hand, microphones with thinner diaphragm have higher mechanical sensitivity. Mechanical sensitivity is obtained by measuring the central deflection of the diaphragm under a given applied pressure as shown in Fig. 2. It can be seen that the mechanical sensitivity increases when the diaphragm thickness is reduced. In this study, the design thickness and diameter are 0.6 µm, 1.9 mm, respectively, with polyimide as the diaphragm material.

Diaphragm

Backplate

Backchamber

Acoustic hole

Air gap

B. Air gap

The bias voltage ( b V ) applied to the condenser microphone provides an electric field in the air gap between the diaphragm and the backplate. The sound pressure variations cause a movement in the microphone diaphragm, resulting to a change in the voltage across the air gap. The electrical sensitivity of the microphone can be expressed as

d V d V S b b e     ( 4
)
where d is the air gap thickness. From the above equation, the electrical sensitivity is proportional to the bias voltage and inversely proportional to the air gap thickness. Theoretically, a higher voltage bias can increase the electrical sensitivity. However, the bias voltage is limited by the pull-in voltage ( p V ) of the microphone (

p b V V  ).
The pull-in voltage for a circular diaphragm can be expressed as [START_REF] Li | S e n s i t i v i t y -improved silicon condenser microphone with a novel single deeply corrugated d i a p h r a g m[END_REF]

m p S d V 0 3 27 8   ( 5
)
where 0  is the permittivity in air.

Combining equations ( 1), ( 4) and ( 5), the open-circuit sensitivity of the microphone would be

0 27 8  m o dS S  (6)
At this pull-in voltage, the diaphragm touches the blackplate and causes an electrical short. Although increasing the air gap thickness results to a higher open-circuit sensitivity and pull-in voltage, it also reduces the air resistance of the microphone. Nevertheless, the air gap thickness cannot be increased indefinitely since the capacitance between the diaphragm and the backplate will d e c r e a s e . He n c e , t h e a i r g a p t h i c k n e s s wa s s e t t o 1 0 μ m i n t h i s study.

C. Backplate

The air resistance of the air gap ( a R ) and the backplate hole ( h R ) can be expressed as [START_REF] Wa N G | Mo d e l i n g a n d characterization of a silicon condenser mic r o p h o n e[END_REF] ) 8

3 4 ln 8 2 ( 12 2 3     A A A n d A R eff a   (7) and 4 8 r n tA R eff h    (8)
respectively, where  is the viscosity of air, eff A is the area of the backplate, A is the ratio of the total acoustic holes to the total backplate area, n is the density of acoustic holes, t is the backplate thickness, and r is the radius of the backplate hole. The air damping and frequency response are controlled by the acoustic holes of the backplate. Figure 3 illustrates the effect of the area fraction of acoustic holes on the frequency response. It can be seen that the cut of frequency and sensitivity increases when the area fraction of acoustic holes is raised. Thus, when the area fraction of the acoustic holes is increased to 25%, the microphone sensitivity is raise d . to backplate, and so on. To optimize the sensitivity, the parameters used for the proposed MEMS condenser microphone are shown in Table 1.

III. Fabrication Process

A. Diaphragm

The diaphragm and the backplate were fabricated in different wafers. Figure 4 (a) shows the fabrication process of t h e d i a p h r a g m. Ab o u t 0 . 5 μ m t h i c k o f S i O 2 film was deposited on the p-type (1 0 0) wafer by LPCVD. Then, a polyimide layer is spin-coated on the oxide layer and patterned by photolithographic techniques. The 80 μ m t h i c k nickel frame was formed by electroplating after the electrode material (Cr/Au) was deposited on the diaphragm. Finally, the diaphragm chip was lifted off by the BOE.

B. Backplate

The microphone backplate was also fabricated by wafer transfer technique. Figure 4 (b) illustrates the backplate process which begins with SiO 2 deposition by LPCVD, and followed by the deposition of 200 nm of Cr/Au seed layer on the SiO 2 film using the evaporation process. This seed layer is patterned and etched to form the square-shaped acoustic holes.

