axial-flux permanent magnet machines for micro power generation

54 JOURNAL OF MICROELECTROMECHANICAL SYSTEMS, VOL. 14, NO. 1, FEBRUARY 2005

Axial-Flux Permanent Magnet Machines forMicropower Generation

Andrew S. Holmes, Member, IEEE, Guodong Hong, and Keith R. Pullen

Abstract—This paper reports on the design, fabrication, andtesting of an axial-flux permanent magnet electromagnetic gen-erator. The generator comprises a polymer rotor with embeddedpermanent magnets sandwiched between two silicon stators withelectroplated planar coils. Finite element simulations have beencarried out using ANSYS to determine the effects on performanceof design parameters such as the number of layers in the statorcoils, and the rotor-stator gap. The effect of including soft mag-netic pole pieces on the stators has also been studied. A prototypedevice with a diameter of 7.5 mm has been tested, and shownto deliver an output power of 1.1 mW per stator at a rotationspeed of 30 000 rpm. The generator has been integrated with amicrofabricated axial-flow microturbine to produce a compactpower conversion device for power generation and flow sensingapplications. [1265]

Index Terms—Axial-flow microturbine, energy harvesting, lasermicromachining, permanent magnet machine, power generator,power MEMS.

I. INTRODUCTION

THE proliferation of portable electronic devices in recentyears has led to increasing interest in alternative power

sources that might eliminate the need for chemical batteries. Themotivation for developing such sources depends on applicationarea and power requirements. For very low power applications,there is the possibility of powering electronics by energy “scav-enging” or “harvesting,” i.e., extracting energy from the sur-roundings. The power available is generally low (typically at the

level), but there is no need for the periodic recharging andreplacement associated with batteries, leading to lower mainte-nance. At higher power levels, there is interest in replacing bat-teries by fuel-burning devices that offer longer shelf life and/orhigher power density.

Energy scavenging power generators based on micro-electromechanical systems (MEMS) technology have beenresearched extensively over the past 10 yr. These have mostlybeen based on mass-spring-damper systems, where the damperis a transducer that converts mechanical energy to electricalenergy. Devices of this type have been developed based onelectrostatic [1], electromagnetic [2], [3], and piezoelectric [4]transduction mechanisms. Being resonant, these generators

Manuscript received February 3, 2004; revised June 1, 2004. This work wassupported by the U.K. Engineering and Physical Sciences Research Council, byGrant GR/N18895 “Microengineered Axial-Flow Pumps and Turbines.” SubjectEditor E. Obermeier.

A. S. Holmes and G. Hong are with the Department of Electrical and Elec-tronic Engineering, Imperial College London, London SW7 2BT, U.K. (e-mail:[email protected]).

K. R. Pullen is with the Department of Mechanical Engineering, ImperialCollege London, London SW7 2BX, U.K.

Digital Object Identifier 10.1109/JMEMS.2004.839016

work most efficiently when the applied excitation is a low am-plitude, reciprocating motion at a frequency close to resonance.This makes them potentially very effective for extracting en-ergy from vibrating machinery. Alternative devices containingonly a mass and a damper are also under development. Thesehave been shown to be more efficient in the case of large,low-frequency excitations such as might be experienced by aportable or wearable electronic device [5].

Work on fuel-burning MEMS power generators has beenfocused mainly on microturbines. The most highly developedwork in this area is that of MIT, where there is an extensive gasturbine program [6], [7]. This project aims to develop a fullyintegrated device that combines compressor, burner, turbine,and electrical generator in a single chip fabricated as a mul-tiwafer stack. Other workers have focused on turbogeneratorsubsystems. For example, in [8] a radial-flow turbine and anelectrical generator were combined on a single chip.

Recently at Imperial College, we have been developing a tur-bine-based power generator that combines an axial-flow turbinewith an axial-flux electromagnetic generator, as shown in Fig. 1.The idea of this device is that it should be able to extract powerdirectly from an externally generated gas flow, rather than re-lying on an internal burner and compressor. The axial-flow ge-ometry is the obvious choice for this kind of generator, becauseit can operate at low pressure ratios [9]. Axial-flow microtur-bines are difficult to realize by conventional microfabricationmethods, because to produce the necessary curved profiles onthe rotor blades and guide vanes [see Fig. 1(b)] it is necessaryto fabricate structures where the sidewall angle evolves withdepth in a controlled manner. We have solved this problem byusing excimer laser micromachining with a variable aperturemask to define the required blade profiles in SU8 polymer partspreformed by UV lithography. Applications for our device areenvisaged in areas such as wireless sensing, where the turbinecould be used to power a remote sensor and a short-range radiotransmitter. In the case of a flow sensor or an air-speed sensor,the turbine could also perform the sensing function, with theadvantage over other sensor types of generating a frequencyoutput.

