Characterisation of heat spreader materials for pulsedIGBT operation

D.J. Lim and S.H. Pulko

Abstract: Insulated gate bipolar transistors (IGBTs) have a very high output power and generatecorrespondingly large amounts of heat. If not dissipated efficiently, this heat will destroy the IC(integrated circuit). Furthermore, since the input to the IGBT is often in the form of a pulsedwave, the rapid repeated heating and cooling of the chip and the surrounding packaging causephysical stresses, which in turn eventually lead to delamination and breakdown. Reducing themagnitude of thermal excursion in pulsed mode operations reduces the amount of stress causedby expansion and contraction, thereby reducing delamination and maintaining component effi-ciency for a longer period of time. It is therefore important to maintain a low rate of thermalexpansion, or have a slow enough change in temperature for the physical stresses not to be dama-ging. This is normally done with heat sink assemblies, which form an integral part of IGBTdesign. This study investigates, via simulations using the transmission line matrix method, thethermal responses of some of the popular heat spreader materials. Material combinationswithin the layered structure of the heat sink assembly will give different thermal responses,and thus an analysis of operational behaviour of these components, with attention given to theinput frequency as well as duty cycle, would provide a better guide to designing more suitableand efficient packaging assemblies and heat sinks.

1 Introduction

Over the last two decades, the insulated gate bipolar transis-tor (IGBT) has become one of the most popular power semi-conductor devices for motor control, inverter and laserwelding applications [1–6]. This has been the result ofthe IGBT’s unique ability to handle similar high voltageand current levels to a bipolar transistor while retainingthe ease of operation, via voltage-control, which is normallyassociated with a MOS-field-effect-rransistor (MOSFET).Although the IGBT typically has a low switching frequency(�50 kHz) compared to a MOSFET, the IGBT’s othercharacteristics make it apt for use in high power and lowfrequency applications. It is this capacity to handle largevoltages (.1000 V), and the ability to function even witha high junction temperature (slightly above 1008C),coupled with ease-of-use akin to that of a MOSFET,which have made the IGBT so popular with designers oftraction drives, welding laser assemblies and other high-voltage switching applications.

IGBTs generally have a very high output power, oftenin excess of 5 kW, and generate correspondingly largeamounts of heat. This can result in temperatures far inexcess of 1008C [7], which cause device damage. If notdissipated efficiently, this heat will destroy the IC.Furthermore, since the input of the IGBT is often in theform of a pulsed wave, the rapid repeated heating andcooling of the chip and the surrounding packaging causephysical stresses, which can in turn eventually lead to

# The Institution of Engineering and Technology 2007

doi:10.1049/iet-cds:20050227

Paper first received 9th August 2005 and in revised form 4th December 2006

The authors are with the Department of Engineering, University of Hull,Cottingham Road, Kingston upon Hull, HU6 7RX, UK

E-mail: s.h.pulko@hull.ac.uk

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breakdown. In some cases the solder which holds theconnections in place either melts, cracks or lifts from thebase-plate [8]. This process, which is known as delamina-tion, is a major concern of IGBT packaging manufacturers.Delamination also occurs between the IGBT and the coolingstructure, which is normally in the form of multiple layers ofthermally conductive material bonded to the chip as shownin Fig. 1 [8, 9]. When delamination occurs, the temperaturein the chip increases further, since the chip now has reducedthermal contact with the heat sink assembly. Ultimately, thewhole physical structure of the silicon may break down asthe increased temperature changes the electrical and phys-ical properties of the materials, causing thermal runaway.The end result of this is usually a catastrophic failure ofthe IGBT. There are, of course, many other factors thatcause component failure in an IGBT [10], but thermalexpansion stresses and the magnitude of the thermalpower involved in typical IGBT operations to a greatextent determine the lifetime of the component. It is forthis reason that the coefficient of thermal expansion(CTE) is used in many cases as a parameter, which dictatesthe choice of material in an IGBT structure. If the bondedand bonding materials around the chip have very similarCTE values, the tendency to delaminate can be reduced,thereby increasing component lifespan. However, evenwith similar CTE values, there will be a time delay in thesystem, as heat needs time to propagate through the sur-rounding areas. This means that with pulsed input, thechip itself could be cooling down and contracting as thearea surrounding it is still heating up and expanding fromthe pulse of heat that was emitted previously.It is therefore critically important to remove heat from the

area surrounding the active region quickly in order toachieve a satisfactory component lifetime. Quickly remov-ing the heat from the source ensures that, even if slight dela-mination occurs, thermal runaway can be significantly

