mems-based power disconnect for 42-v automotive power systems

Ma

HJWUD

5TE

1

FtcsdthTbmtgpHvdecsT�

pscM

1

J. Micro/Nanolith. MEMS MOEMS 7�1�, 013010 �Jan–Mar 2008�

J

Downloaded F

EMS-based power disconnect for 42-Vutomotive power systems

enry K. Chuames K. Millsilliam L. Cleghornniversity of Torontoepartment of Mechanical and Industrial

EngineeringKing’s College Road

oronto, Ontario, Canada M5S 3G8-mail: [email protected]

Abstract. We present the design of a micro-electro-mechanical-system�MEMS�–based power disconnect for use with the proposed 42-V auto-motive power systems. High-voltage power systems are prone to arcing,which occurs during power disconnections. The high arcing current maylead to severe damage on the systems and may expose electric shockhazards to humans. The objective of the MEMS-based power disconnectwe propose is to eliminate arcing occurrence in the systems. To elimi-nate arcing, one alternative is to electronically terminate the power sup-ply to the system prior to the physical disconnection. The integratedMEMS force sensor on the power disconnect will be activated as a ser-vice technician disconnects the connector. The power supply to the sys-tem will be electronically shut off to prevent arcing during the physicalinterruption. The MEMS force sensor on the disconnect has an overalldimension of 3600 �m�1000 �m�10 �m and is fabricated with theMicragem fabrication process. A displacement reduction mechanism isincorporated into the sensor design to increase the sensitivity of theforce sensor. Results show that the sensor is capable of measuring amaximum force input of 10.7 mN, resulting from a 20-�m displacementon the sensing structure. © 2008 Society of Photo-Optical Instrumentation Engineers.�DOI: 10.1117/1.2896107�

Subject terms: microelectromechanical systems; sensors; integrated circuits;noise.

Paper 07055 received Jul. 31, 2007; accepted for publication Dec. 4, 2007; pub-lished online Mar. 20, 2008.

Introduction

rom indicator lights on a dashboard and power door lockso more advanced antilock braking systems and tractionontrol systems, the battery in an automotive vehicle is theole power source to all of the electrical components andevices inside the vehicle. However, as more and more ofhese electrical devices have been installed onto the ve-icle, the battery power demand has gradually increased.he present 200 A /3 kW, 14-V battery systems will soone insufficient to satisfy increasing automobile power de-ands. Hence, the automotive industry proposes increasing

he battery voltage from 14 V to 42 V for the next-eneration vehicle. This new battery standard will be ca-able of providing electrical power up to 200 A /9 kW.owever, one major challenge with such a high batteryoltage is that there will be an increased possibility of arcischarge during physical connection and disconnection oflectrical devices in the vehicle. The high arc dischargeurrent may cause severe damage to the terminals and theystems and may pose electric shock hazards to humans.1,2

o resolve the problem, a micro-electro-mechanical-systemMEMS�–based power disconnect system is proposed.

MEMS devices have become one of the fastest growingroducts in the automotive industry. The advantages ofuch devices are their small physical size and low poweronsumption. Many modern vehicles are now installed withEMS devices to enhance the comfort and safety of the

537-1646/2008/$25.00 © 2008 SPIE

. Micro/Nanolith. MEMS MOEMS 013010-

rom: http://nanolithography.spiedigitallibrary.org/ on 05/17/2015 Terms of U

vehicles. For instance, MEMS-based pressure sensors maybe installed in the tire pressure monitoring system �TPMS�to consistently monitor the pressure level inside the tire.MEMS-based accelerometers may be found in some airbagdeployment systems for crash detection.3

The power disconnect system proposed in this paperconsists of a MEMS sensor, a readout circuit, and a relay.Figure 1 shows a schematic of the power disconnect sys-tem. The MEMS sensor may be installed onto the custom-made package cover described in Ref. 4. The cover servesas an interface for the force input as well as a mechanism totransform the macroscopic input force imparted by a ser-vice technician into a fine motion on the order of a fewmicrons to be applied to the sensor. Preliminary resultsshow that a grasping displacement of 30.86 �m will resultin response to a 60-N macroscopic force input.4 In the pres-ence of human contact, the disconnect system will trigger asignal to an electronic relay to electronically terminate thepower supply to the electric load. As a result, arcing willnot occur during the physical interruption of the connectorterminals.

