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Mechanics of Materials 38 (2006) 11–24

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Numerical analysis of ultrasonic wire bonding: Effectsof bonding parameters on contact pressure

and frictional energy

Yong Ding, Jang-Kyo Kim *, Pin Tong

Department of Mechanical Engineering, Hong Kong University of Science and Technology, Clear Water Bay,

Kowloon, Hong Kong

Received in revised form 27 April 2004

Abstract

The elasto-plastic large deformation taking place in ultrasonic wire bonding is analysed by means of 2-D and 3-Dfinite element method. A special focus has been placed on how the important wire bonding parameters, such as bondforce and power, affect the contact pressure along the wire-bond pad interface. It is shown that the contact interface hada long elliptical shape, and the maximum contact pressure occurred always at the periphery of the contact interface,which is consistent in the current 2-D and 3-D finite element analyses. The normalised real contact area as well asthe maximum frictional energy intensity varied in a similar manner to the contact pressure, with the maximum valuesoccurring at the periphery of contact interface, where weld is preferentially formed in practical wire bonding. A higherbond force does not result in a higher contact pressure, or higher frictional energy intensity, suggesting that a high bondforce is not directly correlated to better wire bondability.� 2005 Elsevier Ltd. All rights reserved.

Keywords: Ultrasonic wire bonding; Wire bondability; Finite element method; Contact pressure; Frictional energy intensity

1. Introduction

Wire bonding is by far the most popular first-level interconnection technology used between

0167-6636/$ - see front matter � 2005 Elsevier Ltd. All rights reservdoi:10.1016/j.mechmat.2005.05.007

* Corresponding author. Tel.: +852 23587207; fax: +85223581543l.

E-mail address: [email protected] (J.-K. Kim).

the die and package terminals. Amongst variouswire bonding processes, the ultrasonic wire bond-ing (UWB) has the advantages of fast bondingprocess, high productivity, excellent electricalperformance, good heat conductivity and goodcorrosion resistance (Tummala, 2001). It has beenwidely used in various electronic packages, such aschip on board (COB), chip scale package (CSP)

ed.

mailto:[email protected]

12 Y. Ding et al. / Mechanics of Materials 38 (2006) 11–24

and ball grid array (BGA) packages. The UWB iscarried out by pressing the metal wire on a bondpad of semiconductor and vibrating it using highfrequency ultrasound, generating frictional energy.The energy applied onto the bond pad breaksdown the surface oxide film, allowing the forma-tion of intermetallics between the wire and thebond pad (Tummala, 2001; Harman, 1997). Ad-vances in wire bonding technology require higherreliability, bonding speed and product yield (Har-man, 1997; Winchell and Berg, 1978; Hamidi et al.,1999). However, the details of bonding mecha-nisms taking place in UWB remain ill-understoodand most our understanding is qualitative and/orexperimental.

The wire bond strength is one of the main crite-ria used to ensure the reliability of UWB. Thebond strength is dependent on many process andmaterial variables, such as ultrasonic power, ap-plied force, welding time, bond pad surface hard-ness and roughness, interface temperature (Kimand Au, 2001; Chan et al., in press). The bondstrength can be improved by increasing the ultra-sonic power, but too high a power can lead to pre-mature failure at the wire neck (Sheaffer andLevine, 1991). Ultrasonic power is a dominant fac-tor for bond formation (Sheaffer and Levine, 1991;Hu et al., 1991) and a proper range of bond forceis required for efficient transmission of ultrasonicpower. The influences of hardness and surfaceroughness of bond pad have been extensively stud-ied (Jeng and Hong, 2001; Chan et al., 2004; Jenget al., 2001). Although a high surface roughnesscan reduce the bonding time for successful bondsat a given bond force and power (Jeng and Hong,2001), too high a surface roughness did not furtherimprove the wire bondability (Chan et al., 2004;Jeng et al., 2001). The bond pad temperature alsoplayed an important role in determining the wirepull strength of metallization on an organic sub-strate (Hu et al., 1991; Jeng et al., 2001). An opti-mal preheat temperature produced the maximumbond strength, and too high a temperature wasdetrimental to wire bonding (Chan et al., 2004).