The backplate structure was fabricated with nickel using the micro-electroplating technique. A 10 μ m t h i c k S U8 s p a c e r i s patterned and used to define the air gap between the diaphragm and the backplate. The microphone components were separated from the substrate by the BOE.

C. Prototype microphone

Figure 5(a) and (b) show the fabricated diaphragm and backplate chips, respectively. The two components of the condenser microphone are mechanically clamped together. To make the prototype condenser microphone, the chip is packaged in a metal housing with a preamplifier circuit as shown in Figure 6.

IV. Results and Discussion

A. Stress of the diaphragm

The mechanical sensitivity m S values obtained are 13, 15, and 17 nm/Pa for the diaphragm thickness values of 1, 0.8, a n d 0 . 6 n m, r e s p e c t i v e l y . S u b s t i t u t i n g t h e d i a p h r a g m' s a r ea, thickness, and mechanical sensitivity values into equation ( 3), the calculated average internal stress of the diaphragms are 8.68, 7.52, and 6.64 MPa. According to equation (3), mechanical sensitivity can be increased by thinning the diaphragm and decreasing the internal stress. Stress is a very important property of a microphone diaphragm. The thermal process should also be controlled in order to decrease the internal tensile stress of the diaphragm. Several materials were tried in the fabrication of diaphragm, but only polyimide has the advantage of having a low temperature fabrication process.

B. Electrical measurements

Using the mechanical sensitivity value of 17 nm/Pa for the d i a p h r a g m t h i c k n e s s o f 0 . 6 μ m ( F i g . 4 ) , t h e p u l l -in voltage is calculated to be about 44 V. Using an impedance analyzer (HP-4284), the capacitances of the packaged microphone were determined. Figure 7 plots the capacitance against the voltage bias without an applied static pressure. The capacitance increased as the voltage bias increased because the diaphragm was pulled towards the backplate. It is clear that the microphone chip operates stably when the voltage bias was less than 25 V.

Figure 8 shows the relationship between capacitance and static pressure at two different voltage biases. The complete set-up for measuring the capacitance change involves a pressure vessel. When the static pressure is varied with a pressurization cylinder, the diaphragm of the condenser microphone attached to the backplate moves and results to a change in capacitance. The capacitance increased linearly with pressure up to 100 Pa. The capacitance can also be increased by increasing the voltage bias due to the decrease of the air gap. Although, it can raise the sensitivity of the condenser microphone, its disadvantage is that the band of flat frequency response is reduced because the resonance frequency increases as the square root of the mechanical stiffness.

C. Frequency response

In order to test the sensitivity of the microphone, the device was placed in an anechoic room as shown in Fig. 9. The fabricated microphone is fitted with a preamplifier (B&K 2669), and the acoustic response is measured using an audio analyzer (B&K 2012). Voltage biases of 12 V and 24 V were used in the measurement.

The frequency response of the prototype microphone is shown in Figure 10. The voltage biases of 12 V and 24 V showed that the sensitivities of the condenser microphone were at -50.2 and -45.3 dB/Pa (at 1 kHz), respectively. Obviously, the sensitivity of the microphone increases with the voltage bias because the electrical sensitivity increases. However, a flat frequency response was obtained over the audio frequency range.

V. Conclusion

The sensitivity of the MEMS condenser microphone can be increased by increasing the mechanical and electrical sensitivities. Any attempt to optimize it will likely result into a compromise somewhere. For example, increasing the diaphragm area and voltage bias, as well as decreasing the diaphragm stress and thickness will raise the sensitivity, but the band of the flat frequency response is reduced. In this paper, the design and fabrication of a high sensitivity microphone is presented. The condenser microphone is fabricated by the MEMS technology through the wafer transfer and micro-electroplating techniques in order to avoid many extra processes. Based on the simulation results, when the area fraction of the acoustic holes is increased to 25%, outstanding microphone sensitivity could be obtained. The mechanical sensiti v i t y a t t h e d i a p h r a g m t h i c k n e s s o f 0 . 6 μ m i s 17 nm/Pa. Moreover, a sensitivity of up to -50.2 dB/Pa was measured for the fabricated microphone with a diaphragm d i a me t e r o f 1 . 9 mm a n d a t h i c k n e s s o f 0 . 6 μ m, u s i n g a v o l t a g e bias of 12 V. Moreover, the fabricated microphone shows a flat frequency response in the audio frequency range.