The MEMS sensor installed on the power disconnectsystem is a capacitive-type force sensor. The capacitivesensing principle is selected for this MEMS-sensor applica-tion because of its simple structure, high sensitivity, lowtemperature dependence, and low power consumption.5

Over the past several years, researchers have investigateddifferent capacitive-type sensors for force and strain mea-surement. Despont et al.6 micromachined a capacitive force
sensor. This sensor could measure force between 0.01 N
Jan–Mar 2008/Vol. 7�1�1

se: http://spiedl.org/terms

afitdgrbbtittptsh

chsMstn5ir

2

2FboRstl2ps

rtmpd

Chu, Mills, and Cleghorn: MEMS-based power disconnect for 42-V automotive power systems

J

Downloaded F

nd 10 N. The differential configuration of the comb-ngers in the sensor helped to improve linearity and sensi-

ivity. Sun et al.7 proposed a capacitive 2-degrees of free-om �DOF� force sensor for micro-force measurement. Theroup used six sets of comb drives in an orthogonal ar-angement to resolve forces in the x and y directions. Ae-ersold et al.8 published a capacitive sensor design forending strain measurement. Experimental results showedhat this version of capacitive sensor was capable of detect-ng over 1600-�� strain. In general, the preceding capaci-ive sensor designs employ the space-variation type capaci-ive sensing principle. This principle has the advantage ofroviding high output capacitance. However, the sensitivityends to decrease as the force or displacement to be mea-ured increases. To resolve this issue, a MEMS-based,igh-sensitivity, capacitive sensor is proposed in this paper.

This paper mainly presents the design of a MEMS-basedapacitive sensor and its associated readout circuitry forigh-sensitivity force detection. Section 2 provides a de-cription of the sensor design and its fabrication process.odeling equations of the capacitive force sensor are pre-

ented in Sec. 3. Section 4 describes the details regardinghe design of the readout circuit. Results from the signaloise analysis conducted in this study are discussed in Sec.. Design of the experiments and the results will be coveredn Sec. 6. The last section of this paper summarizes theesults and findings obtained from this study.

MEMS Capacitive Force Sensor

.1 Sensor Architectureigure 2 shows the architecture of the proposed MEMS-ased space-variation type capacitive force sensor. Detailsf the development of this sensor design are described inef. 9. The overall dimensions of the entire MEMS sensor

tructure are 3600 �m�1000 �m. The sensor consists ofwo identical comb-drive mechanisms �labeled #1 and #2�ocated at the center part of the sensor, as shown in Fig.�a�. The comb-drive mechanisms will alter the sensor ca-acitance in response to an external force imparted onto theensor.

This design of the capacitive force sensor also incorpo-ates a displacement reduction mechanism. The outer struc-ure of the sensor in Fig. 2�a� is the displacement reduction

echanism. The mechanism is designed to convert the im-arted displacement from the input force in the vertical

Fig. 1 Schematic of the MEMS-based power disconnect system.

irection to a smaller lateral displacement onto the comb-



drive mechanisms. The presence of such a mechanismhelps prevent damage of the comb-drive mechanism andimproves the sensor sensitivity.

2.2 Comb-Drive MechanismEach of the two comb-drive mechanisms consists of 59 setsof movable electrodes for force sensing. Each comb-fingeris 60 �m in length and 4 �m in width, with an overlappingarea of 560 �m2. The comb-drive mechanisms detect theinput force imparted onto the sensor through the changes inthe air gap between the comb-fingers. When a force is ap-plied vertically onto the top and bottom beams of the sensorstructure, the comb drive will experience a reduced relativemotion in the lateral direction, as illustrated in Fig. 2�a�. Asa result, the gap between the comb-fingers changes, as il-lustrated in Fig. 3. Due to the differential comb-drive con-figuration employed in this sensor, the capacitance value onone side decreases while the capacitance value on the otherside increases. By measuring the change in the capacitancevalue, the applied force can be determined. Details of thedetection principle will be discussed subsequently.

Fig. 2 Schematic of the sensor: �a� top view; �b� side view.

Fig. 3 Details of the force sensors illustrating the capacitance de-crease due to a force input: �a� sensor at original position; �b� sensor
at new position.
Jan–Mar 2008/Vol. 7�1�2


2MMmmpTsatFt1

tscf

S

Tttrrma

Fi


J

Downloaded F

.3 Displacement Reduction Mechanismotion amplification mechanisms are often incorporated inEMS-based actuators to increase the magnitude of theotion generated from these microdevices.10,11 A similarechanism can be designed to reduce the displacement ex-

erienced by the comb drive in response to a force input.he MEMS sensor proposed in this paper has been de-igned with a 10:1 displacement reduction mechanism tochieve this goal. When a force is applied vertically ontohe top and bottom beams of the sensor, as demonstrated inig. 4, the relative motion in the lateral direction between

he center mass and the comb-finger sidebars is reduced to/10 of the vertical input displacement.