For such a complex problem with many param-eters affecting the wire bond quality, to identifykey mechanisms is essential to understanding thebonding mechanisms. Fortunately, the lift-off fail-

ure of bond area provides a valuable insight: it isshown (Harman, 1997; Winchell and Berg, 1978)that a strong bond was formed preferentially atthe periphery of the contact interface, whereasthe central region remained largely unbonded.This phenomenon is not unique for the ultrasonicbonding and is also observed in the thermocom-pression wire bonding (Harman and Albers,1977), suggesting a non-uniform application ofenergies across the contact interface. An improvedunderstanding is needed as to what mechanicalconditions in terms of stress, contact pressureand frictional energy at the wire-bond pad inter-face produce such a non-uniform bond.

The analytical solutions for the displacementand stress distribution in the wire and bond padduring wire bonding are very complicated due tothe complexity in bonding geometry and materialproperties. Recognizing the significance of stressdistribution along the wire-bond pad interface,increasing efforts are directed towards employingnumerical methods. The use of the finite elementmethod (FEM) in particular allows a more accu-rate description of wire deformation. In addition,the specific loading geometries as well as varyingmechanical properties of bond pads can be prop-erly taken into account. The stress distribution atthe wire-bond pad interface has been studied (Ike-da et al., 1999) for the ball bonding based on theaxisymmetric 2-D FEM (Ikeda et al., 1999; Taka-hashi and Inoue, 2002). The maximum equivalentstresses were determined for the periphery of theinterface assuming a rigid bond pad in the contactanalysis (Ikeda et al., 1999). Few studies have hith-erto been reported based on the 3-D FEM analysisthat can reflect accurately the global deformationand stress distribution of wire bonding.

Although the above FEM analyses were able tocalculate the stress distributions at the wire-bondpad interface, the changes in bond pad surfacecharacteristics and ultrasonic energy cannot beincorporated in the analysis. Experiments (Har-man and Albers, 1977; Joshi, 1971; Harman andLeedy, 1972) suggest that the UWB is a frictionalbonding process. A micro-contact approach wasadopted previously (Jeng and Hong, 2001; Jenget al., 2001) to take into account the surface rough-ness of bond pad. Real contact area and frictional

Y. Ding et al. / Mechanics of Materials 38 (2006) 11–24 13

energy were calculated to interpret the experimen-tal results of wire bondability. However, the non-uniform stress distribution which in turn formsthe bond along the periphery, have not been spe-cifically studied. The non-uniform stress distribu-tions at the interface may cause non-uniformcontact pressure, uneven real contact area and fric-tional energy intensity, all of which affect thebondability.

With the above findings in mind, the objectivesof this paper are (i) to develop a 3-D and a precise2-D models to simulate the ultrasonic wire bond-ing, and (ii) to investigate the bonding mechanismsin terms of contact pressure, real contact area andfrictional energy intensity generating at the wire-bond pad interface. Elasto-plastic large deforma-tion contact analyses are performed based on theFEM.

2. Finite element analysis (FEA)

Both the 3-D and 2-D models of ultrasonicwedge bonding are considered in the present FEanalysis. The geometry, the loading method andthe boundary conditions were selected to representthose of the actual experimental technique, asshown in Fig. 1. A mesh was created for the sym-metric loading geometry using the FE code ABA-QUS, as shown in Fig. 2. Due to the symmetry,only one quarter of the structure was adoptedfor the 3-D model which was composed of

R l

Wire

Pad

Substrate

bond wLoad

Fig. 1. Schematic of we

20,352 brick elements and 24,551 nodes. This en-sures sufficient resolution and thus accuracy ofthe results while maintaining a reasonable timeneeded for computation. The model consisted ofa wedge tool, wire, bond pad and substrate. Thewedge tool was assumed to be a rigid body becauseit is usually made of titanium carbide, which ismuch harder than the Au wire. The Au wire hada cylindrical shape of radius 25.4 lm. The bondpad had a layered structure consisting of Au, Niand Cu layers and was electrolytically plated onthe FR-4 substrate that was made from glass wo-ven fabric reinforced epoxy laminates. The dimen-sions and material properties of these componentsare presented in Fig. 2 and Table 1 (Callister, 2003;ASM International Handbook Committee, 1990).