This paper focuses on the electromagnetic generator sectionof our device, which is based on an axial-flux, permanentmagnet machine. The axial-flux geometry has been widelyused in larger scale motors and generators [10], and also morerecently in microfabricated motors [11]. It is particularly at-tractive for MEMS devices because it allows the use of planarcoils on the stators, which are easier to fabricate than solenoids.In the following sections, we consider in detail the design andfabrication of the generator. ANSYS simulations are presented

1057-7157/$20.00 © 2005 IEEE

HOLMES et al.: AXIAL-FLUX PERMANENT MAGNET MACHINES 55

Fig. 1. Microengineered turbo generator based on axial-flow turbine and axial-flux permanent magnet generator. (a) Schematic cross section through device.(b) Details of turbine section, showing rotor blade and guide vane shapes.

that show the effects on generator performance of designparameters such as the number of layers in the stator coils andthe rotor-stator gap. The effect of including soft-magnetic polepieces on the stators is also considered. The fabrication processis then described, and experimental results are presented for aprototype generator with a diameter of 7.5 mm. Finally, issuesassociated with future downscaling of the device are discussed.

II. DEVICE DESIGN

A. Basic Design Procedure

Fig. 2(a) is a plan view of the permanent magnet generator,showing just one of the pole pairs, while Fig. 2(b) is a crosssection through the same part of the device. The generator coil isa two-layer structure, with spiral coils arranged in an an-nulus in each layer. Other stator geometries based on overlappedcoils (see, for example, [12]) could be more efficient in terms ofpower generation, but were avoided because of potential processyield issues associated with producing many interlayer vias. Inour design, each spiral coil is connected at its outer edge to anadjacent coil in the same layer, and by a single via at the centerto the coil above or below it. For example, in Fig. 2, coils 1 and2 and coils 3 and 4 are connected by vias [shown in Fig. 2(b)],while coils 2 and 3 are connected by an in-plane link [shown inFig. 2(a)]. In this way, all the coils are connected in series, ex-cept at one point where the in-plane link is replaced by tracks toexternal contacts. The various design parameters are defined in

Fig. 2. Schematic of permanent magnet generator, showing design parameters.(a) Plan view. (b) Cross section along arc passing through coil centers.

Table I, where the values used in the first prototype devices arealso shown.

The overall scale of the first prototypes was set by the deci-sion to use conventional ball race bearings (BOCA type SMF681,

) and commercial neodymium boron iron (NdBFe)permanent magnets (CERMAG grade N30H). The bearing usedhad a flange with an outer diameter of 3.8 mm, and the value of

was chosen to allow a 200- clearance between the outeredge of the bearing flange and the inner radius of the coil annulus.The minimum radius for the magnet trajectory, based on the re-quirement that the magnets should lie inside the coil annulus formaximum induction, is , while the max-imum number of pole pairs at this radius is bounded roughly by

. When , ,these expressions give and . In termsof output power, it is generally advantageous to maximize thenumber of pole pairs. However, a lower number of pole pairs

was used to improve the mechanical robustness ofthe polymer rotors, and to facilitate rotor assembly. Onceand had been defined, the spiral coils were constructed, withthe number of turns being set by the interturn pitch and the coilsize in the circumferential direction, allowing for a via post at thecenter. A relatively large pitch of 60 was chosen for ease ofprocessing, giving an value of 12 turns per spiral.

