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A fast mechanical test technique for life time estimation of micro-joints
Microelectronics Reliability 48 (2008) 1822–1830
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A fast mechanical test technique for life time estimation of micro-joints G. Khatibi a,*, W. Wroczewski a, B. Weiss a, T. Licht b a b
Faculty of Physics, University of Vienna, Boltzmanngasse 5, A-1090 Vienna, Austria Infineon Technologies, Max-Planck-Strasse 5, D-59581 Warstein, Germany
a r t i c l e
i n f o
Article history: Received 22 November 2007 Received in revised form 2 September 2008 Available online 17 October 2008
a b s t r a c t A novel accelerated mechanical testing method for reliability assessment of micro-joints in the electronic devices is presented as an alternative to time consuming thermal and power cycling test procedures. A special experimental set-up in combination with an ultrasonic resonance fatigue testing system and a laser Doppler vibrometer is used to obtain fatigue life curves of micro-joints under shear loading. Using this method fatigue life curves of Al wire bonded micro-joints were obtained up to 109 number of loading cycles and discussed with regard to micro-mechanisms of the bond failure. Failure analysis of the fatigued micro-joints showed that the predominant failure mechanism of power cycling tests, bond wire lift-off, was reproduced by the mechanical testing procedure. Life time of the micro-joints was modelled using a Coffin–Manson type relationship and showed a good correlation to lifetime curves obtained by power cycling tests. The major advantage of the proposed fast mechanical testing method is the significant reduction of the testing time in comparison with conventional thermal and power cycling tests. Furthermore subsequent examination of the failure surface provides a reliable tool for improvement of the bonding process. The proposed high frequency fatigue testing system can be applied as a rapid qualification and screening tool for various kinds of interconnects in electronic packaging. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction With increasing demands for innovative products and the implementation of new materials as well as short time-to-market requirements in packaging industry the challenge of producing highly reliable components raises. Life time of electronic devices, manufactured of multilayered structures of various physical and electrical properties, is mainly affected by failure induced due to thermomechanical fatigue as a result of mismatch of coefficient of thermal expansion in the interfaces of different layers. Interconnects, which serve as electrical connection as well as mechanical support between various components of the devices, are critical sites of failure due to thermal cycles as well as mechanical vibrations during the service life. Reliability of a system or component is defined as the ability to perform its required function under given conditions for a specified period of time [1]. Failure renders the device non-operational due to damage caused by a failure mechanism, actuated generally by external and/or internal stresses [2,3]. There are very few cases in which components may be tested under real time conditions with realistic stress–strain temperature histories. Common electronic system reliability programs include various board level tests and inspections in course of manufacturing process. Quality of the components is evaluated by a variety of established isothermal and * Corresponding author. Fax: +43 1 4277 9513. E-mail address: [email protected] (G. Khatibi). 0026-2714/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.microrel.2008.09.003
thermal tests based on relevant standards. These range from simple shear or pull test of single joints to the temperature and power cycling of the devices (e.g. [4–6]). Resistance of semiconductor devices and components and/or board assemblies to withstand mechanical stresses induced by alternating high and low temperature extremes is assessed by temperature cycling tests [5]. Power cycling is performed to determine the resistance of semiconductor devices to thermal and mechanical stresses due to cycling the power dissipation of the internal semiconductor die and internal connectors [6]. During the ‘‘on/off” periods of power cycling test procedures, the junction temperature between the die and the interconnect increases and decreases simulating the stresses induced during the actual circuit applications. Commonly used test procedures are accelerated power and temperature cycling with extreme temperature excursions, reducing the number of cycles to failure. One of the main limitations of accelerated testing is the possibility to invoke failure mechanisms at accelerated stress levels which are not those in the equipment operating range [3]. In comparison with thermal cycling, isothermal mechanical fatigue testing has been referred to be a great aid in microelectronic design due to the significantly shorter duration of testing time. By using an appropriate mechanical testing set-up, equivalent range of mechanical stress induced by temperature or power cycling procedure maybe achieved in significantly shorter periods of time. In the following a brief literature survey of selected mechanical testing techniques used in combination with monitoring and
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analysis of deformation and failure mechanisms during the fatigue process is given. These exemplary studies might give insight into advantages and some restrictions of mechanical fatigue testing in microelectronics. Uegai et al. [7] performed comparative thermal cycling and displacement controlled tension compression mechanical fatigue testing of SMT (surface mount technology) solders joints under various temperature ranges and frequencies up to 1 Hz. The strain range was evaluated by finite element analysis by considering the effect of temperature and strain on the stress–strain curve of the solder in the mechanical fatigue tests. Under these conditions, a linear relationship between maximum strain range of solder joints and the number of loading cycles to fracture was found. Based on these results a method for fatigue life prediction of solder joints by using mechanical fatigue testing was proposed. Ramminger et al. [8] presented a displacement controlled mechanical fatigue testing set-up to study heel crack failure in thick wire bonds of Aluminium accompanied by finite element modelling (FEM). Fatigue life was determined by using a Coffin– Manson type equation to correlate the calculated plastic strain to the number of loading cycles to fracture. Zhao et al. [9] conducted shock and harmonic vibration tests at room and at elevated temperature with different acceleration and frequency levels (50–200 Hz) to study the inelastic behavior of solder joints of BGA (ball grid array) packages. The experimental set up was a combination of an environmental chamber with a mechanical shaker and laser moire interferometry to measure the complete deformation field of the prepared specimen surface. The results showed that at elevated temperature, both shock and vibration induce significant inelastic shear deformation in solder joints, thus shorten the fatigue life of joints. The solder joints were found to suffer more severe damage at lower vibration frequencies than at higher ones. Xia Liu et al. [10] suggested a technique for visual inspection of BGAs during high cycle vibration testing consisting of a cyclic controlled curvature cantilever device with controls for varying the cycling frequency (300 Hz) and the magnitude of the applied load. The failures of solder interconnects was recorded by a direct visual monitoring method using a stroboscopic video system. Comparison of measured crack growth curves against number of cycles to failure for various loads and frequencies showed discrepancies with life prediction models based on damage mechanics. Results from this investigation show the inapplicability of damage mechanics models when conditions are different from those used to generate the models. Life time prognosis based on accelerated mechanical as well as thermal data is influenced by restriction of time and temperature dependent material properties such as strain sensitivity, stress relaxation, creep and microstructural changes [11]. It can be concluded that prediction of life time by mechanical test as an alternative to thermal tests procedures is then reliable, if the effect of these material dependent factors in data evaluation are taken into account. In present study a novel accelerated mechanical testing method for reliability assessment of micro-joints is presented and its applicability for life time evaluation of interconnects in electronic packaging is demonstrated. We have investigated failure of ultrasonically wedge bonded Al wires on silicon devices due to wire bond lift-off caused by mechanical fatigue. The results are compared with a similar failure case in silicon power devices caused by thermomechanical fatigue. For better understanding of our investigation, a brief introduction to wire bond lift-off failure mechanism in insulated gate bipolar transistors (IGBT) chips is given below. IGBTs are power semiconductor devices, applied for controlling large currents at high voltage and at high switching frequencies in
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Fig. 1. Schematic of multilayered structure of an IGBT module.