The wire bonding process is carried out at 25 �Cin this study. When the Au wire was being pressedby the wedge tool toward the bond pad, there ex-ists two contact interfaces, one between the wireand bond pad, the other between the wire andwedge. A typical bond force–time history of ultra-sonic wire bonding is shown in Fig. 3 (Harman,1997; Ikeda et al., 1999), where the applied forceincreases linearly in the first few ms (Step 1), fol-lowed by a constant force for the rest of wirebonding (Step 2). The first step takes place onlyfor a very short period of time, say about 1 ms(Ikeda et al., 1999), whereas the second step takesup the most time of wire bonding. During the firststep, the wire undergoes large elasto-plastic defor-mation under the applied force and the interface is

l = 51um

R = 15umR

edge

dge wire bonding.

Fig. 2. 3-D FEM model for wire bonding.

Table 1Material properties used in the present analyses (Ikeda et al., 1999; Callister, 2003; ASM International Handbook Committee, 1990)

Material Young�s modulus (GPa) Poisson�s ratio Rate of strain hardening (MPa) Initial yield stress (MPa)

Au 68.6 0.44 Eq. (1) 32.7Ni 207 0.31 – –Cu 115 0.308 – –

Bond Force

(Newton)

Bond Time

(msec)

Fb

0 20

Ultrasound application

Step 1 Step 2

Fig. 3. Bond force-time history during ultrasonic wire bonding.


formed between the wire and bond pad. Duringthe second step, a bond is formed between the wireand bond pad due to the applied ultrasonic vibra-tion which in turn generates frictional energy. The

large elasto-plastic deformation contact analysiswas performed for the second step using theFEM to simulate wire deformation and stress dis-tribution at the contact interface. The real contactarea was also calculated based on the micro-con-tact approach. The frictional energy consumed atthe interface during the second step was adoptedas the criterion for bondability.

One of the most important material propertiesrequired for proper simulations of wire bondingis the plastic property of Au wire. Judging fromthe maximum bond force of 0.2 N and the loadingtime of 1 ms for Step 1, the loading speed approx-imates to 0.2 N/ms. The extremely fast wire bond-ing process may result in the deformation of Auwire at a very high strain rate, with a correspond-ing increase in yield strength of Au wire due to thestrain rate effect. According the Hopkinson pres-

Table 2Material properties of FR-4 substrate (Yao and Qu, 1999)

E1 (GPa) E2 (GPa) E3 (GPa) G12 (GPa) G13 (GPa) G23 (GPa) m12 m32 m13

22.4 1.6 22.4 0.2 0.63 0.2 0.14 0.14 0.002


sure bar test (Ikeda et al., 1999), the yield stress,ry, of Au wire is expressed as:

ry ¼32:7þ 0:057_e ðMPaÞ for Step 1;

32:7þ 1459e ðMPaÞ for Step 2.

�ð1Þ

The first equation takes into account the strainrate effect upon linear increase in bond force (Step1), whereas the second equation considers the plas-tic strain hardening effect during constant loading(Step 2). Ni and Cu were assumed linearly elasticin this study because they are much harder thanAu. The FR-4 substrate was also assumed linearlyelastic because it has a large supporting area and isfar away from the contact interface. Its elasticproperties in the in-plane and out-of-plane direc-tions are shown in Table 2 (Yao and Qu, 1999).

To compare the accuracy of stress distributionobtained from the 3-D model as well as to accu-rately calculate the real contact area, a 2-D analy-sis was also performed using a refined meshespecially for the wire and bond pad. Due to thesymmetry, only one half the structure was mod-elled with a total of 23,559 quadrilateral elementsand 23,965 nodes, as shown in Fig. 4. The layered

Fig. 4. 2-D FEM model of wire bonding.

structure of the bond pad and their material prop-erties were the same as those used in the 3-D mod-el. A plane strain condition was assumed for thecross-section of the structure.