Fig. 1 .

 1 Fig. 1. Structure of the condenser microphone.

Fig. 2 .Fig. 3 .

 23 Fig. 2. Mechanical sensitivity at different diaphragm thicknesses.

9- 11

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

Fig. 4 .Fig. 5 .

 45 Fig. 4. Process flow for the condenser microphone: (a) diaphragm; (b) backplate.

Fig. 7 .Fig. 8 .

 78 Fig. 7. The capacitance characteristics for a microphone at different voltage biases

Fig. 9 .Fig. 10 .

 910 Fig. 9. Schematic diagram of the acoustic response measuring system.

9- 11

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

Table 1

 1 Design parameters of the MEMS condenser microphone.

	Parameters	Value
	Diaphragm material	Polyimide
	Diaphragm diameter	1.9 mm
	Diaphragm thickness	0 . 6 μ m
	Air gap	1 0 μ m
	Acoustic hole size	8 0 μ m × 8 0 μ m
	Backplate thickness	1 0 0 μ m
	Bias voltage	12 V


M Worgull

M Heckele

T Mappes

B Matthis

G Tosello

T Metz

J Gavillet

P Koltay

H N Hansen

Sub-µ structured Lotus Surfaces Manufacturing

Sub-micro structured surfaces allow modifying the behavior of polymer films or components. Especially in micro fluidics a lotus-like characteristic is requested for many applications. Structure details with a high aspect ratio are necessary to decouple the bottom and the top of the functional layer. Unlike to stochastic methods, patterning with a LIGA-mold insert it is possible to structure surfaces very uniformly or even with controlled variations (e.g. with gradients). In this paper we present the process chain to realize polymer sub-micro structures with minimum lateral feature size of 400 nm and up to 4 µm high.

I.

INTRODUCTION

A new generation of passive, capillary driven micro-fluidics systems is expected to enable advanced management of liquid and gas. In such enhanced flow structures, fluids are guided through channels with geometrically modified surfaces. In particular, surface sub-µ structuring of micro fluidic systems can be designed to introduce a structural gradient which can generate a driving force to move liquid samples along channel structures (see Figure 1). The realization of Lotus-like patterns on surfaces requires precise control of the structuring method. The manufacturing method of such structured surfaces should also enable mass-fabrication capability, for example by polymer replication. To demonstrate the moulding of sub-µ structures, an indirect tooling technology for the manufacturing of the mould insert and the hot embossing process for polymer replication has been selected. In this paper we present the process chain to realize polymer sub-micro structures with minimum lateral feature size of 400 nm and up to 4 µm high.

WETTING BEHAVIOR OF STRUCTURED SURFACES

The liquid repellency of a surface is principally governed by a combination of its chemical nature (i.e. surface energy) and, in case of stochastic surfaces, by its topography at the micro-scale (i.e. surface roughness). Although flat low surface energy materials can often exhibit high water contact angles [START_REF] Tsibouklis | Poly(perfluoroalkyl methacrylate) Film Structures: Surface Organization Phenomena, Surface Energy Determinations, and Force of Adhesion Measurements[END_REF] this is normally not sufficient to yield the reppellency of superhydrophobic surfaces. In order to obtain this, the difference between the advancing and the receding contact angle (contact angle hysteresis) must be minimal. Effectively, contact angle hysteresis is the force required to move a liquid droplet across a surface. In the case of little or no hysteresis, very little force is required to move a droplet, hence it rolls off easily [START_REF] Furmidge | Studies at phase interfaces. I. The sliding of liquid drops on solid surfaces and a theory for spray retention[END_REF] [START_REF] Miwa | Effects of the surface roughness on sliding angles of water droplets on superhydrophobic surfaces[END_REF].