IET Circuits Devices Syst., 2007, 1, (2), pp. 126–136

delayed [8]. Furthermore, reducing the magnitude ofthermal excursion in pulsed mode operations reduces theamount of stress caused by the rapid expansion and contrac-tion in the first place, thereby reducing delamination, andmaintaining component efficiency for a longer period oftime. It is, therefore, important to maintain a low rate ofthermal expansion, or at least have a slow enough rise intemperature for the physical stresses not to be damaging.In other words, the smoother or slower the thermal transientto steady-state operation, whether that be with a pulsedinput or otherwise, the less stress there will be on thebonded and bonding elements of the IGBT packaging.Although work has been done on coupled electrothermal

simulation of devices such as IGBTs [11], and sophisticatedapproached to implementation have been developed [12, 13],we are unaware of any systematic study of the influenceof material choice on the development and behaviourof thermal field in these devices. Here, we describe asimulation-based investigation of the thermal effects ofusing different spreader materials in IGBT structures underpulsed operation. Three heat spreader materials are con-sidered: copper (Cu), copper–molybdenum (Cu-Mo) andaluminium–silicon–carbide (AlSiC). Simulations suggestthat the spreader material needs to be selected with careand that the thermal conductivity, the specific heat and thedensity are all relevant to the choice.

2 IGBT heat sink assembly structure

As is suggested in Fig. 1, the heat sink assembly of an IGBTpackage is a layered structure. The IGBT itself only com-prises a small fraction of the physical dimensions of an

Fig. 1 Generic IGBT-layered heat sink assembly structure

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IGBT assembly. There is a top plate, which covers andprotects the chips, but the main heat dissipation assemblyis under the chips. Direct bond copper (DBC) or othersimilar methods are used to braze the layers together.Most IGBT devices have a ceramic substrate that is sand-wiched between two ‘heat spreader’ layers, which areusually around 300 mm thick [10]. These layers aretypically made of copper or other similar materials. Thissandwich is then brazed onto a base-plate, which is inturn secured to the heat sink structure, which has a highsurface area for heat transfer to ambient. The thermalcontact between the heat sink and the base-plate isensured by thermal grease, paste or other heat sink com-pound. The substrate layer ensures that the base-plate iselectrically isolated from the active components, whichare individually wire bonded for top-connect operation.

The layered structure of the assembly makes the tempera-ture variation during IGBT operation quite complex. Thechip must be kept within operational temperatures for theIGBT to continue functioning. Additionally, minimisingstress in the solder layer due to repeated and drasticthermal expansion and contraction should be consideredin the design of an IGBT assembly, in order to prevent orreduce the likelihood of delamination. The rate of heat dis-sipation in the chip and the area surrounding the chip is ofprime concern, and therefore a thermal dissipation systemthat removes heat quickly and efficiently from the chip,but has a slightly higher average temperature can be prefer-able to a system that has a lower overall temperature but ishotter at the area surrounding the chip. In order to explorethe possibilities and practicalities of designing such asystem, the thermal simulations in this study focus on thearea of the heat sink assembly closest to the IGBT chip,namely, the chip itself and the top heat spreader layer.

A range of materials commonly used in IGBT construc-tion is shown in Table 1. These materials are normallychosen based on either their thermal conductivity (Kt) ortheir thermal diffusivity, which is related to thermal conduc-tivity via

D ¼Kt

rSpð1Þ

where Sp and r are the specific heat capacity and density ofthe material, respectively. Table 1 also includes values forthe production of specific heat capacity and density (rSp)as representing the heat necessary to raise the temperatureof a unit volume of material. Consideration also tends tobe given to the CTE, even though there is no explicit rule

Table 1: Material properties of IGBT heat sink assembly structure

Material Density (r),

kg/m3

Specific heat (Sp),

J/kg.K

Thermal conductivity (Kt),

W/m.K

rSp Diffusivity

(Kt/rSp)

Substrate

Diamond 2K 3510 620 2000 2.18 � 106 9.19 � 1024

Spreader

Cu 8960 276 393 2.47 � 106 1.59 � 1024

CuMo 9985 678 197 6.77 � 106 2.91 � 1025

AlSiC 2980 722 180 2.15 � 106 8.37 � 1025

Other materials

Silicon 2320 700 148 1.62 � 106 9.11 � 1025

Solder 7400 160 40 1.18 � 106 3.38 � 1025

Top plate 10 220 255 138 2.61 � 106 5.30 � 1025

Base plate 2980 722 180 2.15 � 106 8.37 � 1025

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that determines that choice [6, 14, 15]. This paper focuseson the choice of heat spreader material for pulsed operationof IGBTs.