a broad range of industrial applications with high reliability demands. In power electronic packages electrical interconnection from IGBT chips to out put pins is provided by thick Al bonding wires. With up to 450 wires and 900 wedge bonds in high power multichip IGBT modules wire bond failure is reported to be one of the main lifetime concerns of the devices [12]. A schematic cross section of the multilayered structure of an IGBT chip with an Al wedge bonded wire is represented in Fig. 1. The Al wire is welded on thin Aluminium pads by using ultrasonic bonding technology which is one of the frequently applied interconnection methods in electronic packaging. From a technological point of view, the Al–Al bonding system has been reported to be extremely reliable and suitable for use in any thermal environment applicable for semiconductor devices [13]. Though the metallurgical aspects of the bonding process is not fully understood a phenomenological explanation of the ultrasonic bonding process is given in [14]. It was suggested that during the bonding process, the Al wire and the metallized bonding pad soften by ultrasonic energy and undergo plastic deformation due to the clamping force of the bonding tool. Vibration of the tool sweeps the brittle surface oxides and contaminants aside leaving clean metallic surfaces in contact. Contrary recent investigation shows partial incorporation of nanosized aluminum oxide particles in the interface [15]. Since this bonding process is without application of external heat, the required activation energy for interdiffusion is presumably supplied by ultrasonic energy. It had been reported that the mechanism of adhesion of wire bonds is the formation of interfaces similar to grain boundaries in polycrystalline metals [16]. The bonding interface consists of a layer of high dislocation density and high strength followed by a region of small recrystallized grains formed between the wire and pad interface [16]. Quality and strength of the bonding between the Al wire and the metallized surface of the silicon chip is controlled by bonding parameters and determines the reliability of the interconnection [12,17,18]. Due to the significant difference between the coefficients of thermal expansion of the silicon chip (approx. 2.6 ppm/K) and the Al wire (about 23.8 ppm/K), thermal cycles induced during operation of the devices cause shear stresses in the bonded interface of both materials resulting in thermomechanical fatigue in the bonded area. The highest stress concentration is induced at the terminations of the wire where cracks are initiated and grow into the bonded interface. Reduction of contact area results to an increase of current density in the remaining part of the wire and leads to acceleration of the fatigue process and fracture due to wire lift-off (Fig. 2) [12,18,19]. Other possible failure mechanisms in IGBT power modules such as bond wires heel cracks, reconstruction of Al-metallization on the chips or corrosion of wires is discussed in the relevant literature [20,21]. The life time of Al wire bonds in the modules is commonly evaluated by power cycling tests in which a load current is passed cyclically through several paralleled IGBT chips, the junction
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Fig. 2. Wire bond lift-off due to thermomechanical fatigue.
temperature is raised and lowered within short cycles until failure occurs. The failure criteria is usually defined as an increase of forward voltage (collector – emitter) by more than 5%. The tests are usually carried out at a semiconductor junction temperature excursion DTj in the range between 40 K and 80 K with a cycle time of about 10 s. For a temperature excursion of 40 K, about 10 million cycles to failure are expected, i.e. a test time of more than three years is necessary to reach degradation in the module properties [12,13]. Evaluation of power cycling tests is based on the presentation of the results by a Coffin–Manson type power law relationship in which DTj is related to number of cycles to failure [11,20]. In the following, the applied experimental procedure is described. Fatigue life curves obtained from Al bonded model joints are presented and related to involved fatigue failure mechanisms. The validity of our testing method is discussed by comparing the results with conventional power cycling test data of Al wire bonds in IGBT devices.
2. Experimental procedure 2.1. Principle of testing method The set-up of mechanical shear fatigue testing method for micro-joints is based on forced periodic motion of a system attached to a moving support. In this set-up, an ultrasonic resonance fatigue testing device is used to induce forced cyclic vibrations in a microcomponent specimen attached there-on. The ultrasonic fatigue testing system consists of a power supply in form of an oscillatoramplifier combination, a sonic energy generator, an acoustic horn and a specimen holder. The mechanical part consists of half-wavelength pieces which are excited to symmetrical longitudinal pushpull vibrations at a testing frequency of about 20 kHz. Distribution of displacement and strain varies sinusoidally along the parts of the system, with maximum strain occurring in mid-section and maximum displacement at the end of specimen holder [22]. The micro-component specimen is attached to the free end of the holder in a manner that coupling between specimen and holder is provided solely by a micro-joint (Fig. 3a). During the loading, the coupled holder and the micro-component specimen are exited to forced longitudinal vibrations at a frequency of 20 kHz at room temperature. Due to the inertia, the micro-component is accelerated relative to the holder generating a cyclic shear strain in the joint (Fig. 3b). The amplitude of the induced shear strain in the joint is related to the mass of the micro-component, the stiffness of micro-joint and their geometry. Depending on excitation amplitude and number of loading cycles, a micro-component specimen fails due to fatigue fracture of the micro-joint.
Fig. 3. (a) Schematic illustration of a micro-component attached to specimen holder, (b) schematic illustration of shear deformation of the micro-joint and (c) schematic illustration of specimen set up for shear fatigue testing of micro-joints.
The shear force F acting on a component with mass of m and acceleration a is given by
F¼m a
ð1Þ
The shear stress at a micro-joint with a certain contact area A can be calculated and with G being the shear Modulus of the joint material, shear strain can be derived
s ¼ F=A s¼G c
ð2Þ ð3Þ
Acceleration of the specimen and holder was determined by application of a laser doppler vibrometer (LDV) for measurements of the vibration velocity [23]. The velocity was measured parallel to the direction of motion at the end of the holder and micro-component as shown schematically in Fig. 3c. Peak acceleration is derived from the measured vibration velocity tmax by using Eq. (4), where f is the frequency.
amax ¼ 2pfvmax
ð4Þ
By variation of the location of the micro-component along the holder, the proposed set-up allows life time measurements of joints under various cyclic loading modes including shear, mixed mode of
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Fig. 4. (a) Typical wire bonded specimen and (b) specimen glued to holder or set up picture.
Table 1 Characterization of model joints bonded with Al wire of 400 lm diameter Lot No.
Al-metallization thickness (lm)
Si chip thickness (lm)
Si chip mass (mg)
Bonded area ave. (mm2)
A B C
4 4 3–5
280 280 140
40 40 20
0.13 ± 0.02 0.15 ± 0.015 0.20 ± 0.026
followed by a fine microstructure further toward the centre of the wire and the interface. The remaining microstructure consists of a mixed fibre structured and rather coarse grained annealed grains. Dark lines between the Al wire and Si chip are indications of non-bonded regions along the bonding interface. In the following the experimental results for selected Al wire bonded specimens are presented omitting the details of bonding parameters which are manufacturer’s internal data. 3. Results
Table 2 Mechanical properties of the Al wires used for various lots
3.1. Calibration measurements
Lot No.
Rp0.2 (MPa)
Rm (MPa)
A (%)
A B C
39 ± 1 45 ± 1 37 ± 1
45 ± 1 52 ± 1 50 ± 1
14 ± 3 12 ± 3 18 ± 3
shear – tension – compression and tension – compression. A detailed description of this method can be found in [24]. 2.2. Specimen geometry and characterization The micro-components used in this study consisted of single Al wires bonded to metallized silicon chips using various parameters as shown in Fig. 4 (Table 1). The bonding wires were glued in a groove machined near the free end of a specimen holder, so that coupling of chip to holder was provided by the bonded joint and the chip could vibrate freely during loading (Fig. 4b). The tensile properties of the Al wires used for the preparation of the joints were determined by using a micro tensile machine in combination with a laser speckle extensometer [10] (Table 2). Metallographic cross sections were prepared from selected asbonded and tested specimens to characterize the microstructure and quality of the bonded area and to study crack growth behaviour of the fatigued micro-joints by SEM and optical methods. A typical overview of specimens is presented in Fig. 5a showing the cross section of the Al bond wire on the silicon chip. Etched micro-sections of the wires reveal mostly an inhomogeneous grain size distribution in the Al bond wire with distinctive fine and coarse grained microstructural regions as shown in Fig. 5b. Few large grains are usually observed at both ends of the bonded region
Calibration measurements were performed to verify the applicability of the LDV system (Polytec model CLV-1000 with a peak velocity of 2.5 m/s and a resolution of 4 lm/s) for velocity measurements of the ultrasonic fatigue system. The vibration velocity of the system at the free end of the specimen holder was measured as a function of excitation amplitude and compared with the calculated velocity. The velocity amplitude was derived from the measured strain amplitude at the mid-section of the holder by using standard miniature strain gauges. Distribution of strain in a longitudinally vibrating free- free bar in half-wave resonance is given by Eq. (5) with k being wave length and x angular frequency. The maximum strain amplitude in the mid-centre of the bar is given by (6) from which displacement amplitude A0 is calculated and the peak velocity t at the end of the holder is derived (7).
2p 2p A0 sin x sin xt k k 2p e ¼ A0 k t ¼ x A0
e¼
ð5Þ ð6Þ ð7Þ
Measured values comply fairly well with calculated values with a deviation of approximately 6.5% (Fig. 6a). The influence of the specimen geometry on vibration behaviour of the coupled system as measured by a LDV is shown in Fig. 6b. The plot shows the relationship between peak velocity of two micro-components of different mass and joint area as a function of peak velocity of holder with increasing excitation amplitude. The measurements show that the holder and the chip vibrate at the same frequency and as expected an increase of mass of the
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Fig. 5. (a) Typical overview of cross section of Al bond wire on silicon chip and (b) Grain size distribution of Al in the bonded area.