3. Results and discussion

3.1. 3-D model

In the 3-D analysis, the bond force was in-creased linearly until it reached a plateau value,Fb, which remained constant until the end of wirebonding (Fig. 3). Fig. 5 shows the deformationand von Mises stress distributions for bond forcesFb = 0.1 and 0.2 N. At the low bond force of0.1 N, the wire has just started to deform with highstress concentrations taking place at the peripheryof the contact interface between the wedge andwire, as well as between the wire and bond pad.The minimum stress occurred in the central regionof contact interface, and there were hardly stressesbuilt up in the substrate underneath the bond pad.When the bond force was increased to 0.2 N, thestress concentrations occurred at the same contactregions although the magnitudes of stress distribu-tion generally increased.

The stress distributions are further studied inmore details with reference to Fig. 6, which plotsthe contact pressure distribution over the wholecontact interface between the wire and bond padin different view directions. Fig. 6(a) presents a3-D perspective view, whereas Fig. 6(b)–(d) arethe 2-D views seen in the z, x and y axes, respec-tively. The wire-bond pad contact interface had along elliptical shape, with the contact width inthe lateral direction being larger in the central areathan at the wire ends (Fig. 6(b)). It is clearly seenthat the maximum contact pressure occurred atthe periphery of the contact interface, which isconsistent with the previous analysis as far as thebond pad is sufficiently thick (Takahashi and

Fig. 5. Deformation and Mises stress (in · 106 MPa) distributions at bond forces of (a) 0.1 N and (b) 0.2 N.


Inoue, 2002). There is striking analogy between theregion corresponding to maximum contact pres-sure and the real bonded area identified by themicroscopic analysis (Harman, 1997; Winchelland Berg, 1978). Fig. 7 shows the SEM photo-graphs taken of a typical wedge bond fracturesurface. Much of the central area was unbonded,which is consistent with the generally low contactpressure seen in Fig. 6(a) and (b). It is also notedthat the contact pressure in the central region ofcontact interface (near y = 0) was even higher than

that at the wire edges (y = 40–50 m), see Fig. 6(c).The large contact pressure concentrated in themid-region of �25 lm < y < 25 lm along the wiredirection, which is a reflection of the initial contactbetween the wire and wedge over the wedge lengthof l = 51 lm (see Fig. 1).

3.2. 2-D model

The effects of bond force on bonded interfacearea and stress distributions were further studied

Fig. 6. Contact pressure distributions at a bond force of 0.2 N: (a) 3-D view; (b) top view; (c) view from the lateral direction and (d)view from the wire longitudinal direction.


based on the 2-D model. Fig. 8 shows the wiredeformation and the contact pressure distributionsalong the contact length between the wire andbond pad at varying bond force, Fb. It is obviousthat the maximum contact pressure occurred al-ways at the periphery of the contact interface, con-sistent with the results from the 3-D analysis. Theabsolute magnitudes of the maximum contactpressure were also almost identical between the2-D and 3-D models, justifying the applicabilityof 2-D analysis in liu of the more time-consuming3-D analysis.

Fig. 9 shows the changes in contact length, bn,when the bond force, Fb, varied from 0 to 0.2 N,indicating that the nominal contact length is line-

arly proportional to bond force. Fig. 10 showsthe maximum and average contact pressures atthe contact interface taken from Fig. 8. It is foundthat the maximum contact pressure seen at theperiphery increased initially with increasing thebond force, followed by a constant plateau valuefor bond forces above 0.07 N, whereas the meancontact pressure remained almost a constant overthe whole bond force increment. There were slightfluctuations in the maximum and average contactpressure values due to marginal changes in contactarea for these bond forces. Judging from the factthat the real bonded area is confined to the periph-ery of the contact interface (Harman, 1997; Winc-hell and Berg, 1978) and the majority of central

Fig. 7. SEM photographs of wedge wire bonds at different bond time: after (a) 0 ms, (b) 4 ms, (c) 7 ms and (d) 10 ms. After Harman(1997).


contact region is not bonded, the maximum con-tact pressure at the perimeter is more importantthan the average value. According to the observa-tion that there were negligible changes in maxi-mum contact pressure at bond forces above0.07 N, a moderately high bond force was alreadysufficient to make wire bonds, and too high a bondforce is not necessarily beneficial.