Copper (Cu) and copper–molybdenum (CuMo) heatspreaders are used for this study, as they not only representcommon material types [6], but also provide an interestingcontrast of properties. Cu, a standard heat spreader materialin contemporary IGBT heat sink assemblies, has a highthermal conductivity, a low specific heat capacity (Sp) anda high density (r) compared to other candidate materials,such as aluminium–silicon–carbide (AlSiC). On the otherhand, Cu–Mo has about half the Kt of Cu and aboutdouble the specific heat capacity. Both Cu and CuMohave similar densities. This in turn means a much higherthermal diffusivity for Cu than for CuMo.

The substrate is typically [6] a metal matrix composite(MMC) material. A popular MMC is aluminium nitride oralumina (AlN), which has a low thermal conductivity anddensity but a very high specific heat capacity. Berylliumoxide (BeO) is another popular choice. In this study,however, the substrate used is chemical vapour deposited(CVD) diamond, which has a much higher thermal conduc-tivity (Kt ¼ 2000 W/mK) compared to AlN, while retaininga similar specific heat capacity and density. This will havethe effect of passing the heat through the substrate withminimal resistance, thereby allowing clearer examinationof thermal transients within the chip, solder and spreaderlayers in the IGBT assembly. The thermal properties ofthe various materials used in this investigation are presentedin Table 1 [15].

3 Model structure and the transmission linematrix method

The simulations are based on the transmission line matrix(TLM) method [16]. TLM is widely used to model electro-magnetic problems but also finds application in thermal dif-fusion problems [17, 18], where it has been used innumerous industries [19], including glass lens pressing[20], thermal management of electronic networks [21] anddevices and ceramic drying and firing [22].

TLM is an iterative technique in which the modellingvolume is divided into spacial elements and the modellingperiod into iteration timesteps [23]. In a TLM model forthermal diffusion current and electrical potential in theTLM network are analogous to heat flow and temperature;thermal conductivity is represented by electrical resistorswhich are clustered around calculation nodes, and elementalheat capacity (Spr) is modelled by electrical capacitance.Fig. 2 shows a lumped RC network such as would representthe thermal behaviour of a cubic element of material.

Fig. 2 Lumped RC network representing the thermal propertiesof a cube of material

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Considering the flow of heat between opposite sides ofthe cube side of Dl, the value of each resistance is given by

R ¼Dl

2KTðDlÞ2¼

1

2KTðDlÞð2Þ

When the lumped network is transformed into a TLMnetwork, the resistors remain clustered around the nodeand retain their value; capacitance is represented by trans-mission line impedance [23]. Two types of transmissionlines can be used, link transmission lines and stub trans-mission lines. As illustrated, for one dimension, in Fig. 3,link lines connect each node to its neighbours and carrytemperature pulses between nodes during the timestep, Dt.If elements are cubic and of side, Dl, then each link linehas length Dl. The corresponding stub transmission lines,on the other hand, have length 0.5Dl. At each node, apulse scattered from the node into the stub line travels tothe open circuit termination, is reflected, and then traversesthe same stub line again to be incident once more on thesame node at the next iteration. The total capacitanceassociated with the element is

C ¼ SprðDlÞ3

ð3Þ

If we choose to allocate a capacitance C� to the stub line[24], then the remainder of the nodal capacitance is associ-ated with the link lines. Since, for a one-dimensional node,each element contains one complete link line the impedanceof that line is

Z1�d ¼Dt

ðC � C�Þð4Þ

The extension of this argument to 2- and 3-dimensionalnodes is straightforward.The impedance of the stub line is associated only with C

�

,the capacitance assigned to the stub, so that

Zs ¼Dt

2C�ð5Þ

the factor of 2 occurring because the pulse traverses the stubtwice during each iteration.In operation of a TLM routine [23, 25], pulses are

incident on each node from all branches, including thestub, at every iteration. They give rise to a nodal potential(temperature) according to

F ¼X 2V i

Rþ Zþ2V i

s

Zs

" #1

Y

where

Y ¼Xl

1

Rþ Zþ

1

Zsð6Þ

In (7), F represents the nodal temperature and V themagnitude of pulses, the superscript i denotes an incidentpulse, the subscript l refers to the link line port numbersas illustrated in Fig. 3, and the subscript s refers to thestub. Reflected pulses on the link lines are then calculated