Fig. 6. (a) Comparison of measured and calculated vibration velocity of fatigue testing system and (b) Influence of specimen geometry on vibration behaviour of the coupled system as measured by a LDV.
micro-component by a factor of two and also a slight reduction in contact area (curve with dashed lines) leads to higher peak velocity ratios. The laser doppler vibrometer has proven to be a suitable tool for non- contacting velocity measurements with high spatial resolution at ultrasonic frequencies providing the prerequisites for vibration measurements on small scaled structures [23]. 3.2. Fatigue life curves Fatigue life curves for three series of specimens are presented in Fig. 7 showing calculated values of shear strength as a function of number of loading cycles up to 108 and 109. Specimens of lot B with improved bonding parameters, demonstrate higher fatigue life particularly in the low cycle region. At higher numbers of loading cycles lots A and B show more or less similar behaviour while noticeable lower values are obtained for specimens of lot C. The shear stress range corresponding to lower number of loading cycles for all charges are between 40 MPa and 15 MPa while values from 32 MPa to 12 MPa are obtained for loading cycles higher than 107 showing relatively high scatter of stress values for all speci-
mens. The scatter in fatigue data of wire bonded joints is dependant on the production process and is also observed in power cycling tests of IGBT devices [20]. Similar trend was also observed in standard bond shear tests of the tested specimens. Evaluation and interpretation of fatigue life data of investigated specimens may be discussed with regard to the fracture mechanism incorporated during fatigue process which is primarily dependent on the quality of bonding interface and the mechanical properties of the used Al bonding wires. Metallographic investigations and fracture surface analysis of the fatigued specimens revealed three main types of failure mechanisms: fracture due to bond wire fatigue, mixed mode of wire and interface fracture, and fracture due to interface separation. Crack initiation and growth in fatigued specimens was studied on selected specimens loaded to certain number of cycles and amplitudes and on run-out specimens. SEM investigations of cross sections of several specimens revealed that as expected independent of amplitude or loading cycles. Fatigue crack initiates at the integration point of the bonding wire and silicon chip known to be the location of highest stress concentration of the joint (Fig. 8a). A micrograph with higher magnification shows that the
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Fig. 7. Fatigue life curves of selected model joints.
fatigue crack started from the transition area from non-bonded to bonded area along the interface and propagated along the grain boundaries into the bond wire (Fig. 8b). It is known that crack propagation in Aluminium, even in the early stages of cyclic deformation and low strain values occurs preferably along the grain boundaries as observed in most of the tested specimens which failed due to Al wire fatigue. Similar features were observed by Onuki et al. [17] who studied the effect of grain size on crack growth rate and failure of Al wire bonds. They concluded that because of the high strength of the interface between the Al wire and the pad, failure of the Al bonded joints should be due to the degradation of Al bonding wire [17]. It was suggested to improve the life time of the wire bonds by modification of the grains in the boundary region by means of annealing treatments. The crack growth path observed in the cross section of specimens can be correlated to the topographic features observed in SEM micrographs of fractured specimens. The topography of the fracture surface of a specimen failed due to Al wire fatigue can be observed in Fig. 9a with remnants of wire on the chip side of the component corresponding to missing material on fracture surface of the wire side (Fig. 9b). This mechanism seems to be the dominant mode of failure for long life wire bonded micro-joints with an almost defect free interface and high bonding quality. Pres-
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ence of clusters of pores or non-bonded areas generated during the ultrasonic bonding process in the interface of the micro joints may act as additional crack initiation sites mostly resulting in a combination of interface failure and Al wire fatigue (Fig. 9c and 9d). Interface fracture is characterized by a flat surface at both the wire side and the chip side resulting in premature fracture due to insufficient bonding strength between Al wire and metallization. Fig. 9e and 9f shows an example of a short life specimen with small amount of Al wire remnants in the centre and large flat areas at both ends of the foot print of the wire. These areas either correspond to not bonded or weakly bonded regions resulting in interface separation. Fracture is accelerated due to reduced contact area and additional stress concentration sites. It can be resumed that the main failure mechanism in our study is recognized to be bond wire lift-off due to Al wire fatigue, while the presence of defects and non-bonded areas mainly caused scatter in fatigue life data. In the case of fatigue due to degradation of Al, reliability of micro-joints is mainly influenced by the mechanical properties of the Al bonding wires, as demonstrated by improved fatigue properties of specimens of lot B with better mechanical properties of the bonding wires (Table 2). Since the bonding wires undergo deformation and recovery due to the ultrasonic bonding process, therefore the mechanical properties of Al in the bonded region might slightly deviate from those of non-bonded wires. Values given in Table 2 serve as a guide for tensile properties of Aluminium wire in the bonded region and not as absolute values. 3.3. Comparison of mechanical and thermal fatigue tests Thermal fatigue life of Al bond wires is often modelled by using the simple bimaterial approach where a is the coefficient of thermal expansion and DT the temperature swing in the junction (Eq. 8). It has been suggested that due to the large CTE mismatch between Al and Si, the total strain in the joint at high junction temperatures maybe regarded as plastic strain (De) [19]. Resuming Eqs. (8) and (9) a Coffin–Manson type relationship is obtained which describes the thermal fatigue life of the joints Eq. (10).
De ¼ ðaAl aSi Þ DT De / N f N f ¼ a DT n
ð8Þ ð9Þ ð10Þ
Using the above approach, a log–log plot of the number of loading cycles to failure as a function of DT was obtained from mechanical fatigue test data of lot C specimens and compared with thermal fatigue curves of IGBTs with similar micro-joints. Thermal fatigue data correspond to the life time of IGBTs with single emitter bond wires of 350 lm subjected to power cycles with a cycle period of
Fig. 8. (a) Crack path in a fatigued specimen (run-out specimen (De = 2.8 104, N = 1 109)) and (b) Detail of 8a, Integration point as crack initiation site and intergranular cracking in Al wire chip interface.
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Fig. 9. (a,c,e) SEM micrographs of failed specimens showing characteristic types of fatigue failure (chip side) and (b,d,f) SEM micrographs of failed specimens showing characteristic types of fatigue failure (corresponding wire side).
3 s, corresponding to testing frequency of 0.3 Hz, maximum junction temperature (Tjmax) of 373 K and a voltage drift of 10% from the initial value set as failure criterion [25]. As previously described mechanical fatigue data where obtained from model joints with different production parameters, complete separation of joint as failure criteria and loading frequency of 20 kHz. The plot shows that, cycles to failure is exponentially related to DT from which life time at lower junction temperatures can be predicted by extrapolation. Thermal and mechanical fatigue life lines run almost parallel to each other with an upward shift of the mechanical fatigue line to higher number of cycles to failure for a certain temperature (Fig. 10). Though the mechanical test procedure leads to an overestimation of life time of micro-joints, considering the principle differences between thermal and mechanical fatigue loading,
correlation of both data seems to be promising. An analysis of the test and specimen specific factors and the conditions involved in both methods permits the verification of the applicability of this approach. It is known that specimen specific parameters usually result in a certain variation of data but the influence of testing frequency and temperature dependent material parameters may be more significant. At conventional power cycling temperatures of maximum 373 K, Aluminium might undergo a recovery process leading to a slight decrease in hardness of the wire. One possible effect is the degradation of the Al metallization film known as reconstruction due to thermal cycling [20]. Creep of Aluminium bond wires at these temperature ranges are reported to be negligible [13]. Effect of cyclic test frequency on fatigue life has been subject of several studies and may be understood as an effect of plastic strain
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Fig. 10. Comparison of mechanical fatigue data with thermal fatigue curves of IGBTs.