To verify the accuracy of 2-D model, the pres-sure distribution obtained from the 2-D FEManalysis was compared with that obtained fromthe 3-D FEM model. Two cross-sections at y = 0and y = 17 lm were chosen in the 3-D model forthis purpose, representing the axis of symmetryand the region with moderately high contact pres-sure, both within the main contact interface,respectively. The contact pressures shown inFig. 11 indicates that the results obtained fromthe 2-D and 3-D FEM analyses were almost iden-tical in terms of both pressure distribution and

maximum value. This confirms the validity of the2-D analysis used in this study, which can providethe advantage of a drastically reduced computa-tion time.

3.3. Real contact area

Although it was assumed in the present analysisthat the bond pad surface is perfectly flat, it, inreality, is rough on a microscopic scale. The sur-face roughness causes the real contact area duringwire bonding to become much less than the nomi-nal contact area. The micro-contact phenomenoncannot be simulated properly in the FEM modelbecause of the microscale and random distributionof the roughness. Thus, a micro-contact modelneeds to be introduced here. Since the early devel-opment of a basic elastic model (Greenwood andWillamson, 1966), there have been many studieson contact of rough surfaces (Abbott and Fire-

Fig. 8. Deformations in the wire and the corresponding contact pressure between the wire and bond pad at different bond forces.

0.00 0.05 0.10 0.15 0.200

2

4

6

8

10

12

14

b n (u

m)

Bond Force (N)

Fig. 9. Variation of contact length, bn, as a function of bondforce.


stone, 1933; Pullen and Williamson, 1972; Changet al., 1987; Zhao et al., 2000; Tabor, 1959; John-son, 1968; Lim and Ashby, 1987; Ashby et al.,1991). The rough-surface contact was character-ized by full plastic deformation of asperities (Ab-bott and Firestone, 1933; Pullen and Williamson,1972), whereas some (Chang et al., 1987; Zhaoet al., 2000) developed elasto-plastic contact mod-els. The elasto-plastic models were adopted later inthe study of wire bonding (Jeng and Hong, 2001;Jeng et al., 2001). Although the above contactmodels considered surface profiles on a micro-scopic scale, the effect of shear stress was not

0.00 0.05 0.10 0.15 0.200

50

100

150

200

250

Max Contact Pressure Average Contact Pressure

Con

tact

Pre

ssur

e (M

Pa)

Bond Force (N)

Fig. 10. Variations of maximum and average contact pressuresas a function of bond force.

0 2 4 6 8 10 12 140

50

100

150

200

250 2-D model 3-D model at the cross-section of y=0 3-D model at the cross-section of y=17um

Con

tact

Pre

ssur

e (M

Pa)

x (um)

Fig. 11. Comparisons of contact pressure between 2-D and 3-DFEM analyses.


specifically taken into account. In ultrasonic orthermosonic wire bonding, the contact surfacesslide each other, generating significant shear stres-ses at the asperities of the contact surfaces. The sei-zure load, Fs), developed at the surface asperitiesat the real contact area is expressed as (Tabor,1959):

F s ¼H 0An

ð1þ atl2Þ1=2. ð2Þ

An is the nominal contact area, which was assumedto be identical to the real contact area, Ar in theiranalysis (Tabor, 1959); H0 is the hardness of thesofter material of the two contact surfaces; and lis the coefficient of friction. at = 12 is a constant

determined from experiment (Tabor, 1959; John-son, 1968). Eq. (2) was further modified (Limand Ashby, 1987; Ashby et al., 1991) to explainthe wear mechanisms assuming that the real con-tact area for plastic contact is given by