B. Scaling Considerations

The induced output voltage of any permanent magnet gener-ator is given by , where is the total magneticflux linkage in the stator coils. For our device we can expressthe maximum flux linkage for each stator in the form

(1)


TABLE IDESIGN PARAMETERS FOR PERMANENT MAGNET GENERATOR, WITH VALUES USED IN PROTOTYPE DEVICES

where is the remnant flux density of the permanent magnets,is the area enclosed by the boundary of each coil, and is a

geometrical factor of order unity that depends on (a) the preciselayout of the turns in each coil, (b) the exact form of the fluxdistribution around each permanent magnet, and (c) the gapbetween rotor and stator. The factor is simply the numberof spiral coils on each stator. Assuming has a harmonic timedependence, as will be approximately true for any well-designedgenerator, we can then write the rms output voltage as

(2)

where is the angular rotation speed of the rotor. The availableoutput power per stator is then given by

(3)

where is the total winding resistance, which may be expressedas

(4)

In (4), is the conductivity of the coil metal (copper in ourcase), and and are the width and height of the coil tracks.The term represents the total track length of eachstator coil, being a second geometrical factor depending onthe coil shape and layout of turns.

Using (2)–(4), we can establish some basic scaling rules forthe generator. In particular, if the entire device is scaled in sizewithout any change in geometry, so that only , and areaffected by scaling, then the following rules will apply

(5)

where is any characteristic linear dimension. The scalingof output power imposes some severe limitations on the min-imum useful size for generators of this type. For example, inthis paper we demonstrate cm-scale devices with output powerlevels of about 1 mW. At constant maximum rotation speed,

downscaling of these devices by a factor of only four would re-duce the output power to around 1 , which is on the limitof what is useful with current electronics. The detrimental ef-fects of length-scaling can be offset by increasing the rotationspeed in smaller devices, but this is not feasible with conven-tional bearings.

Some care must be exercised when using the above equationsto explore the scaling effects of parameters other than devicesize, because the combined effects of , , and may not beas first expected. Moreover, (2) does not remain valid under allconditions. To illustrate this point we consider the case wherethe number of poles is varied by adding or removing permanentmagnets. In this case, assuming the stator coils are reshaped as

varies so as to keep the coil annulus filled, then changesin will be largely compensated by changes in , so that theproduct will remain roughly constant. Under these condi-tions (2) predicts . In practice, however, changes inthe shape of the induced voltage waveform as the number ofpoles is varied result in the actual scaling index for beingcloser to 1.5. Under the same conditions shows little varia-tion with (because ), so the availablepower scales roughly as .

C. Numerical Simulations

While (2)–(4) are useful in showing the functional depen-dence of output voltage and power on various design parame-ters, they cannot yield quantitative information unless the valuesof and are known. Such quantitative information is key tothe design of miniature power generators, particularly as and

are likely to be small, and there is a risk that the device maynot generate a useful level of output voltage at the design speed.For self-primed operation (i.e., in the absence of a permanentvoltage source such as a battery), it is helpful if the generatorcan produce 1 V pk-pk or more. Assuming the track width andheight are scaled in inverse proportion to (to maintain con-stant line/space and aspect ratios), the output voltage and powerscale as and respectively, so to maximizepower output we should choose the smallest value of thatwill satisfy any minimum voltage constraint.


Fig. 3. Flux distributions for a one-dimensional (1-D) array of permanent magnets, calculated using ANSYS. (a) and (b) Overall distributions with and withoutsoft magnetic pole pieces. (c) and (d) Effect of increasing the gaps between the permanent magnets and the pole pieces.

Calculation of is simply a matter of geometry once the coilshape and turn layout are known. A very approximate valuefor can also be obtained relatively easily, without solving themagnetic problem, simply by assuming that the magnetic fluxdensity is uniform and equal to in any coil region lyingdirectly over a permanent magnet, and zero elsewhere.is the flux density on the axis at the end of a long cylindricalmagnet. Given this assumption, numerical calculation of the fluxlinkage for the geometry in Fig. 2 is straightforward. This ap-proach is fast, and gives surprisingly good results in spite of thegross nature of the approximations involved.

For more accurate calculations, the actual flux density dis-tribution experienced by the stator coils must be known. Wehave carried out finite element analysis (FEA) simulations usingANSYS software to predict the variation of output voltage withthe rotor-stator gap , and to determine the effect on per-formance of the soft magnetic pole pieces. Similar calculationshave been reported previously for larger scale axial-flux gener-ators [13]. However, these are typically based on a slightly dif-ferent geometry, consisting of a single stator with a pair of cou-pled, permanent magnet rotors placed either side of it. This latterdesign is more efficient because it avoids eddy current losses inthe magnetic materials. However, it is more difficult to realizeby microfabrication, because it requires the stator coils to befabricated on a thin membrane.