rate on fracture and deformation behaviour of materials [26]. The main advantage of high frequency fatigue testing is the rapid and efficient life time determination. The remarkable time saving factor becomes evident by comparing standard accelerated power cycling procedures which require e.g. about 104 days to obtain the demanded 3 106 loading cycles [13] with 2.5 min loading time needed by ultrasonic testing method. Investigations of fatigue and crack growth behaviour of FCC materials at ultrasonic frequency have shown a negligible effect on life time at low homologous temperatures [27]. While these studies include mostly pure copper and Al alloys, reports on the influence of cyclic frequency on fatigue life of pure Al are scarce. Hoffelner et al. [28] studied the dependence of fatigue hardening behaviour of pure annealed Aluminium on testing frequency and strain amplitude by performing cyclic tension- compression tests and a subsequent determination of yield strength. An increase of yield strength by a factor of 2.3 was reported after 104 loading cycles at 20 kHz, while the amount of this effect for a frequency of 2 Hz was about 1.6 times of the initial value. The effect was more pronounced in frequency range from 2 Hz to 20 kHz and lower strain amplitudes (2.8 104) rather than 200 Hz to 20 kHz. Several factors including strain rate sensitivity and condition of material prior to cycling, also testing temperature, frequency and amplitude may affect the rate of cyclic hardening and fatigue life curves. The author’s own experiments on cyclic response of Al bond wires showed only moderate hardening effects. After 108 loading cycles at the strain amplitude of 4 104, an increase of 1.4 times the initial yield strength value for wires in annealed condition was measured while minor hardening effect on properties of hard wires was observed [12]. In present study, it may be expected that ultrasonic testing frequencies result in a slight shift of life time of Al micro-joints relative to thermal data. Investigation of thermal and mechanical fatigue life of identical bonded micro-joints are in progress which may allow a better comparison of life time prognosis of both methods. The plastic strain values calculated by the bimetallic approach and the resulting Coffin–Manson life time models are best approximations and not the true plastic strain value in the joint. FE analysis of thermally induced plastic strain in a single Al wedge bond on a Si chip under power cycling conditions shows a non-uniform
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strain distribution in the bonding area. The bonding interface shows highest strain values, with a maximum at integration points of wire and chip, decreasing toward the centre of the wire [19,29]. One model suggests a combination of mode I and mode II cracking with dominating shear mode due to CTE mismatch of Al and Si. The calculated average plastic strain for a step of 75 K at bond terminations is given to be in the range of 0.2% with a maximum of 2.7% Dl/ l occurring near the crack tip corresponding to crack initiation at 102–103 loading cycles [29]. Power cycling data shows that under actual test conditions, failure at a junction temperature range of 75 K usually occurs above 104 numbers of loading cycles [25]. Non-uniform strain distribution, crack initiation and failure modes of micro-joints due to power cycling tests resemble closely the deformation and fracture mechanisms involved in the mechanical shear fatigue procedure and are well described by FE analysis. Micro-sections of specimens subjected to mechanical shear fatigue reveal deformation gradients from the interface toward the centre of the wire. The development of fatigue cracks is concentrated in a rather narrow area near the interface while the bulk of wire either shows lower degrees of cyclic deformation or remains un-deformed (Fig. 2, 8a and 8b). Analogous to the crack initiation site in lift-off failure due to power cycling [12,18] and FE analysis [29,30], the fatal crack is always induced at the termination of bonded wire growing inwards the bonded area.
4. Summary and conclusion In this paper a novel high frequency mechanical shear testing method for fatigue life determination of micro-joints is introduced which operates at 20 kHz and can be used as an alternative to thermal and power cycling tests. The strain induced in the interfaces of the different materials due to thermal mismatch, is one of the main failure mechanism in microelectronic devices The basic idea of our method is to reduce the testing time by inducing an equivalent thermal strain in the interface of the specimens by mechanical means. The test set-up is based on the principles of forced cyclic vibration of coupled materials by using an ultrasonic resonance fatigue testing device. The developed system, is capable to generate mechanical cyclic shear strain in the interface of various micro-joints and bonded components to obtain fatigue life curves in very short testing times. Further, wire bonded Al micro-joints were used to evaluate the applicability of the proposed testing method for life time prediction of interconnects. The shear strength induced in the microjoints was calculated by using a Laser doppler vibrometry method and fatigue life curves (S–N curves) were plotted. The bimetallic law (D e = a DT) was used to approximate the equivalent temperature swing DT to the shear strain the micro-joints and model the life time by using the Coffin–Manson relationship. This approach enabled us to compare the results with the power cycling data of similar joints. It could be shown that the life time predicted by mechanical shear fatigue and power cycling indicate a very good correlation. The failure analysis showed that the predominant failure mechanism of power cycling tests, bond wire lift-off due to Al wire degradation is reproduced in the proposed mechanical test. The scatter in the fatigue life data could be related to the presence of defects and non-bonded areas. It can be concluded that the proposed method can be used not only to predict life time of micro-joints in significantly short periods of time, but also allows fast examination of failure surface. This provides useful information about the influence of bonding/interconnection technology and interaction of constituent materials on the reliability of the joint. The proposed method can be applied to various kinds of micro-joints such as wire bonds and solder
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joints as an alternative to conventional thermal cycling tests. It can be used as a reliable rapid product qualification and screening tool in the interconnect technology, reducing the time and associated costs during the design and production process. Acknowledgements The authors wish to gratefully thank FFG – Bridge program for financial support (Project No. 811022/9455 – SCK/SAI), Dr. G. Mitic, Dr. P. Zimprich, Prof. V. Groeger and Prof. R. Stickler for valuable discussions and reviewing of the manuscript. We are grateful to Mr. G. Lefranc for initiating this investigation and providing Fig. 2. References [1] Martin P. Electronic failure analysis handbook. USA: McGraw-Hill; 1999. [2] Dasgupta A, Pecht M. Material failure mechanisms and damage models. IEEE Trans Reliab 1991;40(5):531–6. [3] Lall P. Tutorial: temperature as an input to microelectronics – reliability models. IEEE Trans Reliab 1996;45(1):3–9. [4] EIA/JESD22-B116 July 1998. Wire bond shear test method. [5] IEC 60749-25 ED. 1.0 B: 2003. Semiconductor devices-mechanical and climatic test methods-Parts 25: Temperature cycling. [6] IEC 60749-34 Ed. 1.0 B: 2005. Semiconductor devices-mechanical and climatic test methods-Parts 34: Power cycling. [7] Uegai Y, Tani S, Inoue A, Yoshioka S, Tamura K. A method of fatigue life prediction for surface mount solder joints of electronic devices by mechanical fatigue test. Proc 2nd ASME Int Elect Packag Conf 1993;1:493–8. [8] Ramminger S, Seliger N, Wachutka G. Reliability model for Al wire bonds subjected to heel crack failures. Microelect Reliab 2000;40:1521–5. [9] Zhao Y, Basaran C, Cartwright A, Dishongh T. Thermomechanical behaviour of micron scale solder joints under dynamic loads. Mech Mater 2000;32(3):161–73. [10] Xia Liu K, Valmiki A, Sooklal A, Melod B, Verges C, Michael AL. Experimental study and life prediction on high cycle vibration fatigue in BGA packages. Microelect Reliab 2006;46:1128–38. [11] Plumbridge WJ, Matela RJ, Westwater A. Structural integrity and reliability in electronics. London: Kluwer Academic Publishers; 2003.
[12] Lefranc G, Weiss B, Klos C, Dick J, Khatibi G, Berg H. Microelect Reliab 2003;43:1833–8. [13] Berg H, Wolfgang E. Advanced IGBT modules for railway traction applications, reliability testing. Microelect Reliab 1998;38:1319–23. [14] Harman G. Wire bonding in microelectronics: materials, processes, reliability and yield. New York: McGraw-Hill; 1997. [15] Geissler U, Engelmann H-J, Urban I, Rooch H. FIB preparation and TEM analytics on AlSi1 bond pads. Pract Metallogr 2006;42(10):520–32. [16] Krzanowski J. A transmission electron microscopy study of ultrasonic wire bonding. IEEE Trans Components Hybrids Manuf Technol 1990;13(1):176–81. [17] Onuki J, Koizumi M, Echigoya J. Microstructural analysis of the bonding interface between thick Al wire and Al-1 mass%Si electrode film. Mater Trans JIM 1996;37:1324–31. [18] Onuki J, Koizumi M, Suwa M. Reliability of thick Al wire bonds in IGBT modules for traction motor drives. IEEE Trans Adv Packag 2000;23(1):108–12. [19] Ramminger S, Türkes P, Wachutka G. Crack mechanism in wire bonding technology. Microelect Reliab 1998;38:1301–5. [20] Ciappa M. Selected failure mechanisms of modern power modules. Microelect Reliab 2002;42:653–67. [21] Hua Ye, Minghui Lin, Basaran C. Failure modes and FEM analysis of power electronic packaging. Finite Elements Anal Des 2002;38:601–12. [22] Roth LD. Ultrasonic fatigue testing. Metals handbook, 9th ed., vol. 8. Ohio: ASM; 1985. p. 241–57. [23] Castellini P, Martarelli M, Tomasini EP. Laser Doppler vibrometry: development of advanced solutions answering to technology’s needs. Mech Syst Signal Process 2006;20:1265–85. [24] Khatibi G, Groeger V, Weiss B, Lefranc G, Mitic G. Patent No. 10 2005 016038B3 2006.12.28. [25] Cova P, Fantini F. On the effect of power cycling stress on IGBT modules. Microelect Reliab 1998;38:1347–52. [26] Tien JK. The state of ultrasonic fatigue. In: Ultrasonic Fatigue, ed. Wells, Buck, Roth, Tien, Conf Proc AIME; 1982. p. 1–14. [27] Weiss B, Mullner H, Stickler R, Lukas P, Kunz L. Influence of frequency on fatigue limit and fatigue crack growth behavior of polycrystalline Cu. Adv Fract Res 1984;3:1783–90. Fracture 84. ICG6. [28] Hoffelner W. The effect of test frequency on the fatigue behaviour of aluminium. Scripta Metallurgica 1975;9:1157–61. [29] Ramminger S, Mitic G, Türkes P, Wachutka G. Thermo-mechanical simulation of wire bonding joints in power modules. Proc Int Conf Model Simul Microsyst; 1999. p. 483–6. [30] Wilde J. Lebensdauerprognose von Drahtbondverbindungen für die Mechatronik mittels FEM. GMM-Fachbericht 2002;36:117–22.