Theoretical studies for idealized rough hydrophobic surfaces predict that contact angle hysteresis initially increases with surface roughness [START_REF] Johnson | Contact angle hysteresis. I. Study of an idealized rough surface[END_REF] until eventually a maximum value is reached. Greater roughness scales beyond this lead to a decline of the contact angle hysteresis due to the formation of a composite interface (where the liquid is unable to completely penetrate the surface).

The phenomena can be described by the Cassie-Baxter equation [START_REF] Tsibouklis | Poly(perfluoroalkyl methacrylate) Film Structures: Surface Organization Phenomena, Surface Energy Determinations, and Force of Adhesion Measurements[END_REF], where porous surface topography, superimposed to the roughness of a solid surface, causes air to become trapped in voids (i.e. prevents liquid from wicking) [START_REF] Cassie | Wettability of porous surfaces[END_REF]:

1 cos * cos - + ⋅ = f f ϑ ϑ (1)
Where: θ* is the apparent contact angle, f is the surface fraction (i.e. the total area of solid-liquid interface in a unity of plane geometrical area parallel to the surface), and θ is the contact angle of the rough surface but not porous. The surface fraction for a simple porous surface structure such as a columnar structure with square section is defined as follows (see Figure 1): Where: a is the pillar width and b is the spacing between two consecutive pillars (see Figure 1). By increasing the void between two micro structures (i.e. decreasing the width of the structures), a smaller surface fraction can be obtained (see Equation ( 2)) and therefore a larger contact angle (see Equation ( 1)) is achieved. A hydrophobic substrate can become superhydrophobic; under given circumstances even a hydrophilic substrate can act as a hydrophobic one.

DESIGN OF LOTUS STRUCTURES

In this workhoneycomb structures are used to and can be attained over a broad range of geometric parameters. To compare different wetting behaviors depending on the ratio between void and bulk characteristic lengths (see Figure 1), two different designs were investigated and manufactured (see Figure 2). The pitch of the micro-structured pattern was maintained constant (pitch = 4 μm), whereas the diameter of the combs and the thickness of the separating walls varied. The two patterns had the following geometrical characteristics:

1. Wall thickness = 1000 nm and comb diameter = 3000 nm (opening ratio f=0.25). 2. Wall thickness = 400 nm and comb diameter = 3600 nm (opening ratio f=0.10).

The height of the structures was 4 μm resulting on an aspect ratio of the sub-micro walls between 4 and 10. The height of the structures (and therefore the aspect ratio) was chosen in order to obtain superhydrophobic surfaces and also to allow the demolding after the replication by hot embossing, without damaging the separating walls.

The structured surface was of an area of 20x10mm² split into two different areas of 10x10mm² with a different diameter and pitch of the honeycomb structures. The two designs were placed as close as possible to each other to analyze an effect at the interface between both designs (see Figure 2). The process chain consist of the process steps E-beam lithography, mask fabrication, X-ray lithography, electroplating of mold insert and finally replication by hot embossing. (Figure 3) The design was written via electron beam (E-beam) lithography into a 3.2 µm thick photoresist (PMMA). The Ebeam device was operated at maximum acceleration voltage of 100 kV. In order to avoid damage of structures during the demolding step of the hot embossing process, vertical sidewalls on the mold insert have to be manufactured. Hence, to avoid any undercuts in the mold inserts already at the beginning of the process chain, the dose during E-beam lithography had to be optimized. Therefore variations in dose had to be performed to calibrate the E-beam process and to select the beam parameters for the writing procedure. Finally, writing was performed with a cu (s ferent honey Figure 4a: Intersection of the two dif comb designs after E-beam lithography Detailed view of the honeycomb structures after Figure 4b.

E-beam lithography. The thickness of the resist is 3 µm.

MASK FABRICATION

ickness of the gold mask was in the range of 2.20 ± 0.05 µm.