In qualitative terms, it is expected that the soft magnetic polepiece will tend to reduce the lateral flux leakage from eachmagnet pole, thereby increasing the induced output voltage. On

the other hand, increasing the rotor-stator gap is expected to re-duce the output voltage, because the flux density will decay withaxial distance from the magnet poles. These effects are illus-trated in Fig. 3, which shows magnetic flux plots for an arrayof permanent magnets (a) without a pole piece, and (b-d) with apole piece at different distances. These plots were produced bya two-dimensional (2-D) simulation, but a three–dimensional(3-D) simulation would produce similar results in qualitativeterms.

The plots in Fig. 3 neglect the effects of induced eddy currentsin the pole pieces. Such currents will inevitably arise when-ever the rotor is moving if the pole pieces are conducting. Eddycurrents lead to power dissipation, reducing the generator ef-ficiency, so conducting pole pieces should be avoided if max-imum efficiency is to be achieved. Eddy currents also produce anadditional magnetic field that modifies the flux distribution bothinside the pole piece and, to a lesser extent, in the rotor-statorgap. However, the change in the output voltage amplitude dueto this effect is less than 1% for our device at a rotation speedof 30 000 rpm. Consequently we have ignored eddy currents inour output voltage simulations.

Fig. 4 shows the ANSYS model set up for full 3-D simu-lation. Only the permanent magnets and the pole pieces wereincluded in the model; all other parts of the structure—statorsilicon, SU8 polymer layers, and copper coils—were assumedto have and modeled as air. Based on the manufac-turer’s data, the permanent magnets were assumed to have a

value of 1160 mT, and a coercivity of 880 kA/m, while the


Fig. 4. Physical model of permanent magnet generator used in ANSYS calculations. Silicon, copper and SU8 are assumed to have � = 1, and are treated as air.Generator output is assumed to be open circuit.

Fig. 5. Calculated variation of output voltage with rotor-stator gap, with andwithout pole pieces. Y axis is peak-to-peak induced voltage in mV for one stator,divided by rotation speed in rad=s.

soft magnetic material was assumed to be linear with a rela-tive permeability of 600 (Nickel). The magnetic problem wassolved using the Vector Potential method [14]. This calculationwas done once for each rotor-stator gap. The flux linkage for asingle coil was then determined for a series of angular positions,and differentiated numerically to determine the rate of changeof flux linkage with angle. From this function the output voltagefor the complete stator was calculated.

The simulation results are shown in Fig. 5, where the peak-to-peak output voltage per stator has been plotted as a function ofrotor-stator gap, with and without a soft magnetic pole piece.Note the output voltage is normalized to unit angular rate, andexpressed in units of . The corresponding values,inferred from (2) (assuming a sinusoidal waveform), are shownon the right-hand y axis. It is clear from Fig. 5 that the outputvoltage falls off steeply with increasing rotor-stator gap, and that

the rate of this decay is not affected significantly by the pres-ence of the soft magnetic pole piece. The pole piece enhancesthe output voltage by roughly 30% at a gap of 100 , corre-sponding to a 70% increase in output power. This effect gradu-ally becomes less pronounced as the gap increases. The outputvoltage per stator at a speed of 30 000 rpm (nominal shaft speedfor prototype turbine) for a generator with a pole piece and arotor-stator gap of 120 is expected to be 1.88 V pk-pk. Witha stator winding resistance of 45 , this corresponds to an avail-able power of 2.45 mW, which is more than adequate for powergeneration in many low-power applications. The above windingresistance value is calculated using the parameters in Table I, as-suming copper coils with a conductivity of .

III. FABRICATION

In this section, we briefly describe the processes used to fab-ricate the rotors and stators for the first prototype generators.These rely on a combination silicon deep reactive ion etching(DRIE), multilevel electroplating, SU8 processing, and lasermicromachining.

A. Stator Fabrication

The stators were produced by double-sided processing of4 -dia, 500- -thick silicon substrates. First, 300- -deepannular cavities were etched in the backside of the wafer byDRIE using a photoresist mask. After reoxidation and sputterdeposition of a Cr/Cu seed layer, the cavities were filled withnickel by electroplating [see Fig. 6(a)]. A dry-film photoresistmask was used to confine the electrodeposition to the cavities.