With the proposed displacement reduction mechanism,he sensitivity of the sensor is substantially improved. Theensitivity of a capacitive force sensor is equal to thehange in the capacitance value, C, over the change in theorce, F, or displacement, D, input, and thus,

ensitivity =�C

�F=

�C

�D. �1�

he capacitance value of the sensor is inversely related tohe initial gap distance and the sensor displacement. Withhe proposed mechanism design, the sensor displacement inesponse to a force input is reduced. Hence, it is possible toeduce the initial gap distances by a factor of 10. With thisodification, the sensitivity is improved by the same factor

s well.

Fig. 4 Schematic of the displacement reduction mechanism.

ig. 5 Comb-drive mechanism of �a� the original design and �b� the
mproved sensor design.


In addition, as the sensor gap distances are reduced,more comb-fingers can be integrated into the design. Forthis sensor design, it is found that the sensor capacitance aswell as the sensor sensitivity can be increased by a factor of5.7 with this modification. Figure 5 shows a schematic ofthe improvement made on the comb-drive mechanism de-sign with the displacement reduction mechanism.

The overall sensitivity of the sensor with this reductionmechanism is approximately 57 times higher than the de-sign without the reduction mechanism.

2.4 Sensor FabricationThe sensor design has been fabricated using a multiuser,silicon-on-insulator �SOI�–based MEMS fabrication pro-cess called Micragem.12 The fabrication process startedwith a 500-�m-thick glass wafer as the base substrate. Theglass surface was patterned and etched to form 10-�m-deepcavities using wet etching. Single crystal silicon �SCSi�with 10-�m thickness was used as the structural layer andwas anodically bonded to the glass substrate, leaving onlythe area above the cavities to be suspended. A thin metallayer was deposited on top of the silicon surface for betterelectrical conductivity. Last, the sensor structure was pat-terned and etched using deep reactive ion etching �DRIE�.Figure 6 shows an image of the fabricated MEMS sensorunder the optical microscope and the sensor parameters aresummarized in Table 1.

Table 1 Parameters of the capacitive sensor.

Description Value

Sensor Dimension 3600 �m�1000 �m�10 �m

Number of electrodes 118 sets

Electrode length 60 �m

Electrode width 4 �m

Electrode thickness 10 �m

Overlap area 560 �m2

Nominal gap distance 3 �m and 6 �m

Maximum input displacement 20 �m

Displacement reduction ratio 10:1

Fig. 6 Image of the MEMS sensor under an optical microscope.

Jan–Mar 2008/Vol. 7�1�3


3

3Iitt

C

wa

iosccmsc

Cg

C

C

wip

ettcsois


J

Downloaded F

Sensor Modeling

.1 Sensor Capacitancen space-variation-type capacitive sensors, the capacitances inversely proportional to the gap that separates the elec-rodes. The value of the capacitance can be evaluated usinghe parallel plate equation:

=�A

g, �2�

here � is the relative permittivity, A is the overlappingrea, and g is the gap distance between the electrodes.

Considering the existence of nonlinear capacitive behav-or, the force sensor proposed in this paper employs the usef a differential-style comb-drive configuration. Figure 7hows a schematic representation of a differential-styleomb-drive configuration. The advantage of this type ofomb-drive configuration is that for small sensor displace-ent, nonlinear effects are negligible and a linear relation-

hip between sensor displacement and output capacitancean be achieved.13

With the notation in Fig. 7, the capacitance value,increase and Cdecrease can be evaluated by considering bothaps gS and gL on the two sides of the electrode:

increase = n� �A

�gS − x�+

�A

�gL + x�� , �3�

decrease = n� �A

�gS + x�+

�A

�gL − x�� , �4�

here n is the total number of comb-finger sets, g is thenitial gap distance on the left and right, and x is the dis-lacement of the sensor.

For this sensor design, the sensor output capacitance isqual to the difference between Cincrease and Cdecrease. Whenhe sensor is idle, it has a capacitance of zero. The capaci-ance value increases as a force is applied on the sensor,ausing a displacement of the comb-fingers. The relation-hip between the capacitance value and the gap distance forne set of the comb-finger is plotted in Fig. 8, with thenitial gap distances of gL=6 �m and gS=3 �m. It is ob-

Fig. 7 Detailed view of the comb drive.

erved that when the sensor displacement is less than



1 �m, the relationship between comb-finger displacementand its resulting capacitance value is relatively linear, witha slope of 0.8 fF /�m, as shown in Fig. 8.