The generated polymer grooveswere filled with gold by electroplating (see Figure 5) and developed. To obtain a gold layer with a homogeneous absorption a gold sulfidic bath was employed. The th f = 0,15 f = 0,30

Figure 5 -Electroplated gold mask containing honeycomb sub-µ structures with two different aspect ratios. Height of the structures is 2.5 µm. A different surface fraction is obtained by decreasing the wall thickness from 1200 nm (aspect ratio = 2.1) to 600 nm (aspect ratio = 4.2) (left and right respectively) with a unit cell feature pitch of 4 µm.

X-RAY LITHOGRAPHY

The mask was used in a next step for the transfer of the design to a 4 µm and a 10 µm thick resist of PMMA. Here different aspect ratios of the structures up to 20 (referring to the sidewalls) were achieved. As shown in Figure 6, a deformation of the structures of 10 µm thickness occured due to a shift of the mask during the X-ray lithography process.

REPLICATION BY HOT EMBOSSING

The replication of the mold inserts into PMMA was carried out by hot embossing (Figure 9). Hot embossing is well suited for the replication of large structured areas with high aspect ratio sub-μ structures on a thin residual layer [START_REF] Heckele | Review on Micro Molding of Thermoplastic Polymers[END_REF].

Especially the short flow distances and the moderate flow velocities produce low stress in the molded part. Furthermore, the precise vertical demolding, guaranteed by the precise vertical guidance of the crossbars of the machine, is essential for successful demolding of the filigree lotus structures. Honeycomb structures in the submicron range with an aspect ratio of 3.5 were already replicated successfully by this replication technology [START_REF] Worgull | Modeling and optimization of the hot embossing process for micro-and nanocomponent fabrication[END_REF]. Nevertheless, because of the influence of shrinkage of the polymer, a challenging task is the demolding without deformation or damage of the structures. This holds in particular for the honeycomb structures with sidewalls in the range of 400 nm and structure height in the range of 10 µm.

The replication of high aspect ratio micro-nano structures into polymer with a very low defect rate is essential for the functionality of the superhydrophobic surface. The lotus structures can be replicated in a wide range of polymer materials, beginning from low temperature materials like PS or PMMA up to semi-crystalline high temperature materials like LCP or PEEK. For the first approach PMMA was employed. An example of sub-micron honeycomb structures molded in PMMA by hot embossing is shown in Figure 10. The structure size of these honeycombs is in the same range like as structures fabricated in this work.

SUMMARY

A new process chain for the replication of sub-μm structured surfaces into polymer was established. The tooling process includes E-beam writing, mask fabrication and X-ray exposure with subsequent nickel electroplating. Hot embossing was used to replicate on a polymer substrate the micro-nano honeycomb structures.

Surface sub-µm structuring can be employed to provide particular properties to the polymer surface for enhanced flow functionality suitable for microfluidics applications. In this research work, in particular, sub-μm honeycomb structures were used to provide hydrophobic properties to a hydrophilic polymer surface. Alternatively, hydrophobic substrates can be turned into super-hydrophobic by using the presented surface structuring.

Advanced design solutions employing this new technology may include micro fluidic systems whose surfaces incorporates structural gradients generating a driving force to move liquid samples along channel structures. Figure 12 shows a gradient design which will be proceed with the same process chain described above.

9- 11

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

Figure 1 -

 1 Figure 1 -Passive droplet activation by means of a linear gradient, i.e. a change of the surface fraction f along the moving direction obtained by different openings b on the surface's substrate.

Figure 2 :

 2 Figure 2: Design and arrangement of honeycomb structures

Figure 3 -

 3 Figure 3 -Process chain for fabrication of sub-micron Lotus structures by Ebeam lithography, mask fabrication, X-Ray lithography, electro forming and finally replication by hot embossing.

Figure 6 .

 6 Figure 6. Via X-Ray lithography the structures were transferred into 4 µm high resist (top picture) and 10 µm high resist (bottom picture).