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Microelectronics Reliability journal homepage: www.elsevier.com/locate/microrel
A fast mechanical test technique for life time estimation of micro-joints G. Khatibi a,*, W. Wroczewski a, B. Weiss a, T. Licht b a b
Faculty of Physics, University of Vienna, Boltzmanngasse 5, A-1090 Vienna, Austria Infineon Technologies, Max-Planck-Strasse 5, D-59581 Warstein, Germany
a r t i c l e
i n f o
Article history: Received 22 November 2007 Received in revised form 2 September 2008 Available online 17 October 2008
a b s t r a c t A novel accelerated mechanical testing method for reliability assessment of micro-joints in the electronic devices is presented as an alternative to time consuming thermal and power cycling test procedures. A special experimental set-up in combination with an ultrasonic resonance fatigue testing system and a laser Doppler vibrometer is used to obtain fatigue life curves of micro-joints under shear loading. Using this method fatigue life curves of Al wire bonded micro-joints were obtained up to 109 number of loading cycles and discussed with regard to micro-mechanisms of the bond failure. Failure analysis of the fatigued micro-joints showed that the predominant failure mechanism of power cycling tests, bond wire lift-off, was reproduced by the mechanical testing procedure. Life time of the micro-joints was modelled using a Coffin–Manson type relationship and showed a good correlation to lifetime curves obtained by power cycling tests. The major advantage of the proposed fast mechanical testing method is the significant reduction of the testing time in comparison with conventional thermal and power cycling tests. Furthermore subsequent examination of the failure surface provides a reliable tool for improvement of the bonding process. The proposed high frequency fatigue testing system can be applied as a rapid qualification and screening tool for various kinds of interconnects in electronic packaging. Ó 2008 Elsevier Ltd. All rights reserved.
1. Introduction With increasing demands for innovative products and the implementation of new materials as well as short time-to-market requirements in packaging industry the challenge of producing highly reliable components raises. Life time of electronic devices, manufactured of multilayered structures of various physical and electrical properties, is mainly affected by failure induced due to thermomechanical fatigue as a result of mismatch of coefficient of thermal expansion in the interfaces of different layers. Interconnects, which serve as electrical connection as well as mechanical support between various components of the devices, are critical sites of failure due to thermal cycles as well as mechanical vibrations during the service life. Reliability of a system or component is defined as the ability to perform its required function under given conditions for a specified period of time [1]. Failure renders the device non-operational due to damage caused by a failure mechanism, actuated generally by external and/or internal stresses [2,3]. There are very few cases in which components may be tested under real time conditions with realistic stress–strain temperature histories. Common electronic system reliability programs include various board level tests and inspections in course of manufacturing process. Quality of the components is evaluated by a variety of established isothermal and * Corresponding author. Fax: +43 1 4277 9513. E-mail address: [email protected] (G. Khatibi). 0026-2714/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.microrel.2008.09.003
thermal tests based on relevant standards. These range from simple shear or pull test of single joints to the temperature and power cycling of the devices (e.g. [4–6]). Resistance of semiconductor devices and components and/or board assemblies to withstand mechanical stresses induced by alternating high and low temperature extremes is assessed by temperature cycling tests [5]. Power cycling is performed to determine the resistance of semiconductor devices to thermal and mechanical stresses due to cycling the power dissipation of the internal semiconductor die and internal connectors [6]. During the ‘‘on/off” periods of power cycling test procedures, the junction temperature between the die and the interconnect increases and decreases simulating the stresses induced during the actual circuit applications. Commonly used test procedures are accelerated power and temperature cycling with extreme temperature excursions, reducing the number of cycles to failure. One of the main limitations of accelerated testing is the possibility to invoke failure mechanisms at accelerated stress levels which are not those in the equipment operating range [3]. In comparison with thermal cycling, isothermal mechanical fatigue testing has been referred to be a great aid in microelectronic design due to the significantly shorter duration of testing time. By using an appropriate mechanical testing set-up, equivalent range of mechanical stress induced by temperature or power cycling procedure maybe achieved in significantly shorter periods of time. In the following a brief literature survey of selected mechanical testing techniques used in combination with monitoring and
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analysis of deformation and failure mechanisms during the fatigue process is given. These exemplary studies might give insight into advantages and some restrictions of mechanical fatigue testing in microelectronics. Uegai et al. [7] performed comparative thermal cycling and displacement controlled tension compression mechanical fatigue testing of SMT (surface mount technology) solders joints under various temperature ranges and frequencies up to 1 Hz. The strain range was evaluated by finite element analysis by considering the effect of temperature and strain on the stress–strain curve of the solder in the mechanical fatigue tests. Under these conditions, a linear relationship between maximum strain range of solder joints and the number of loading cycles to fracture was found. Based on these results a method for fatigue life prediction of solder joints by using mechanical fatigue testing was proposed. Ramminger et al. [8] presented a displacement controlled mechanical fatigue testing set-up to study heel crack failure in thick wire bonds of Aluminium accompanied by finite element modelling (FEM). Fatigue life was determined by using a Coffin– Manson type equation to correlate the calculated plastic strain to the number of loading cycles to fracture. Zhao et al. [9] conducted shock and harmonic vibration tests at room and at elevated temperature with different acceleration and frequency levels (50–200 Hz) to study the inelastic behavior of solder joints of BGA (ball grid array) packages. The experimental set up was a combination of an environmental chamber with a mechanical shaker and laser moire interferometry to measure the complete deformation field of the prepared specimen surface. The results showed that at elevated temperature, both shock and vibration induce significant inelastic shear deformation in solder joints, thus shorten the fatigue life of joints. The solder joints were found to suffer more severe damage at lower vibration frequencies than at higher ones. Xia Liu et al. [10] suggested a technique for visual inspection of BGAs during high cycle vibration testing consisting of a cyclic controlled curvature cantilever device with controls for varying the cycling frequency (300 Hz) and the magnitude of the applied load. The failures of solder interconnects was recorded by a direct visual monitoring method using a stroboscopic video system. Comparison of measured crack growth curves against number of cycles to failure for various loads and frequencies showed discrepancies with life prediction models based on damage mechanics. Results from this investigation show the inapplicability of damage mechanics models when conditions are different from those used to generate the models. Life time prognosis based on accelerated mechanical as well as thermal data is influenced by restriction of time and temperature dependent material properties such as strain sensitivity, stress relaxation, creep and microstructural changes [11]. It can be concluded that prediction of life time by mechanical test as an alternative to thermal tests procedures is then reliable, if the effect of these material dependent factors in data evaluation are taken into account. In present study a novel accelerated mechanical testing method for reliability assessment of micro-joints is presented and its applicability for life time evaluation of interconnects in electronic packaging is demonstrated. We have investigated failure of ultrasonically wedge bonded Al wires on silicon devices due to wire bond lift-off caused by mechanical fatigue. The results are compared with a similar failure case in silicon power devices caused by thermomechanical fatigue. For better understanding of our investigation, a brief introduction to wire bond lift-off failure mechanism in insulated gate bipolar transistors (IGBT) chips is given below. IGBTs are power semiconductor devices, applied for controlling large currents at high voltage and at high switching frequencies in
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Fig. 1. Schematic of multilayered structure of an IGBT module.