In a second process step, two-layer planar coils were fabri-cated on the front side of the wafer by multilayer UV lithog-raphy and copper electroplating [Fig. 6(b)]. The insulation layerbetween the two layers of each coil, and a top protection layer,were made by SU8 photolithography. When defining the coils,


Fig. 6. Stator cross sections. (a) After backside processing. (b) Afterfabrication of two-layer coils. (c) After through-wafer etching for fluid channel.

Fig. 7. SEM photograph showing top layer of a stator coil prior to depositionof the protective SU8 layer.

an overplating method based on thin photoresist masks was usedto form mushroom-shaped tracks. These were coated by evap-oration with Cr to improve the adhesion of subsequent SU8layers. The coil tracks were nominally 12 high, with a widthof 30 and a pitch of 60 . Fig. 7 shows an SEM photo-graph of a finished coil prior to deposition of the top protectionlayer.

MEMS coil fabrication processes have been reported previ-ously for numerous other devices such as magnetic sensors [15],[16], magnetic actuators [17], micromotors [18], inductors, andtransformers [19]. The problem was complicated in our work bythe fact that parts of the wafer surface had to be left free of SU8at the end of the coil fabrication process to allow through-waferetching by DRIE. This was essential for the intended integrationof the generator with an axial-flow turbine. In early prototypes,windows were opened in the SU8 following coil fabrication byplasma etching using an electroplated nickel mask [20]. Sub-sequently, however, a more efficient process was developed inwhich the windows were defined lithographically in each suc-cessive layer of SU8 during coil fabrication. This required the

Fig. 8. Rotor fabrication process. (a) After deposition and patterning ofCu sacrificial layer. (b) After two-layer SU8 lithography. (c) After sacrificialrelease.

second level of coil plating to be carried out on a surface withrelatively extreme topography, but eliminated the need for anSU8 plasma etching step. Both processes were shown to be ca-pable of fabricating viable two-layer stator coils.

B. Rotor Fabrication

The rotors were fabricated by a combination of SU8 lithog-raphy and laser micromachining. The lithography stage wasperformed on 4 -dia silicon wafers on which a 10- -thickpatterned copper sacrificial layer had previously been defined[Fig. 8(a)]. Two layers of SU8, each 500- -thick after pro-cessing, were deposited and patterned to produce a 1-mm-thickrotor structure with cylindrical cavities for the permanent mag-nets and a stepped cavity at the center to accommodate the shaft[Fig. 8(b)]. After the second layer of lithography, the rotorswere released from the wafer by wet etching of the sacrificiallayer [Fig. 8(c)].

The turbine rotor blades, which required curved profiles, wereproduced by excimer laser micromachining with a dynamicallyvariable mask. Crosslinked SU8 is well suited to machining bylaser ablation at 248 nm wavelength, producing relatively cleanstructures with good surface finish. With the setup available, theblades had to be machined individually, with two machining op-erations being required for each blade, one for the convex sideand another for the concave side. A fixed mask was used to pro-tect other regions of the rotor from exposure, and a moving maskwas translated, under computer control, in front of the fixedmask so that different parts of the blade received different expo-sures. This process has been described in detail elsewhere [21].Fig. 9 shows an SEM image of a rotor after the laser machiningprocess.

Following laser machining, permanent magnets were manu-ally inserted into the cavities in the rotor, and secured in placewith SU8 which was applied using a needle and then cured byUV exposure and heating. For rotors to be used in turbo gen-erator tests, a precision-machined steel shaft was also attachedby a similar method. Each turbo generator was assembled ona metal jig with integral pins to align the two stators. An SU8spacer was used to define the gap between the stators, as shownin Fig. 1(a).


Fig. 9. SEM photograph showing an SU8 rotor after laser micromachining ofthe turbine blades.

Fig. 10. Test setup used to measure the variation of output voltage withrotor-stator gap.

IV. CHARACTERIZATION

Fig. 10 illustrates the setup used for characterization of proto-type generators. For these tests, the rotor was driven by a smalldc motor. A soft magnetic disc was attached to the motor shaft,and magnetic forces due to the embedded permanent magnetswere used to hold the rotor in position against the disc. Themotor was mounted with its axis horizontal, and positioned ad-jacent to a vertically mounted stator. The axial position of thestator could be adjusted by means of a micropositioner. Fig. 11is a captured oscilloscope trace showing the time variation ofthe generator output voltage at a rotation speed of around 7200rpm. The time variation is close to sinusoidal, with very lowharmonic distortion, confirming that the assumptions made inderiving (2) are valid for our device.