3.2 Spring ConstantIn this capacitive sensor, the input force applied on thesensor deforms the sensor. When the input force is small,the relationship between the force and the deformation canbe expressed using Hooke’s Law:14

F = kx , �5�

where F is the input force, k is the stiffness of the sensor,and x is the deformation.

Simulation from finite element analysis software, COM-SOL, shows that the sensor has a stiffness of 537 N /m inthe input direction and a stiffness of 4.75 kN /m in the out-put direction.

3.3 Resonant FrequencyThe resonant frequency of the sensor is crucial in determin-ing factors such as the operating bandwidth and is esti-mated as:

�o = � k

mef f�1/2

, �6�

where meff is effective mass of the structure.The effective mass of the structure represents the actual

portion of the overall mass attributed to the structure vibra-tion. For a fixed-fixed cantilever beam, the effective masscan be evaluated as:15

mef f =mL�0

Ly�x�2dx

yF2 , �7�

where mL is mass of the sensor per unit length, y is thedisplacement of the beam, and yF is maximum displace-ment of the sensor.

For the sensor proposed in this paper, the mass of thestructure is estimated to be 9.68�10−9 kg. The effectivemass is calculated to be approximately 0.37 of the overallmass, which is 3.58�10−9 kg. The resonant frequency is

Fig. 8 The capacitance value versus gap distance for one set of thecomb-finger.

found to be 61.6 kHz.

Jan–Mar 2008/Vol. 7�1�4


3

Tddn

idMbAfipa

b

wss

piwtootdl

b

waTo

s

b

Af


J

Downloaded F

.4 Damping

wo types of damping mechanisms are active as the combrives experience displacement: slide-film and squeeze-filmamping. Figure 9 illustrates these two damping mecha-isms.

Slide-film damping occurs at a result of a sliding motionn a viscous medium. The damping action will exert aamping force against the sliding motion. For the proposedEMS sensor, slide-film damping occurs at the air gap in

etween the comb-drive structure and the glass substrate.s the structure displaces, a sliding motion and hence slide-lm damping will result. At low frequency, the velocityrofile of the fluid underneath the moving plate is linear,nd the damping constant is expressed as:16

slide =�A

h, �8�

here � is the viscosity of medium, A is the area of theensor mass, and h is the oil layer thickness between theensor mass and the substrate.

Squeeze-film damping takes place when two parallellates in a viscous medium displace relative to one anothern the normal direction. The lateral movement of the sensorill deflect the comb-fingers and reduce the gap between

he comb-fingers. The air within the gap will be squeezedut of the gap to other areas. Due to the air viscosity, anpposing force is generated with the force proportional tohe velocity of the sensor mass. For a small comb-fingereflection, the squeeze-film damping constant can be calcu-ated as:17

squeeze =��W/L�LW3

g3 , �9�

here W and L are the width and length of the comb-finger,nd ��W /L� is a function of the aspect ratio of the finger.he value of ��W /L� can be obtained from Fig. 10 basedn the width and length of the comb-finger.17

The total damping factor of the sensor, b, is equal to theum of the squeeze-film and slide-film damping:

= bslide + bsqueeze. �10�

t a temperature of 25 °C, the sensor has a total damping−6

Fig. 9 �a� Slide-film damping and �b� squeeze-film damping.

actor of 3.15�10 Ns /m.



3.5 Quality FactorThe quality factor, Q, is the ratio between the total energystored in the system to the energy lost per cycle due to thedamping effects present. The quality factor of the system isgiven as:18

1

Q=

1

Qa+

1

Qs+

1

Qi, �11�

where Qa is the energy lost to the surrounding fluid, Qs isthe energy dissipated through the supports to the surround-ing solid, and Qi is the energy dissipated within the internalstructure.

The energy lost to the surrounding fluid is found to bethe dominant factor in the sensor design and is evaluatedas:

Q =k

�ob. �12�

The calculated quality factor of the sensor is approximately440.

4 Readout CircuitA schematic of the capacitance readout circuit is presentedin Figure 11. The circuit utilizes a differential charge am-plifier, taken from Ref. 19, and a voltage comparator with avariable resistor to generate an on or off signal to the ve-hicle battery supply system. The differential charge ampli-fier in the readout circuit converts the sensor capacitance toan electrical voltage. A sinusoidal voltage input of 2 Vp-p,60 kHz is connected to the capacitive sensor, Csensor. Ahigh-frequency input signal is selected as the power sourceof the sensor to minimize the noise within the inputsignal.20 A replica of this sinusoidal input is phase-shifted180 deg using an inverting amplifier to generate anothervoltage input to the reference capacitor, Cref. The voltageoutput of the charge amplifier, Vout, is then calculated as:19

Vout =�Csensor − Cref�

CdemVin, �13�

where Cdem is the capacitance value across the charge am-

Fig. 10 Influence of the aspect ratio on the damping constant.

plifier.