Ar

An

¼ FF s

; ð3Þ

where F is the normal contact force. When F > Fs,Ar becomes identical to An. In 2-D model, An andAr are represented by the lengths of bn and br,respectively. Considering the non-uniform pres-sure distribution between the wire and bond pad,Eq. (2) can be reduced:

ps ¼H 0

ð1þ atl2Þ1=2; ð4Þ

where ps is the seizure pressure. Here, m(x,y) isdefined as the ratio of the real contact area tothe nominal contact area:

mðx; yÞ ¼ DAr

DAn

¼ pðx; yÞps

; ð5Þ

where DAr and DAn are the micro real and micronominal contact areas at the point (x,y); andp(x,y) is the corresponding contact pressure.When p(x,y) > ps, m(x,y) becomes unity. For a2-D analysis, m(x,y) can be simplified to m(x).When H0 = 245 MPa (ASM International Hand-book Committee, 1990) for the Au wire, l = 0.38(Mayer et al., 2002), and the corresponding seizurepressure, ps = 170 MPa.

The normalised real contact area, m(x,y) pre-sented in Fig. 12 varied along the wire lateraldirection in a similar manner to the contact pres-sure shown in Fig. 8, except the very edge wherethe normalised contact area consistently showeda plateau constant for a length of 0.5–0.8 lm forall bond forces studied. The plateau constant valuerepresents the most intimate real contact whereweld is most likely to be made. The maximum va-lue of the normalised real contact area was almostidentical for all bond forces studied, which is con-sistent with the summary drawn from Fig. 10 inthat a high bond force does not necessarily meana large area of intimate contact. Fig. 13 summa-rises the real contact length, br, which increasedlinearly with increasing the bond force.

(c)

0 2 4 6 8 10 12 140.0

0.2

0.4

0.6

0.8

1.0

Bond Force = 0.05 N

m(x

)

x (um)(a)

0 2 4 6 8 10 12 140.0

0.2

0.4

0.6

0.8

1.0

Bond Force = 0.10 N

m(x

)

x (um)(b)

0 2 4 6 8 10 12 140.0

0.2

0.4

0.6

0.8

1.0

Bond Force = 0.15 N

m(x

)

x (um)0 2 4 6 8 10 12 14

0.0

0.2

0.4

0.6

0.8

1.0

Bond Force = 0.20 N

m(x

)

x (um)(d)

Fig. 12. Normalised real contact area, m(x), along the wire lateral direction for different bond forces.

0.00 0.05 0.10 0.15 0.200

2

4

6

8

10

12

14

b r (um

)

Bond Force (N)

Fig. 13. Variation of real contact length, br, as a function ofbond force.


3.4. Frictional energy intensity

A key factor essential to successful bonds inultrasonic wire bonding is the sufficient frictionalenergy between the wire and bond pad to form asolid state weld through local melting. It was con-firmed that bond is made preferentially along the

periphery of contact interface and the vast major-ity of central region remains unwelded, see Fig. 7(Harman, 1997; Winchell and Berg, 1978). Thisobservation indicates that the local frictional en-ergy intensity plays an important role in wirebondabilty. Considering the non-uniform pressuredistribution at the contact interface, the frictionalenergy intensity, Eif, in 3-D model is defined:

Eif ¼ ulpðx; yÞ; ð6Þwhere u is the average speed due to the reciprocat-ing motion in ultrasonic wire bonding. The aver-age speed can be defined as u = 4bf, where b andf refer the amplitude and frequency of vibration,respectively. p(x,y) is the contact pressure thatcan be obtained from the FEM analysis. For a2-D model, p(x,y) can be simplified to p(x).

The friction energy intensities were calculatedfor an ultrasonic frequency of 60 kHz and a vibra-tion amplitude of 1 lm (Harman, 1997), which areplotted for different bond forces as shown inFig. 14. The maximum frictional energy intensity

(a)

0 2 4 6 8 10 12 140

1 x105

2 x105

3 x105

4 x105

Bond Force = 0.05 N

Fric

tion

ener

gy in

tens

ity (J

/m2 )

x (um)(b)

0 2 4 6 8 10 12 140.0

1.0 x105

2.0 x105

3.0 x105

4.0 x105

Bond Force = 0.10 N

Fric

tion

ener

gy in

tens

ity (J

/m2 )

x (um)

(c)

0 2 4 6 8 10 12 140

1 x105

2 x105

3 x105

4 x105

Bond Force = 0.15 N

Fric

tion

ener

gy in

tens

ity (J

/m2 )

x (um)

(d)

0 2 4 6 8 10 12 140

1 x105

2 x105

3 x105

4 x105

Bond Force = 0.20 N

Fric

tion

ener

gy in

tens

ity (J

/m2 )

x (um)

Fig. 14. Variations of frictional energy intensity along the wire lateral direction for different bond forces.