Fig. 3 One-dimensional TLM network for modelling diffusion

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according to

V rl ¼

FZ

Rþ Zþ V i

l

R� Z

Rþ Zð7Þ

where the superscript r refers to reflected quantities. Thestub line is traditionally resistance free so that R ¼ 0 and(7) becomes

Vrs ¼ F� V

is ð8Þ

Pulses on link lines then travel along those lines, to arriveat the relevant ports of neighbouring nodes a timestep later.The pulse on the stub line travels to the end of the stub trans-mission line and is reflected to be incident on the same nodea timestep later.TLM models solve equations of the form

r2V ¼ 4RdCd

dV

dtþ 2LdCd

d2V

dt2ð9Þ

where V is the voltage, Rd represents the distributed resist-ance, Cd the distributed capacitance and Ld the distributedinductance. As can be seen from the description of the scat-tering process, the routine is explicit, requiring informationonly from neighbouring nodes at any iteration, so thatmemory requirements are modest. It is also unconditionallystable, so that the results of a simulation will not tendtowards infinity, regardless of the time-step used, but theerror term associated with the second term on the right-handside of (9) will increase and the results become wavelike[26]. Therefore, the accuracy of the transient simulationcan be varied according to the requirements, making highprecision, short-time step simulations for situations withfast transients, as well as low precision, long-time stepsimulations for slow transient or steady-state situationsequally viable [23, 25]. This flexibility is of particularbenefit to the simulations in this study, as both transientand steady-state simulations were required. As is the casefor the Du Fort–Frankel method, it is necessary to ensurethat a TLM model is consistent with the diffusion equationby arranging for Dl and Dt to approach zero appropriately asthe mesh is refined; the condition for consistency of a TLMmodel is that (Dt)2 tend to zero as (Dl) tends to zero [27].All data for this study were generated using in-house

developed, well-validated software implementing theTLM technique. Stub transmission lines were used torepresent variations in elemental capacitance from onematerial to another, and care was taken to ensure that thediscretisation regime and network parameterisation wereappropriate [24].Two simulation models were used in this series of

studies. The first, more complex model mimics the structureof a typical IGBT assembly [15], but uses a heat transfercoefficient to represent the high surface area fin structure.Many common IGBT modules have six subassemblies,arranged in a 3�2 matrix structure. In the interests of com-putational efficiency, the model has been simplified to onlyrepresent a quarter of one such subassembly, which is6.8 cm � 6.2 cm in area with a thickness of 0.64 cm, asillustrated in Fig. 4. All boundaries of this model, whichrepresent the model’s contact with the ambient, are main-tained at 08C. The two faces that correspond to the sectionslinked to the rest of the assembly are represented by reflec-tive boundaries. The active region is subjected to 1.7 kWpulses of various durations, and the temperature rise isplotted at various points along the central axis through theassembly.

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The second model is a simplified version of the first,with thermally reflective boundaries on all four sides, and08C ambient conditions on the upper and lower surfaces.This second model represents a small section of theIGBT from the middle of the assembly. This ‘apple core’model (thus named for its similarity to an extracted applecore) provides a vastly faster running, albeit one-dimensional, model which allows structural or materialchanges to be made and simulated within a reasonabletimeframe. The ‘apple core’ model represents a cross-section measuring 0.6 cm � 0.6 cm in area with a thicknessof 0.64 cm. Both models represent an IGBT assembly witha CVD diamond (Kt ¼ 2000 W m21 K21) substrate. Theheat spreaders studied are copper and a copper–molyb-denum alloy blend. When comparing the first 3D modeland the second ‘apple core’ model, it is found that bothmodels have very similar trends, with the ‘apple core’model maintaining a slightly higher temperature overall,as shown in Fig. 5. The discrepancy in temperature is con-sistent with the fact that the 3D model has upper, lower andtwo other faces where heat transfer to ambient occurs, withonly two reflective boundaries, whereas the ‘apple core’model only has heat transfer to the ambient from the topand bottom faces. The discrepancy between the models isnegligible for the purpose of this investigation, and thusthe 1D model is used to generate all the results, unlessotherwise stated.