Figure 7

 7 Figure 7 Electroplated microstructured Nickel shim. Nickel shims with a typical thickness of about 300 µm are well suited for the replication of nanostructures . Typical diameters are in the range of 4 inch or 6 inch.

Figure 8 .

 8 Figure 8. Honeycomb structures of the microstructured nickel mold insert. In the pictures the electroplated shim from the 4 μm thick resist is shown. The detail view on the bottom picture shows the rough surface on the top side of the structures as an result of the roughness of the adhesion and starting layer TiOx.

Figure 9

 9 Figure 9 Schematic view of the hot embossing process. The process is characterized by heating of a polymer film via heat conduction, a two step embossing cycle, a convective cooling and finally a vertical demolding.

Figure 11 .

 11 Figure 11. Different wetting behaviors due to the surface sub-micro structuring of the polymer substrate: honeycomb with wall thickness of 400nm (right), honeycomb with wall thickness of 1000nm (left).

Figure 12 :

 12 Figure 12: Lotus gradient design generating a driving force to move liquid samples along channel structures

7. ACKNOWLEDGEMENTS The presented research was performed within the framework of the European Network of Excellence "Multi Material Micro Manufacture: Technology and Applications" (4M) (European Community founding FP6-500274-1; www.4m-net.org) and within the activity of the Cross Divisional Project "MINAFLOT" (Micro-and Nano-structured Surfaces for Liquid and Gas Management in Microstructured Flowfields) supported by 4M and the European Community (Project no. FP6-500274-2). Furthermore the collaboration of the Polymer Technology Division (4M Work package 4) and of the Microfluidics Application Division (4M Work package 10) is gratefully acknowledged.


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  Effect of temperature and of dietary lipid source on female broodstock performance and fatty acid composition of the eggs of rainbow trout Geneviève Corraze, Laurence L. Larroquet, Gérard Maisse, D. Blanc, Sadasivam S. Kaushik To cite this version: Geneviève Corraze, Laurence L. Larroquet, Gérard Maisse, D. Blanc, Sadasivam S. Kaushik. Effect of temperature and of dietary lipid source on female broodstock performance and fatty acid composition of the eggs of rainbow trout. 4. International Symposium Fish nutrition and feeding, Jun 1991, Biarritz, France. hal-02777948


The maturation îs an important stcp as the gcnnination and the plant1er recoverics arc dcpcnding on the qualîty of thl.': somalie cmhryos. For the conifcrs. as for the angiospcnns. a great improvcrncnt was obtaincd by the addition of ah~l:i.sic acid (ABA) 10 the cultuw medium. This gave fisc 10 wdl~dcvdopcd somalie cmbryos with typical morphology (considcring th~zygotic embryo), synchronous developmenl and no precocious germination (1). Changes in content in ABA during maturation of hyhrid larch somalie embryos were followed using an enzyme immunoassay (ELISA).

Materials and Methods

Somatie embryogenic lines of hybrid larch (uJrix kaempferi x L. riecirilla) were ohlained in 1990 from immature zygotic emhryos and since suoculturcd evelY two weeks on hasal MSG mediulll as pn;viously dcscrihed (2). The maturation consislCd in the obtention of cOIylcdonary somali..: emhryos well-devclopcd. The maturation medium consisling in MSG medium supplemenled with 0.2 M sucrose. 1 ~M ISA and 60 ~M ASA (2). Artel' 3 weeks of culture; cotyledonary somatic emhryos wen: transferrcd to germination medium whicb is a growth regulalnr-frce MSG medium. After 5 woeks of germination, germinatcd somalie emhryos which devduped need1cs were considered as plantlcLS. For the hormonal analyses, the somatic emhryos wll1l.:h undcrwcnt cmhryogeny were removed al'ter 0, l, 2, 3, 4 and 5 wecks of maturation. Analyses using an enzyme immunoassay (ELISA) were realised as previously descrihcd (1).