Jan–Mar 2008/Vol. 7�1�5


uc1atav

aiataDvrvcp

ic of th


J

Downloaded F

According to Eq. �13�, a small capacitance should besed for Cdem in order to maximize the voltage output of theharge amplifier. In this design of the readout circuit, a-pF capacitor is selected as the Cdem. The 1-nF capacitornd the 2.5-M� resistor in front of the charge amplifier areo block away any DC voltage in the signal into the chargemplifier.19,21 A feedback resistor of 2.5 M� is used to pre-ent saturation of the charge amplifier.19,22

The voltage output from the charge amplifier is thenmplified by a gain factor of 11 through the following non-nverting amplifier. The amplified voltage is then passed to

peak detector design, taken from Ref. 23. The peak de-ector determines the peak value of the sinusoidal voltagend outputs a steady DC voltage at this peak value. ThisC voltage will be compared to another voltage from aariable resistor at the voltage comparator. The variableesistor allows adjustment on the threshold voltage of theoltage comparator. The output from the entire readout cir-uit is normally an on signal. When the voltage from the

Fig. 11 Schemat

eak detector increases and exceeds the threshold value as



a result of an external force imparted onto the sensor, an offsignal will be generated to notify the battery power systemto terminate the power supply.

A prototype of the readout circuit was fabricated usingthe conventional printed circuit board �PCB� fabricationprocess. The PCB is a double-layered board with a0.062-in. FR-4 laminate as the board material. It has anoverall dimension of 3.5 in. by 2.5 in. and has an LED toindicate the on and off signals of the board output. Thesensitivity of this circuit, using Eq. �13� with an amplifica-tion factor of 11, is calculated to be 22 V /pF.

5 Noise AnalysisSystem noise affects the overall performance of the sensor.Since MEMS devices are very small and compact, the out-put signal is usually low. The signal-to-noise ratio of thesystem has an impact on factors such as the minimum de-

e readout circuit.

tectable signal of the sensor.

Jan–Mar 2008/Vol. 7�1�6


5

Tiiwia

X

wa

q

X

TCcfnr�me0n

nant fr


J

Downloaded F

.1 Noise in the Sensor

he thermodynamics noise, or the Johnson-Nyquist noise,s one of the noise sources in the MEMS sensor. This noises due to the energy dissipation of the thermal energy storedithin the mechanical structure. The resultant noise-

nduced displacement on the sensor structure can be evalu-ted using the Brownian noise equation:24,25

n2 =

4kBTb

k2 � 1

1 − ��/�o�2 + ��/Q�o�2� , �14�

here kB is the Brownian constant, 1.38�10−23 Nm /K,nd T is the temperature.

For sensors to be operated well below the resonant fre-uency, the Brownian noise equation can be simplified as:

n2 =

4kBTb

k2 =4kBT

woQk. �15�

he resonant frequency of the sensor was simulated usingOMSOL. Considering the sensor structure as a lumpedantilever beam, the first-mode resonant frequency wasound to be 61 kHz, as shown in Fig. 12. The equivalentoise level at this frequency for a Q factor of 440 and atoom temperature 298 K was determined to be 4.24

10−16 m /Hz1/2. Since the displacement of the comb-driveechanisms are 1 /10 of the sensor displacement, the

quivalent noise level of the comb drive is approximately.43�10−16 m /Hz1/2, which is equivalent to a voltage

1/2

Fig. 12 First-mode reso

oise level of 0.088 nV /Hz .



5.2 Noise in the Readout Circuit

At the low-frequency region, flicker noise or 1 / f noise isthe dominating noise source. Flicker noise is present in allmaterials used in electronics and can be evaluated by thefollowing equation:26

NoiseFlicker = �K2 lnfH

fL�1/2

, �16�

where K is the normalized device constant at 1 Hz, and fHand fL are the lowest and highest frequency of the band-width.

If the input voltage noise spectral density of the compo-nent is known, the K value can be evaluated using the fol-lowing equation:27

equency of the sensor.

Fig. 13 Simulation result of the voltage noise spectral density.

Jan–Mar 2008/Vol. 7�1�7


K

wio

e

iorm

e

w

maflnl

F�


J

Downloaded F

2 = �eL2 − eF

2��fL� , �17�

here eL is the noise voltage at lowest possible frequencyn the 1 / f noise region at frequency fL, and eF is the whiter flat noise voltage.

Using Eq. �17�, the K values of the operational amplifi-rs, AD-743, are calculated to be 24.8 nV.