Fig. 12 compares the measured and simulated variations ofoutput voltage with rotation speed for stators with and withoutsoft magnetic pole pieces. The rotor-stator gap was 120throughout. In all cases the output voltage is proportional to ro-tation speed, as expected from (2). Furthermore, the enhance-ment due to the pole piece is similar in the experimental andsimulated cases, being 33% in the former and 27% in the latter.However, the simulated voltages are consistently higher thanthe experimental ones, the ratios of simulated to measured volt-ages being 1.53 and 1.61, respectively, for the stators with andwithout pole pieces. This discrepancy is attributed primarily tolower than expected magnetization in the permanent magnets.Evidence for this comes from remnance measurements madeon a random sample of ten magnets taken from the same batch

Fig. 11. Oscilloscope trace showing output waveform from a single stator.

Fig. 12. Comparison between simulated and measured variations of outputvoltage with rotation speed for stators with and without soft magnetic polepieces. Rotor-stator gap is 120 �m in all cases.

as those used in the prototypes. The ten magnets analyzed hada mean Br value of 767 mT with a standard deviation of 46 mT.The mean value is lower than the manufacturer’s stated Br of1160 mT by a ratio of 1.51, which is very close to the sim-ulated-to-measured voltage ratio of 1.53 obtained for the de-vice with pole pieces. The small difference between the sim-ulated-to-measured voltage ratios observed with and withoutpole pieces is probably due to differences in the rotor-stator gap,which was measured only to within .

Fig. 13 compares measured and simulated variations of thenormalized output voltage with rotor-stator gap for a stator witha soft magnetic pole piece. In each case, the voltage has beennormalized to give a value of 1 at the minimum gap of 120 .The extremely close agreement between the simulated and ex-perimental trends lends further support to the suggestion that thediscrepancies in the absolute voltage levels in Fig. 12 are due toa lower value rather than, for example, geometrical discrep-ancies between the simulated and the experimental devices.

In addition to the generator tests reported here, preliminarytests of completed turbo generators have been carried out, usinga test jig for measuring the induced voltage and pressure dropas a function of nitrogen flow rate. Operation at pressure dropsdown to a few mbar has been demonstrated, indicating that the


Fig. 13. Comparison between simulated and measured variations of normal-ized output voltage with rotor-stator gap for a stator with a soft-magnetic polepiece.

Fig. 14. Measured variations of rotation rate and output voltage for a prototypeturbo generator driven by nitrogen.

devices are well suited to extraction of power from ambient gasflows. The results of these tests have been reported in detail in aseparate paper [22]. Fig. 14 shows the measured variation withnitrogen flow rate of the rotation speed and output voltage fora typical device. This particular turbine achieves a rotation rateof 30 000 rpm at a pressure drop of around 8 mbar and a flowrate of about 35 L/min. At 30 000 rpm the open-circuit voltageper stator is 1.19 V pk-pk, corresponding to an available powerof 1.1 mW per stator into a matched (40 ) load. This generatorperformance is marginally worse than the best achieved in thegenerator tests (290 mV at 7000 rpm, corresponding to 1.24 Vat 30 000 rpm) because the assembly method used in the com-pleted devices did not allow rotor-stator gaps down to 120 .

V. DISCUSSION

This paper has presented design calculations, simulations andinitial experimental results for a mm-scale axial-flux electro-magnetic generator aimed at easy integration with an axial-flowmicroturbine. A fabrication process has been developed that in-volves deep silicon etching and multilayer electroplating for thestator parts, and SU8 processing for the rotors. Experimentalmeasurements on prototype generators show close agreement

Fig. 15. Variation of output power with number of coil layers, assumingGrc = 120 �m and Gcc = 25 �m. Y axis is normalized to output power fortwo layers.

with ANSYS simulations in terms of trends, although the mea-sured output voltages are consistently lower than the simulatedones by around 35%. This discrepancy is due to the average rem-nant flux density of the permanent magnets being lower thanstated in the manufacturer’s data.