Results

Developmcnt of cotylcdonary somalie embryos was greaùy improwd hy the pl"èSenCC 01" ASA 111 the culture mèdium (1). l-Iowever in hyhrid lareh. the extension of Ihe period 01" malllrJ.lioll in presence of ASA from ::\ wecks to 4 or 5 weeks oesultcd in a significant dCl:rcase ln hoth germination and plantlct frequcncics (figure 1). The highest germination rates (9()Ç() \Vere obtained for 3 wccks-old somatie embryos. In contrast prolongation of maturation pcJiod 10 5 wccks rcsu1ted in oeduced gCl1l1ination (38%); this was aiso obscrvcd for the plantlct rccovery :

the development of the aerial pans was reduced from 70 to 10% for respectively 1 and 5 wcàsnid somatic cmbryos. In presence of ASA in the culture medium, the ASA levcl of the sornatil: emhryos increased during their dcvelopmcnt 10 reach a maximum valllc at \Vcck 5 (15.p moUcmbryo. figure 1). As a consequence. wc assume that il was the levcl of ABA in p/mllo Ihal \Vas fèsponsihk for the inhihition.

Conclusion

ABA actcd in both a stimulatory and an inhibitory manncr. If ABA promoted the ohtention nf bigh quality snmalic cmhryos in large numbcrs. il also had a dcktcrious effeel on the suhscqw.:nt dcvelopmcnt. i.e. genninalion and plantlcl recovcries. The rcsults stresscd lhe importance of the maturation procedure and duration.

Literature

(1) LE LU MA and LABEL P (1994) • : germination rate at 28 days of culture;. plantlet raie at 28 days of clIltllrc.Bar.l. rcprescnt standard errors on the mean (n=5 for ABA. n=60 for rates).

  Comment citer ce document : Lelu-Walter, M.-A., Lointier, A., Gouez, M.L., Narboux, C., Cornu, D., Joseph, C., Label, P. (1995). Somatic embryogenesis : control of somatic embryo maturation in hybrid larch. In: Mechanisms and markers of regeneration and genetic stability. Extended abstracts (p. 12-13). Presented at 1. Meeting of the Cost 822 Working group, Prague, CZE (1995-04-27 -1995-04-30). document : Lelu-Walter, M.-A., Lointier, A., Gouez, M.L., Narboux, C., Cornu, D., Joseph, C., Label, P. (1995). Somatic embryogenesis : control of somatic embryo maturation in hybrid larch. In: Mechanisms and markers of regeneration and genetic stability. Extended abstracts (p. 12-13). Presented at 1. Meeting of the Cost 822 Working group, Prague, CZE (1995-04-27 -1995-04-30).

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  Comment citer ce document : Lelu-Walter, M.-A., Lointier, A., Gouez, M.L., Narboux, C., Cornu, D., Joseph, C., Label, P. (1995). Somatic embryogenesis : control of somatic embryo maturation in hybrid larch. In: Mechanisms and markers of regeneration and genetic stability. Extended abstracts (p. 12-13). Presented at 1. Meeting of the Cost 822 Working group, Prague, CZE (1995-04-27 -1995-04-30).

Figure 1 :

 1 Figure 1: Rclalionship octwoen cndogenous ABA levcls and gro\Vth rates of somatÎl' cmhryos.

  :.:;,/1 i (? -, ~l' Somatie cmhl)'ogcllcsis of hyhtid larch (Larix x ieptoeuropaea) bas hccn sucoessflllly dcvdnpcd al INRA of Orléans (France) sillec 1990.

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	Somatie embryogenesis	control of somatie embryo maturation ln
	hybrid lareh
	M•A. Lelu. A. Lointier, M.L. Gouez, C. Narboux, D. Cornu, C. Joseph,
	p, Label.
	LN.R.A., Station J'Amélioration des Arbres Forestiers, F-45160 Ardon. Franoe
	Aim

  . Physiol. PianI.. 92. 53-60. (2) LELU MA. BASTIEN C. KLiMASZEWSKA K. WARD C and CHAREST P (1'!94). Plant Ccii Tissue Organ Culture. 36. 107-115.