Similar to the sensor structure, thermal noise also existsn all resistive elements within the readout circuit. The levelf noise is proportional to the temperature as well as theesistance value. At frequencies below 100 MHz, the ther-al noise can be estimated with the following equation:26

Thermal = �4kBTR�1/2, �18�

here R is the resistance value in ohms.For the readout circuit proposed in this work, the ther-

al noise from the two 2.5-M� resistors in the chargemplifier is determined to be the most significant. Theicker noise in the circuit is comparatively small and iseglected from the calculation. Each resistor has a noise

1/2

ig. 14 The 6-DOF manipulator in the experiment: �a� overall view;b� close-up view of the sensor on the platform stand.

evel of 202 nV /Hz . Adding the two noncorrelated noise



sources together, a total noise level of 290 nV /Hz1/2 is es-timated. This noise signal will be amplified as it passesthrough the noninverting amplifier, which has a gain factorof 11. As a result, output noise is expected to be about3190 nV /Hz1/2. Simulation results from TINA-TI softwarein Fig. 13 show that the noise floor of the sensing circuit isabout 3250 nV /Hz1/2 or 3.25 �V /Hz1/2, which is veryclose to the noise level computed from the previously men-tioned equations and correlations.

6 Experimental Results

6.1 Sensor CapacitanceFigure 14 shows the manipulator and the tungsten probewith a tip radius of 25 �m used in the experiment.28 Thesensor was wire-bonded onto a standard ceramic pin guidearray �PGA�, as shown in Fig. 15. The packaged sensor wasmounted onto the movable platform of the manipulator dur-ing the experiment. The movable platform has a resolutionof 0.2 �m in the x-, y-, and z-axis directions.28 The sensorwas pushed toward the probe tip to emulate a force inputonto the sensor. Corresponding to a 1-mm displacement ofthe sensor toward the probe is an input force of 537 �Napplied to the sensor. A capacitance-to-digital converterboard �Model AD7746, Analog Devices� was connected tothe sensor to measure the change in the capacitance at dif-ferent loading conditions. For every measurement, 100 datapoints were sampled, and the peak-to-peak variation wasmeasured to be less than 5 fF. The capacitance values weretransferred to a desktop computer for analysis. Experimen-tal results and the theoretical prediction are plotted in Fig.

Fig. 15 The sensor packaged on a PGA.

16. For a 20-�m displacement, a 188-fF change in the ca-

Jan–Mar 2008/Vol. 7�1�8


pcdditttbc

6TtcsvTcfpr1�tp

Fc


J

Downloaded F

acitance value was measured, whereas a change of 500-fFapacitance is expected from the theoretical equation. Theiscrepancy between two values may be due to out-of-planeisplacement in the comb drive, which results a reductionn the overlapping area. One possible solution to minimizehe effect is to use flexure hinges to ensure that the deflec-ions are in the planar direction.29 Inherent parasitic capaci-ance within the sensor and the testing equipment may alsoe another error source for the discrepancy in the outputapacitance of the sensor.

.2 Readout Circuit Outputo evaluate the performance of the readout circuit, capaci-

ors with known values were connected to the readout cir-uit. The voltage output after the charge amplifier was mea-ured using an HP54600 oscilloscope. Fig. 17 shows theoltage output for a 4-pF capacitor after the readout circuit.he measured capacitance-voltage response of the readoutircuit for a series of capacitors was plotted in Fig. 18. It isound that a relatively linear relationship between the ca-acitance and the voltage output can be achieved with theeadout circuit, providing a sensitivity of approximately.45 V /pF. The theoretical sensitivity calculated using Eq.13� is approximately 2 V /pF. The discrepancy betweenwo values is mainly due to the existence of parasitic ca-acitance in the operational amplifier and in the circuit

Fig. 16 Capacitance response of the MEMS force sensor.

ig. 17 Oscilloscope output of the readout circuit with a 4-pF
apacitor.


board. Since a small value of Cdem is selected, the parasiticcapacitance will increase Cdem and hence lower the sensi-tivity of the circuit.

The noise level of the sensing circuit was measured us-ing an HP/Agilent 35670A Dynamic Signal Analyzer. Thevoltage noise spectral density of the circuit was plotted onFig. 19. The noise floor was measured to be5.21 �V /Hz1/2, compared to the calculated noise of3.25 �V /Hz1/2. For an operating bandwidth of 1 kHz, thenoise voltage will be about 164 �V. Hence, the equivalentminimum detectable capacitance of the readout circuit isapproximately 0.07 fF.