a broad range of industrial applications with high reliability demands. In power electronic packages electrical interconnection from IGBT chips to out put pins is provided by thick Al bonding wires. With up to 450 wires and 900 wedge bonds in high power multichip IGBT modules wire bond failure is reported to be one of the main lifetime concerns of the devices [12]. A schematic cross section of the multilayered structure of an IGBT chip with an Al wedge bonded wire is represented in Fig. 1. The Al wire is welded on thin Aluminium pads by using ultrasonic bonding technology which is one of the frequently applied interconnection methods in electronic packaging. From a technological point of view, the Al–Al bonding system has been reported to be extremely reliable and suitable for use in any thermal environment applicable for semiconductor devices [13]. Though the metallurgical aspects of the bonding process is not fully understood a phenomenological explanation of the ultrasonic bonding process is given in [14]. It was suggested that during the bonding process, the Al wire and the metallized bonding pad soften by ultrasonic energy and undergo plastic deformation due to the clamping force of the bonding tool. Vibration of the tool sweeps the brittle surface oxides and contaminants aside leaving clean metallic surfaces in contact. Contrary recent investigation shows partial incorporation of nanosized aluminum oxide particles in the interface [15]. Since this bonding process is without application of external heat, the required activation energy for interdiffusion is presumably supplied by ultrasonic energy. It had been reported that the mechanism of adhesion of wire bonds is the formation of interfaces similar to grain boundaries in polycrystalline metals [16]. The bonding interface consists of a layer of high dislocation density and high strength followed by a region of small recrystallized grains formed between the wire and pad interface [16]. Quality and strength of the bonding between the Al wire and the metallized surface of the silicon chip is controlled by bonding parameters and determines the reliability of the interconnection [12,17,18]. Due to the significant difference between the coefficients of thermal expansion of the silicon chip (approx. 2.6 ppm/K) and the Al wire (about 23.8 ppm/K), thermal cycles induced during operation of the devices cause shear stresses in the bonded interface of both materials resulting in thermomechanical fatigue in the bonded area. The highest stress concentration is induced at the terminations of the wire where cracks are initiated and grow into the bonded interface. Reduction of contact area results to an increase of current density in the remaining part of the wire and leads to acceleration of the fatigue process and fracture due to wire lift-off (Fig. 2) [12,18,19]. Other possible failure mechanisms in IGBT power modules such as bond wires heel cracks, reconstruction of Al-metallization on the chips or corrosion of wires is discussed in the relevant literature [20,21]. The life time of Al wire bonds in the modules is commonly evaluated by power cycling tests in which a load current is passed cyclically through several paralleled IGBT chips, the junction
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Fig. 2. Wire bond lift-off due to thermomechanical fatigue.
temperature is raised and lowered within short cycles until failure occurs. The failure criteria is usually defined as an increase of forward voltage (collector – emitter) by more than 5%. The tests are usually carried out at a semiconductor junction temperature excursion DTj in the range between 40 K and 80 K with a cycle time of about 10 s. For a temperature excursion of 40 K, about 10 million cycles to failure are expected, i.e. a test time of more than three years is necessary to reach degradation in the module properties [12,13]. Evaluation of power cycling tests is based on the presentation of the results by a Coffin–Manson type power law relationship in which DTj is related to number of cycles to failure [11,20]. In the following, the applied experimental procedure is described. Fatigue life curves obtained from Al bonded model joints are presented and related to involved fatigue failure mechanisms. The validity of our testing method is discussed by comparing the results with conventional power cycling test data of Al wire bonds in IGBT devices.
2. Experimental procedure 2.1. Principle of testing method The set-up of mechanical shear fatigue testing method for micro-joints is based on forced periodic motion of a system attached to a moving support. In this set-up, an ultrasonic resonance fatigue testing device is used to induce forced cyclic vibrations in a microcomponent specimen attached there-on. The ultrasonic fatigue testing system consists of a power supply in form of an oscillatoramplifier combination, a sonic energy generator, an acoustic horn and a specimen holder. The mechanical part consists of half-wavelength pieces which are excited to symmetrical longitudinal pushpull vibrations at a testing frequency of about 20 kHz. Distribution of displacement and strain varies sinusoidally along the parts of the system, with maximum strain occurring in mid-section and maximum displacement at the end of specimen holder [22]. The micro-component specimen is attached to the free end of the holder in a manner that coupling between specimen and holder is provided solely by a micro-joint (Fig. 3a). During the loading, the coupled holder and the micro-component specimen are exited to forced longitudinal vibrations at a frequency of 20 kHz at room temperature. Due to the inertia, the micro-component is accelerated relative to the holder generating a cyclic shear strain in the joint (Fig. 3b). The amplitude of the induced shear strain in the joint is related to the mass of the micro-component, the stiffness of micro-joint and their geometry. Depending on excitation amplitude and number of loading cycles, a micro-component specimen fails due to fatigue fracture of the micro-joint.
Fig. 3. (a) Schematic illustration of a micro-component attached to specimen holder, (b) schematic illustration of shear deformation of the micro-joint and (c) schematic illustration of specimen set up for shear fatigue testing of micro-joints.
The shear force F acting on a component with mass of m and acceleration a is given by
F¼m a
ð1Þ
The shear stress at a micro-joint with a certain contact area A can be calculated and with G being the shear Modulus of the joint material, shear strain can be derived
s ¼ F=A s¼G c
ð2Þ ð3Þ
Acceleration of the specimen and holder was determined by application of a laser doppler vibrometer (LDV) for measurements of the vibration velocity [23]. The velocity was measured parallel to the direction of motion at the end of the holder and micro-component as shown schematically in Fig. 3c. Peak acceleration is derived from the measured vibration velocity tmax by using Eq. (4), where f is the frequency.
amax ¼ 2pfvmax
ð4Þ
By variation of the location of the micro-component along the holder, the proposed set-up allows life time measurements of joints under various cyclic loading modes including shear, mixed mode of
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Fig. 4. (a) Typical wire bonded specimen and (b) specimen glued to holder or set up picture.
Table 1 Characterization of model joints bonded with Al wire of 400 lm diameter Lot No.
Al-metallization thickness (lm)
Si chip thickness (lm)
Si chip mass (mg)
Bonded area ave. (mm2)
A B C
4 4 3–5
280 280 140
40 40 20
0.13 ± 0.02 0.15 ± 0.015 0.20 ± 0.026
followed by a fine microstructure further toward the centre of the wire and the interface. The remaining microstructure consists of a mixed fibre structured and rather coarse grained annealed grains. Dark lines between the Al wire and Si chip are indications of non-bonded regions along the bonding interface. In the following the experimental results for selected Al wire bonded specimens are presented omitting the details of bonding parameters which are manufacturer’s internal data. 3. Results
Table 2 Mechanical properties of the Al wires used for various lots
3.1. Calibration measurements
Lot No.
Rp0.2 (MPa)
Rm (MPa)
A (%)
A B C
39 ± 1 45 ± 1 37 ± 1
45 ± 1 52 ± 1 50 ± 1
14 ± 3 12 ± 3 18 ± 3
shear – tension – compression and tension – compression. A detailed description of this method can be found in [24]. 2.2. Specimen geometry and characterization The micro-components used in this study consisted of single Al wires bonded to metallized silicon chips using various parameters as shown in Fig. 4 (Table 1). The bonding wires were glued in a groove machined near the free end of a specimen holder, so that coupling of chip to holder was provided by the bonded joint and the chip could vibrate freely during loading (Fig. 4b). The tensile properties of the Al wires used for the preparation of the joints were determined by using a micro tensile machine in combination with a laser speckle extensometer [10] (Table 2). Metallographic cross sections were prepared from selected asbonded and tested specimens to characterize the microstructure and quality of the bonded area and to study crack growth behaviour of the fatigued micro-joints by SEM and optical methods. A typical overview of specimens is presented in Fig. 5a showing the cross section of the Al bond wire on the silicon chip. Etched micro-sections of the wires reveal mostly an inhomogeneous grain size distribution in the Al bond wire with distinctive fine and coarse grained microstructural regions as shown in Fig. 5b. Few large grains are usually observed at both ends of the bonded region
Calibration measurements were performed to verify the applicability of the LDV system (Polytec model CLV-1000 with a peak velocity of 2.5 m/s and a resolution of 4 lm/s) for velocity measurements of the ultrasonic fatigue system. The vibration velocity of the system at the free end of the specimen holder was measured as a function of excitation amplitude and compared with the calculated velocity. The velocity amplitude was derived from the measured strain amplitude at the mid-section of the holder by using standard miniature strain gauges. Distribution of strain in a longitudinally vibrating free- free bar in half-wave resonance is given by Eq. (5) with k being wave length and x angular frequency. The maximum strain amplitude in the mid-centre of the bar is given by (6) from which displacement amplitude A0 is calculated and the peak velocity t at the end of the holder is derived (7).