4 Simulation results

Figs. 6–8 show the temperature profile through the centre ofan IGBT assembly at 1 � 1023, 1 � 1022 and 1 � 1021 sfor models with Cu and CuMo heat spreaders and continu-ous heat generation in the 3D models. It is evident that thelayer with the largest thermal gradient per unit thickness inthe IGBT heat sink assembly is the solder layer that linksthe IGBT chip to the upper heat spreader. Unfortunately,there are many thermal, chemical and electrical restrictions,which prevent the type and composition of the solder usedbeing modified at will [28]. However, the region with thesecond highest heat gradient, the upper heat spreader,does not have many of these restrictions. From Figs. 6–8,this region has the largest temperature drop per unit thick-ness in the heat sink assembly after the solder layer, dis-counting the IGBT chip itself. Comparing the layers at1 � 1021 s, the difference in rate of temperature drop perunit thickness in the upper spreader layer for both the Cuand CuMo models is about twice that in the lower layer,and more than ten times that in the substrate. Thus, thislayer is the object of further scrutiny and the subject ofthis study.

Fig. 4 Simulated section of 3D model (top view)

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Fig. 5 Comparison of 3D and 1D simulation models

Transients show chip temperatures

Fig. 9 compares the temperature in the centre of the chipfor Cu, CuMo and AlSiC heat spreaders for uniform heatgeneration throughout the chip over 1 � 1021 s. As isseen from Fig. 9a, AlSiC is less thermally favourable thanCu or CuMo. As far as the Cu and CuMo heat spreadersare concerned, it is apparent that the effect the heat spreadermaterial has on the thermal response of the IGBT chip is notas straightforward as it might initially seem. This is evidentin the thermal transients presented in Figs. 9a and 10a,where the models with Cu and CuMo heat spreaders havepoints where the thermal transients ‘cross-over’. This‘cross-over’ point occurs at different points in time in thedifferent layers of the assembly, as Figs. 9a and 10a. Themodel with the Cu heat spreader has a slightly higher temp-erature very early in the transient, both in the chip and inthe solder layer between the chip and the upper spreaderlayer. This occurs even though Cu has a higher Kt value

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than Cu–Mo. However, at �2.0 � 1023 s, the thermaltransients converge and crossover, the Cu–Mo heat sprea-der now showing slightly higher temperature in the chip,as shown in the detailed transient in Fig. 9a. At�2.0 � 1022 s, the temperature trends reverse again, withthe model with the Cu spreader now showing the highertemperature. Much further into the transient, and beyondthe scope of Figs. 9a and 10a, the thermal profiles crossoveryet again and continue into steady-state, with the CuMospreader model having the higher steady-state temperature,as would be expected of a material with a lower Kt than Cu.Figs. 11–13 show the transient temperature of three con-

secutive pulses, each having a 50% duty cycle, for modelswith Cu and CuMo heat spreaders. From the simulationresults, which have been tabulated in Table 2, it is evidentthat the rate of temperature rise is slower with each consecu-tive pulse, whereas the rate of temperature drop (i.e. the rate

Fig. 6 Temperature profile through IGBT assembly (1 � 1023 s)

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Fig. 7 Temperature profile through IGBT assembly (1 � 1022 s)

at which the temperature in the chip falls after the pulseends) is greater. Comparing the relative temperature risefor the duration of the pulses at the chip at 0.1 and 0.2 s(the end of the first and the second pulse), the temperaturerise caused by the first pulse is about 7.25% higher thanthe temperature rise caused by the second pulse in bothmodels (Cu and CuMo heat spreaders), whereas the relativetemperature rise caused by the second pulse is about 8.75%higher than that caused by the third pulse at 0.3 s. By con-trast, the peak of the second pulse is �8.65% lower than thefirst, whereas the peak of the third pulse is only 0.42% lowerthan the second pulse. The faster drop in temperature, com-bined with the slower pulse rises, has a cumulative effect ofcausing the temperature in the assembly with the CuMo heatspreader to be consistently lower than the one with the Cu

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heat spreader with each consecutive pulse, as is evidencedin Fig. 14.

5 Discussion

The main difference between Cu and CuMo in terms ofthermal properties is that Cu has a significantly higher Kt

and a higher thermal diffusivity than CuMo. On the otherhand, for CuMo the product of specific heat capacity anddensity is higher than is the case for Cu, so that a unitvolume must absorb more heat before its own temperaturerises. Thus, although the CuMo heat spreader cannotperform as well as Cu in removing the heat from theentire region in the steady-state, the situation is morecomplex in the transient and the results presented in

Fig. 8 Temperature profile through IGBT assembly (1 � 1021 s)

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Fig. 9 Comparison of chip temperatures for IGBT assemblies with Co, CuMo and AlSiC heat spreaders

a Comparison for times up to 5 � 1022 sb Details of the thermal crossover

Figs. 9 and 10 suggest that at some stages the lower transi-ent chip and solder temperatures are associated with the useof a CuMo spreader.