7 ConclusionA power disconnect proposed for use with 42-V automotivepower systems is presented in this paper. The design aimsto prevent arc discharge during the disconnection of suchhigh-voltage systems. This power disconnect, integratedwith a MEMS-based force sensor, will terminate the powersupply when a service technician disconnects the connec-tors. A space-variation type capacitive sensor is selected as

Fig. 18 Voltage output of different capacitors.

Fig. 19 Experimental result of the voltage noise spectral density.

Jan–Mar 2008/Vol. 7�1�9


trsmb2srctt1s

ATCt

R

1

1

1

1

1

1

1

1

1

1


J

Downloaded F

he force sensing mechanism. This sensor design incorpo-ates a displacement reduction mechanism to increase theensitivity of the sensor. Results show that a 10:1 displace-ent reduction ratio can be achieved and the sensitivity can

e improved by 57 times with the mechanism. When a0-�m displacement, or a 10.7-mN force, is applied on theensor, the capacitance output is measured to be 188 pF. Aeadout circuit is designed to convert the capacitancehange to voltage and to generate the off signal to terminatehe power supply when the voltage exceeds the presethreshold. This circuit board has a sensitivity of5.95 V /pF after amplification, and the noise floor is mea-ured to be 5.21 �V /Hz1/2.

cknowledgmentshis research work was funded by AUTO21, a Network ofentres of Excellence of Canada program. The authors also

hank CMC Microsystems for MEMS chip fabrication.

eferences

1. A. Graf, “Semiconductors in the 42 V PowerNet,” presented at the 42V Automotive Systems Conference, Chicago, 17–18 September 2001,Intertech Corp., Portland, ME �2001�.

2. A. Jahed, A. Housely, and J. Spoonhower, SEMA 42-volt systemsoverview, Specialty Equipment Market Association �SEMA� �2002�.

3. R. Dixon and J. Bouchaud, “Prospects for MST sensors in automo-tive,” in Advanced Microsystems for Automotive Applications, J. Vall-dorf and W. Gessner, Eds., pp. 3–12, Springer, Berlin, Heidelberg�2006�.

4. H. K. Chu, J. K. Mills, and W. L. Cleghorn, “Stereolithography as ameso-structure for input force reduction to a capacitive force MEMSsensor,” Proc. SPIE 6464, 64640Q �2007�.

5. L. K. Baxter, Capacitive Sensors: Design and Applications, IEEEPress Series on Electronics Technology, R. J. Herrick, Ed., JohnWiley & Sons, New York �1997�.

6. M. Despont, G. A. Racine, P. Renaud, and N. F. de Rooij, “Newdesign of micromachined capacitive force sensor,” J. Micromech. Mi-croeng. 3, 239–242 �1993�.

7. Y. Sun, B. J. Nelson, D. P. Potasek, and E. Enikov, “A bulk micro-fabricated multi-axis capacitive cellular force sensor using transversecomb drives,” J. Micromech. Microeng. 12, 832–840 �2002�.

8. J. Aebersold, K. Walsh, M. Crain, M. Martin, M. Voor, J.-T. Lin, D.Jackson, W. Hnat, and J. Naber, “Design and development of aMEMS capacitive bending strain sensor,” J. Micromech. Microeng.16, 935–942 �2006�.

9. H. K. Chu, J. K. Mills, and W. L. Cleghorn, “Design of a highsensitivity capacitive force sensor,” in 7th IEEE International Con-ference on Nanotechnology �2007� �in press�.

0. N. Lobontiu and E. Garcia, “Analytical model of displacement am-plification and stiffness optimization for a class of flexure-based com-pliant mechanisms,” Comput. Struct. 81, 2797–2810 �2003�.

1. E. Garcia, Y. Nam, and N. Lobontiu, “Design, modeling, and initialexperiments on microscale amplification device,” J. Intell. Mater.Syst. Struct. 16, 1039–1049 �2005�.

2. CMC Microsystems and Micralyne, “Introduction to micragem: asilicon-on-insulator based micromachining process,” Report ICI-138V4.0, CMC Microsystems and Micralyne �2005�.

3. A. S. Tamsir F. Saharil, and B. Y. Majlis, “Finger beam geometricalbehavior analysis of a non-crossing differential capacitive MEMSaccelerometer,” Proc. Intl. Conf. Software Engg. 673–676 �2004�.

4. J. Shigley, Mechanical Engineering Design, 7th ed., pp. 72–104,McGraw Hill, New York �2003�.

5. J. Brotz, “Damping in CMOS-MEMS Resonator,” MASc Disserta-tion, Dept. Electrical and Computer Engineering, Carnegie MellonUniversity, Pittsburgh �2004�.