The mW power levels generated by the prototype devicesare more than adequate for many remote sensing applications,and the output voltages are sufficient to allow self-primed op-eration. Higher output powers could be achieved from devicesof the same size and basic design by a number of routes, forexample by reducing the rotor-stator gap, enhancing the coilfill-factor (ratio of copper to dielectric in the stator coils), in-creasing the number of poles, or increasing the number of coillayers on each stator. All of these changes would make fabrica-tion or assembly of the device more difficult, and each wouldafford only a modest increase in output power. However, thecombined effect could be significant. For example, a device sim-ilar to ours but with a rotor-stator gap of 25 (compared to120 ), a coil fill factor of 50% (compared to 25%) and eightpoles (compared to 5) would generate around 9 mW per statorat 30 000 rpm. Increasing the number of coil layers while main-taining the same interlayer coil spacing could achieve a furtherthreefold increase in power, as illustrated in Fig. 15. In the ab-sence of flux leakage, the output power would be proportionalto the number of coil layers. In practice, however, beyond abouteight coil layers the gap between any additional coils and therotor is so large that flux leakage renders them ineffective. Thus,a device with an output power of around 25 mW per stator at30 000 rpm should be achievable without changing the basicconfiguration. The main factor limiting the output power of sucha device would be the maximum operating speed of the ballbearings.

Commercial permanent magnets and bearings were used inthe first prototypes to shorten the development cycle. In futurework, we will investigate the use of permanent magnets that canbe deposited at wafer level prior to sacrificial release of the rotor.This will eliminate the manual assembly and bonding operationsassociated with the current design; no other changes to the rotorfabrication process will be required. Printed magnets based ona powdered magnetic material in a polymer binder will prob-ably be used here [23], as these appear to be capable of higher


remnant flux densities than electroplated materials [24]. We willalso explore the possibility of replacing the ball bearings withmicrofabricated aerodynamic bearings. Significant progress onbearings of this type has been reported by other workers in re-cent years [7]. With these two changes, downscaling of the de-vices will become a possibility. As noted earlier, scaling lawsdo not favor a significant reduction in size. However, a devicesmaller by a factor of 2 or 3 (in all linear dimensions) couldstill yield output power levels useful for some sensing applica-tions, particularly if downscaling could be accompanied by anincrease in rotation speed.

ACKNOWLEDGMENT

The authors are grateful to Dr. J. Stagg of Imperial Collegefor his help with DRIE.

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Andrew S. Holmes (M’02) received the B.A. degreein natural sciences from Cambridge University,U.K., in 1987, and the Ph.D. degree in electricalengineering from Imperial College London, U.K., in1992.

He is currently a Reader with the Optical andSemiconductor Devices Group, Department ofElectrical and Electronic Engineering, ImperialCollege London. His research interests are in theareas of micropower generation and conversion,MEMS devices for microwave applications, and

laser processing for MEMS manufacture.

Guodong Hong received the B.Eng. and M.Eng. de-grees in material engineering from Shandong Uni-versity, China, in 1984 and 1989, respectively, andthe Ph.D. degree in mechanical engineering from Ts-inghua University, China, in 1998.

He is currently a Research Associate with theOptical and Semiconductor Devices Group, De-partment of Electrical and Electronic Engineering,Imperial College London. His areas of researchinterests include the micropower generation andconversion, micromoulding, rapid prototyping, and

manufacturing for microelectromechanical systems.

Keith R. Pullen received the first degree in 1987and the Ph.D. degree in 1990 from Imperial CollegeLondon, U.K., the latter being funded on full salaryby Rolls-Royce plc for investigations into smallhigh-speed turbines and generators.

He joined Noble Denton in 1990, where he was inthe field of offshore engineering working in variouslocations, including Norway. In 1992 he returned toImperial College as a Lecturer to continue researchinto small high-speed electrical machines and was acofounding member of the Turbo Genset Company

being employed as Chief of Design. He has undertaken engineering consultancywork for more than 15 companies including BP, BOC Edwards, Honeywell, andNortel Networks in turbomachinery. He is an author of 50 papers in interna-tional journals and conferences and has 21 patents granted or pending. He iscurrently a Senior Lecturer and has recently been appointed Director of Engi-neering for Hydroventuri Ltd. As well as teaching subjects such as gas turbinetechnology, he manages a group of eight researchers and an experimental labo-ratory of 250 m , including several large facilities.

Dr. Pullen earned three prizes, including top undergraduate sponsored andgraduate, while with Rolls-Royce.

axial-flux permanent magnet machines for micro power generation

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