2p 2p A0 sin x sin xt k k 2p e ¼ A0 k t ¼ x A0
e¼
ð5Þ ð6Þ ð7Þ
Measured values comply fairly well with calculated values with a deviation of approximately 6.5% (Fig. 6a). The influence of the specimen geometry on vibration behaviour of the coupled system as measured by a LDV is shown in Fig. 6b. The plot shows the relationship between peak velocity of two micro-components of different mass and joint area as a function of peak velocity of holder with increasing excitation amplitude. The measurements show that the holder and the chip vibrate at the same frequency and as expected an increase of mass of the
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Fig. 5. (a) Typical overview of cross section of Al bond wire on silicon chip and (b) Grain size distribution of Al in the bonded area.
Fig. 6. (a) Comparison of measured and calculated vibration velocity of fatigue testing system and (b) Influence of specimen geometry on vibration behaviour of the coupled system as measured by a LDV.
micro-component by a factor of two and also a slight reduction in contact area (curve with dashed lines) leads to higher peak velocity ratios. The laser doppler vibrometer has proven to be a suitable tool for non- contacting velocity measurements with high spatial resolution at ultrasonic frequencies providing the prerequisites for vibration measurements on small scaled structures [23]. 3.2. Fatigue life curves Fatigue life curves for three series of specimens are presented in Fig. 7 showing calculated values of shear strength as a function of number of loading cycles up to 108 and 109. Specimens of lot B with improved bonding parameters, demonstrate higher fatigue life particularly in the low cycle region. At higher numbers of loading cycles lots A and B show more or less similar behaviour while noticeable lower values are obtained for specimens of lot C. The shear stress range corresponding to lower number of loading cycles for all charges are between 40 MPa and 15 MPa while values from 32 MPa to 12 MPa are obtained for loading cycles higher than 107 showing relatively high scatter of stress values for all speci-
mens. The scatter in fatigue data of wire bonded joints is dependant on the production process and is also observed in power cycling tests of IGBT devices [20]. Similar trend was also observed in standard bond shear tests of the tested specimens. Evaluation and interpretation of fatigue life data of investigated specimens may be discussed with regard to the fracture mechanism incorporated during fatigue process which is primarily dependent on the quality of bonding interface and the mechanical properties of the used Al bonding wires. Metallographic investigations and fracture surface analysis of the fatigued specimens revealed three main types of failure mechanisms: fracture due to bond wire fatigue, mixed mode of wire and interface fracture, and fracture due to interface separation. Crack initiation and growth in fatigued specimens was studied on selected specimens loaded to certain number of cycles and amplitudes and on run-out specimens. SEM investigations of cross sections of several specimens revealed that as expected independent of amplitude or loading cycles. Fatigue crack initiates at the integration point of the bonding wire and silicon chip known to be the location of highest stress concentration of the joint (Fig. 8a). A micrograph with higher magnification shows that the
G. Khatibi et al. / Microelectronics Reliability 48 (2008) 1822–1830
Fig. 7. Fatigue life curves of selected model joints.
fatigue crack started from the transition area from non-bonded to bonded area along the interface and propagated along the grain boundaries into the bond wire (Fig. 8b). It is known that crack propagation in Aluminium, even in the early stages of cyclic deformation and low strain values occurs preferably along the grain boundaries as observed in most of the tested specimens which failed due to Al wire fatigue. Similar features were observed by Onuki et al. [17] who studied the effect of grain size on crack growth rate and failure of Al wire bonds. They concluded that because of the high strength of the interface between the Al wire and the pad, failure of the Al bonded joints should be due to the degradation of Al bonding wire [17]. It was suggested to improve the life time of the wire bonds by modification of the grains in the boundary region by means of annealing treatments. The crack growth path observed in the cross section of specimens can be correlated to the topographic features observed in SEM micrographs of fractured specimens. The topography of the fracture surface of a specimen failed due to Al wire fatigue can be observed in Fig. 9a with remnants of wire on the chip side of the component corresponding to missing material on fracture surface of the wire side (Fig. 9b). This mechanism seems to be the dominant mode of failure for long life wire bonded micro-joints with an almost defect free interface and high bonding quality. Pres-
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ence of clusters of pores or non-bonded areas generated during the ultrasonic bonding process in the interface of the micro joints may act as additional crack initiation sites mostly resulting in a combination of interface failure and Al wire fatigue (Fig. 9c and 9d). Interface fracture is characterized by a flat surface at both the wire side and the chip side resulting in premature fracture due to insufficient bonding strength between Al wire and metallization. Fig. 9e and 9f shows an example of a short life specimen with small amount of Al wire remnants in the centre and large flat areas at both ends of the foot print of the wire. These areas either correspond to not bonded or weakly bonded regions resulting in interface separation. Fracture is accelerated due to reduced contact area and additional stress concentration sites. It can be resumed that the main failure mechanism in our study is recognized to be bond wire lift-off due to Al wire fatigue, while the presence of defects and non-bonded areas mainly caused scatter in fatigue life data. In the case of fatigue due to degradation of Al, reliability of micro-joints is mainly influenced by the mechanical properties of the Al bonding wires, as demonstrated by improved fatigue properties of specimens of lot B with better mechanical properties of the bonding wires (Table 2). Since the bonding wires undergo deformation and recovery due to the ultrasonic bonding process, therefore the mechanical properties of Al in the bonded region might slightly deviate from those of non-bonded wires. Values given in Table 2 serve as a guide for tensile properties of Aluminium wire in the bonded region and not as absolute values. 3.3. Comparison of mechanical and thermal fatigue tests Thermal fatigue life of Al bond wires is often modelled by using the simple bimaterial approach where a is the coefficient of thermal expansion and DT the temperature swing in the junction (Eq. 8). It has been suggested that due to the large CTE mismatch between Al and Si, the total strain in the joint at high junction temperatures maybe regarded as plastic strain (De) [19]. Resuming Eqs. (8) and (9) a Coffin–Manson type relationship is obtained which describes the thermal fatigue life of the joints Eq. (10).
De ¼ ðaAl aSi Þ DT De / N f N f ¼ a DT n
ð8Þ ð9Þ ð10Þ
Using the above approach, a log–log plot of the number of loading cycles to failure as a function of DT was obtained from mechanical fatigue test data of lot C specimens and compared with thermal fatigue curves of IGBTs with similar micro-joints. Thermal fatigue data correspond to the life time of IGBTs with single emitter bond wires of 350 lm subjected to power cycles with a cycle period of
Fig. 8. (a) Crack path in a fatigued specimen (run-out specimen (De = 2.8 104, N = 1 109)) and (b) Detail of 8a, Integration point as crack initiation site and intergranular cracking in Al wire chip interface.
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Fig. 9. (a,c,e) SEM micrographs of failed specimens showing characteristic types of fatigue failure (chip side) and (b,d,f) SEM micrographs of failed specimens showing characteristic types of fatigue failure (corresponding wire side).