6. Y. H. Cho, B. M. Kwak, A. P. Pisano, and R. T. Howe, “Slide filmdamping in laterally driven microstructures,” Sens. Actuators, A 40,31–39 �1994�.

7. R. P. van Kampen, M. J. Vellekoop, P. M. Sarro, and R. F. Wolffen-buttel, “Application of electrostatic feedback to critical damping ofan integrated silicon capacitive accelerometer,” Sens. Actuators, A43, 100–106 �1994�.

8. S. Beeby, MEMS Mechanical Sensors, pp. 85–112, Artech House,Boston, �2004�.

9. M. Kraft, “Closed loop digital accelerometer employing oversam-pling conversion,” PhD Dissertation, School of Engineering, Coven-try University, UK �1997�.

. Micro/Nanolith. MEMS MOEMS 013010-1


20. N. Yazdi, H. Kulah, and K. Najafi, “Precision readout circuit forcapacitive microaccelerometers,” in Proc. IEEE Sensors 2004, D.Rocha, P. M. Sarro, and M. J. Vellekoop, Eds., pp. 28–31 �2004�.

21. B. E. Boser, “Electronics for micromachined inertial sensors,” in Int.Conf. Solid-State Sensors and Actuators, Vol. 2, pp. 1169–1172�1997�.

22. W. J. Smith and J. R. LaCourse, “Non-contact biopotential measure-ment from the human body using a low-impedance charge amplifier,”Proc. IEEE 30th Annual Northeast Bioengineering Conference, Vol.30, pp. 31–32 �2004�.

23. Analog Devices, “AD843 Datasheet,” Rev. D, Analog Devices Inc.,Norwood, MA �2006�.

24. G. M. Rebeiz, RF MEMS Theory, Design, and Technology, pp. 429–448, John Wiley and Son, Hoboken, NJ �2003�.

25. M. Suster, J. Guo, N. Chaimanonart, W. H. Ko, and D. J. Young, “Ahigh-performance MEMS capacitive strain sensing System,” J. Mi-croelectromech. Syst. 15�5�, 1069–1077 �2006�.

26. B. Carter, “Op-amp noise theory and applications,” in Op Amps forEveryone, Texas Instruments Literature No. SLOA082, Texas Instru-ments, Inc., Dallas �2001�.

27. “Noise analysis in operational amplifier circuits–application report,”Texas Instruments Literature No. SLVA043B �2007�.

28. N. Dechev, L. Ren, W. Liu, W. L. Cleghorn, and J. K. Mills, “Devel-opment of a 6 degree of freedom robotic micromanipulator for use in3D MEMS Microassembly,” in Proc. 2006 IEEE Int. Conf. Roboticsand Automation, pp. 281–288 �2006�.

29. N. Lobontiu, J. S. N. Paine, E. Garcia, and M. Goldfarb, “Corner-filleted flexure hinge,” J. Mech. Des. 123�3�, 346–352 �2001�.

Henry K. Chu received a BASc degree inmechanical engineering from the Universityof Waterloo, Ontario, Canada, in 2005. Heis currently working toward an MASc degreein mechanical and industrial engineering atthe University of Toronto. His research inter-ests include MEMS force sensor and pack-aging design for automotive applications.

James K. Mills received bachelor’s andmaster’s degrees in electrical engineering in1980 and 1982, respectively, and a PhD de-gree in mechanical engineering in 1987from the University of Toronto, Ontario,Canada. He is currently a professor of me-chanical engineering at the University ofToronto, and his research interests have en-compassed a number of related areas, in-cluding robot control, control of multirobots,control of flexible link robots, design of ac-

tuators, localization, development of fixtureless assembly technol-ogy, design and control of high-speed machines, and developmentof neural network controllers. He has published over 220 journal andconference papers and supervised over 55 master’s and PhD stu-dents and a number of postdoctoral fellows and research engineers.He has been an invited visiting professor at the Centre for ArtificialIntelligence and Robotics in Bangalore, India, in 1997 and the HongKong University of Science and Technology in 1995. More recently,he was a visiting professor in the Department of Automation andComputer Aided Engineering at the Chinese University of HongKong from 2002 to 2003.

William L. Cleghorn received a BASc de-gree in 1975, an MASc degree in 1976, anda PhD degree in 1980, all from the Univer-sity of Toronto, Ontario, Canada, Depart-ment of Mechanical Engineering. He is cur-rently a professor in the same department,and his current areas of research includedynamics, vibrations, and mechanisms. Hehas published over 200 journal and confer-ence papers in these areas. He currentlyholds the Clarice Chalmers Chair ofEngineering Design.

Jan–Mar 2008/Vol. 7�1�0


mems-based power disconnect for 42-v automotive power systems

Documents