0.00 0.05 0.10 0.15 0.200.0

0.5

1.0

1.5

2.0

2.5

Tota

l fric

tion

ener

gy (J

/m)

Bond Force (N)

Fig. 15. Total frictional energy generated at the contactinterface as a function of bond force.


always occurred at the perimeter of the contactinterface regardless of bond force, consistent withthe pressure distribution in Fig. 8. The frictionalenergy intensity at the perimeter was much higherthan that at the central area.

The total frictional energy per unit length wascalculated by integrating the friction energy inten-sity of Eq. (6) over the contact interface:

Ef ¼ZAn

4fblpðx; yÞdA. ð7Þ

Fig. 15 clearly shows that the total frictional en-ergy increases linearly with bond force. It wasshown that while the total frictional energy in-creased consistently with bond force, the high fric-tional energy intensity obtained at the periphery ofcontact interface did not show a similar increase.Thus, it can be said that a high bond force is notnecessarily beneficial for wire bondabilty. Instead,too high a bond force may impair the efficiency ofultrasonic energy transfer and may cause crateringfailures (Harman, 1997). Besides the bond force,

the effects of other wire bonding parameters canalso be envisaged from Eq. (7). A long weldingtime, a high frictional coefficient and a high vibra-tional speed may also increase the frictional energyintensity, which in turn increase the bondability.


The vibrational speed is the product of ultrasonicfrequency and amplitude, u = 4bf, and a highvibrational speed normally requires a high ultra-sonic power (Ramsey et al., 1997). However, to in-crease the ultrasonic frequency while maintainingthe constant ultrasonic power does not necessarilyenhance the vibration speed.

4. Conclusions

The deformation and stress distributions in thewire and bond pad during the ultrasonic wirebonding are analysed using the 2-D and 3-DFEM analyses. The following can be highlightedfrom the numerical study.

(1) The wire-bond pad contact interface had along elliptical shape, with the contact widthin the lateral direction being larger in thecentral area than at the wire ends. The max-imum contact pressure between the wire andbond pad occurred at the perimeter of thecontact interface, which is consistent withthe previous analysis. The absolute magni-tudes of the maximum contact pressureobtained from the 2-D and 3-D analysesagreed well, confirming the validity of thepresent 2-D model. A higher bond force doesnot mean a higher contact pressure, suggest-ing that a high bond force is not necessarilybeneficial to high wire bondability.

(2) The normalised real contact area calculatedbased on the 2-D analysis varied along thewire lateral direction in a similar manner tothe contact pressure. The plateau maximumvalue occurred always at the periphery overa length of 0.5–0.8 lm for all bond forcesstudied, which represents the most intimatereal contact area.

(3) The frictional energy is thought to be one ofthe most important factors determining suc-cessful bonding. The non-uniform pressuredistribution at the wire-bond pad contactinterface resulted in a non-uniform frictionalenergy intensity. The maximum frictionalenergy intensity occurred at the peripheryof the contact interface, where weld is prefer-

entially made as shown by many experimen-tal evidence. The total frictional energyincreased linearly with bond force, but thehigh frictional energy intensity obtained atthe periphery of the contact interface didnot show a similar increase.

Acknowledgements

This paper is presented at the InternationalSymposium on Macro-, Meso-, Micro- andNano-Mechanics of Materials (MM2003), whichis dedicated to Professor Pin Tong on the occasionof his 65th birthday. This work has been sup-ported by the Research Grant Council (RGC) ofHong Kong and the postdoctoral matching fundof HKUST. The first author (YD) was a visitingscholar when this work was performed.

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