3 s, corresponding to testing frequency of 0.3 Hz, maximum junction temperature (Tjmax) of 373 K and a voltage drift of 10% from the initial value set as failure criterion [25]. As previously described mechanical fatigue data where obtained from model joints with different production parameters, complete separation of joint as failure criteria and loading frequency of 20 kHz. The plot shows that, cycles to failure is exponentially related to DT from which life time at lower junction temperatures can be predicted by extrapolation. Thermal and mechanical fatigue life lines run almost parallel to each other with an upward shift of the mechanical fatigue line to higher number of cycles to failure for a certain temperature (Fig. 10). Though the mechanical test procedure leads to an overestimation of life time of micro-joints, considering the principle differences between thermal and mechanical fatigue loading,
correlation of both data seems to be promising. An analysis of the test and specimen specific factors and the conditions involved in both methods permits the verification of the applicability of this approach. It is known that specimen specific parameters usually result in a certain variation of data but the influence of testing frequency and temperature dependent material parameters may be more significant. At conventional power cycling temperatures of maximum 373 K, Aluminium might undergo a recovery process leading to a slight decrease in hardness of the wire. One possible effect is the degradation of the Al metallization film known as reconstruction due to thermal cycling [20]. Creep of Aluminium bond wires at these temperature ranges are reported to be negligible [13]. Effect of cyclic test frequency on fatigue life has been subject of several studies and may be understood as an effect of plastic strain
G. Khatibi et al. / Microelectronics Reliability 48 (2008) 1822–1830
Fig. 10. Comparison of mechanical fatigue data with thermal fatigue curves of IGBTs.
rate on fracture and deformation behaviour of materials [26]. The main advantage of high frequency fatigue testing is the rapid and efficient life time determination. The remarkable time saving factor becomes evident by comparing standard accelerated power cycling procedures which require e.g. about 104 days to obtain the demanded 3 106 loading cycles [13] with 2.5 min loading time needed by ultrasonic testing method. Investigations of fatigue and crack growth behaviour of FCC materials at ultrasonic frequency have shown a negligible effect on life time at low homologous temperatures [27]. While these studies include mostly pure copper and Al alloys, reports on the influence of cyclic frequency on fatigue life of pure Al are scarce. Hoffelner et al. [28] studied the dependence of fatigue hardening behaviour of pure annealed Aluminium on testing frequency and strain amplitude by performing cyclic tension- compression tests and a subsequent determination of yield strength. An increase of yield strength by a factor of 2.3 was reported after 104 loading cycles at 20 kHz, while the amount of this effect for a frequency of 2 Hz was about 1.6 times of the initial value. The effect was more pronounced in frequency range from 2 Hz to 20 kHz and lower strain amplitudes (2.8 104) rather than 200 Hz to 20 kHz. Several factors including strain rate sensitivity and condition of material prior to cycling, also testing temperature, frequency and amplitude may affect the rate of cyclic hardening and fatigue life curves. The author’s own experiments on cyclic response of Al bond wires showed only moderate hardening effects. After 108 loading cycles at the strain amplitude of 4 104, an increase of 1.4 times the initial yield strength value for wires in annealed condition was measured while minor hardening effect on properties of hard wires was observed [12]. In present study, it may be expected that ultrasonic testing frequencies result in a slight shift of life time of Al micro-joints relative to thermal data. Investigation of thermal and mechanical fatigue life of identical bonded micro-joints are in progress which may allow a better comparison of life time prognosis of both methods. The plastic strain values calculated by the bimetallic approach and the resulting Coffin–Manson life time models are best approximations and not the true plastic strain value in the joint. FE analysis of thermally induced plastic strain in a single Al wedge bond on a Si chip under power cycling conditions shows a non-uniform
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strain distribution in the bonding area. The bonding interface shows highest strain values, with a maximum at integration points of wire and chip, decreasing toward the centre of the wire [19,29]. One model suggests a combination of mode I and mode II cracking with dominating shear mode due to CTE mismatch of Al and Si. The calculated average plastic strain for a step of 75 K at bond terminations is given to be in the range of 0.2% with a maximum of 2.7% Dl/ l occurring near the crack tip corresponding to crack initiation at 102–103 loading cycles [29]. Power cycling data shows that under actual test conditions, failure at a junction temperature range of 75 K usually occurs above 104 numbers of loading cycles [25]. Non-uniform strain distribution, crack initiation and failure modes of micro-joints due to power cycling tests resemble closely the deformation and fracture mechanisms involved in the mechanical shear fatigue procedure and are well described by FE analysis. Micro-sections of specimens subjected to mechanical shear fatigue reveal deformation gradients from the interface toward the centre of the wire. The development of fatigue cracks is concentrated in a rather narrow area near the interface while the bulk of wire either shows lower degrees of cyclic deformation or remains un-deformed (Fig. 2, 8a and 8b). Analogous to the crack initiation site in lift-off failure due to power cycling [12,18] and FE analysis [29,30], the fatal crack is always induced at the termination of bonded wire growing inwards the bonded area.
4. Summary and conclusion In this paper a novel high frequency mechanical shear testing method for fatigue life determination of micro-joints is introduced which operates at 20 kHz and can be used as an alternative to thermal and power cycling tests. The strain induced in the interfaces of the different materials due to thermal mismatch, is one of the main failure mechanism in microelectronic devices The basic idea of our method is to reduce the testing time by inducing an equivalent thermal strain in the interface of the specimens by mechanical means. The test set-up is based on the principles of forced cyclic vibration of coupled materials by using an ultrasonic resonance fatigue testing device. The developed system, is capable to generate mechanical cyclic shear strain in the interface of various micro-joints and bonded components to obtain fatigue life curves in very short testing times. Further, wire bonded Al micro-joints were used to evaluate the applicability of the proposed testing method for life time prediction of interconnects. The shear strength induced in the microjoints was calculated by using a Laser doppler vibrometry method and fatigue life curves (S–N curves) were plotted. The bimetallic law (D e = a DT) was used to approximate the equivalent temperature swing DT to the shear strain the micro-joints and model the life time by using the Coffin–Manson relationship. This approach enabled us to compare the results with the power cycling data of similar joints. It could be shown that the life time predicted by mechanical shear fatigue and power cycling indicate a very good correlation. The failure analysis showed that the predominant failure mechanism of power cycling tests, bond wire lift-off due to Al wire degradation is reproduced in the proposed mechanical test. The scatter in the fatigue life data could be related to the presence of defects and non-bonded areas. It can be concluded that the proposed method can be used not only to predict life time of micro-joints in significantly short periods of time, but also allows fast examination of failure surface. This provides useful information about the influence of bonding/interconnection technology and interaction of constituent materials on the reliability of the joint. The proposed method can be applied to various kinds of micro-joints such as wire bonds and solder
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joints as an alternative to conventional thermal cycling tests. It can be used as a reliable rapid product qualification and screening tool in the interconnect technology, reducing the time and associated costs during the design and production process. Acknowledgements The authors wish to gratefully thank FFG – Bridge program for financial support (Project No. 811022/9455 – SCK/SAI), Dr. G. Mitic, Dr. P. Zimprich, Prof. V. Groeger and Prof. R. Stickler for valuable discussions and reviewing of the manuscript. We are grateful to Mr. G. Lefranc for initiating this investigation and providing Fig. 2. References [1] Martin P. Electronic failure analysis handbook. USA: McGraw-Hill; 1999. [2] Dasgupta A, Pecht M. Material failure mechanisms and damage models. IEEE Trans Reliab 1991;40(5):531–6. [3] Lall P. Tutorial: temperature as an input to microelectronics – reliability models. IEEE Trans Reliab 1996;45(1):3–9. [4] EIA/JESD22-B116 July 1998. Wire bond shear test method. [5] IEC 60749-25 ED. 1.0 B: 2003. Semiconductor devices-mechanical and climatic test methods-Parts 25: Temperature cycling. [6] IEC 60749-34 Ed. 1.0 B: 2005. Semiconductor devices-mechanical and climatic test methods-Parts 34: Power cycling. [7] Uegai Y, Tani S, Inoue A, Yoshioka S, Tamura K. A method of fatigue life prediction for surface mount solder joints of electronic devices by mechanical fatigue test. Proc 2nd ASME Int Elect Packag Conf 1993;1:493–8. [8] Ramminger S, Seliger N, Wachutka G. Reliability model for Al wire bonds subjected to heel crack failures. Microelect Reliab 2000;40:1521–5. [9] Zhao Y, Basaran C, Cartwright A, Dishongh T. Thermomechanical behaviour of micron scale solder joints under dynamic loads. Mech Mater 2000;32(3):161–73. [10] Xia Liu K, Valmiki A, Sooklal A, Melod B, Verges C, Michael AL. Experimental study and life prediction on high cycle vibration fatigue in BGA packages. Microelect Reliab 2006;46:1128–38. [11] Plumbridge WJ, Matela RJ, Westwater A. Structural integrity and reliability in electronics. London: Kluwer Academic Publishers; 2003.
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