Temperature differences are larger in the case of theAlSiC spreader, which is associated with higher chip andsolder temperatures throughout the early transient asshown in Figs. 9 and 10. Although the thermal diffusivityof AlSiC falls between that of Cu and that of CuMo, itsthermal conductivity is less than a half of that of Cu andits specific heat capacity–density product is only a thirdthat of CuMo. Although in steady-state chip temperatureswere found to be lowest with a Cu spreader and highestwith an AlSiC spreader with CuMo yielding intermediatevalues, it appears that much more detailed considerationis needed to ascertain which spreader will yield lowesttemperatures in the transient.

‘Crossover phenomena’, shown in Figs. 9–14, indicatethat the choice of heat spreader material should be madewith consideration to the duty cycle of the IGBT in ques-tion. From Figs. 9a and b, it is clear that the thermalresponse can be separated into three distinct ranges. The

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same is even more evident in the solder layer between thechip and the upper spreader (Fig. 10), where the crossoveroccurs slightly earlier (approximately at 1.75 � 1023 sinto the transient) resulting in a marginally larger discre-pancy between the models for Cu and CuMo heat spreaders.It is evident that, within the transient period, at operationalfrequency between 50 and 500 Hz a Cu heat spreader yieldslower temperatures, at least for duty cycles around 50%, atfrequencies outside this range a CuMo heat spreader yieldslower temperatures.The materials used in the heat sink assembly typically

have different magnitudes of thermal expansion and con-traction for a given temperature change, which is numeri-cally expressed as the CTE. Since the assembly is alayered structure, materials with different CTE can causethe layers of the assembly to warp when subjected torapid thermal excitation, which may then lead to delamina-tion. As this occurs, the efficiency of thermal dissipationfrom the active region deteriorates rapidly, leading to a criti-cal failure as the chip overheats. Therefore, componentdesign should not only take into account the thermal and

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Fig. 10 Comparison of upper spreader solder temperatures for IGBT assemblies with Co, CuMo and AlSiC heat spreaders

a Comparison for times up to 5 � 1022 sb Details of the thermal crossover

Fig. 11 Thermal transient comparison of first pulse in an IGBT chip, with Cu and CuMo spreaders

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Fig. 13 Thermal transient comparison of third pulse in an IGBT chip, with the start point of the pulses normalised to 0 for comparison

Cu and CuMo spreaders

Fig. 12 Thermal transient comparison of second pulse in an IGBT chip, with the start point of the pulses normalised to 0 for comparison

Cu and CuMo spreaders

physical problems inherent in the chip-to-heat sink assem-bly connection, but also of the interaction issues betweenthe layers of the heat sink assembly itself. Cu and AlSiChave much higher CTE values (17.2 and 7.0, respectively)making them less favourable compared with CuMo.

It was also observed in Fig. 14 that the cooling rates werehigher with each consecutive pulse. This implies that, as the

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number of pulses increases, the cooling rates of the heat sinkassembly will play a more important role in ensuring anacceptable overall temperature compared to the mainten-ance of a lower peak temperature, as relatively more heatwill be dissipated with each successive pulse.Interestingly, whereas the model with the Cu heat spreaderhas larger differences in the relative pulse peak

Table 2: Temperature rise and fall percentages for three pulses in models with Cu and CuMo heat spreaders, diamond2k substrate

Cu Pulse peak (8C) Pulse trough (8C) % Drop (from peak) % Rise (pulse 1–2) % Rise (pulse 2–3)

1st Pulse 33.46 18.49 44.75 — —

2nd Pulse 49.20 33.35 52.08 9.05 —

3rd Pulse 63.67 47.45 53.50 — 0.38

CuMo

1st Pulse 31.93 16.25 49.11 — —

2nd Pulse 45.83 29.34 56.30 8.28 —

3rd Pulse 58.50 41.63 57.85 — 0.45

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Fig. 14 Graph depicting the thermal transient of three consecutive pulses for a heatsink with a Cu and CuMo spreader

temperatures, that is each consecutive pulse has a relativelylower temperature compared to the CuMo spreader model(�0.76%), the CuMo model has a significantly larger temp-erature drop compared to the Cu spreader model (4.28%).This means that the model with the CuMo heat spreaderalways starts each successive pulse at a lower temperaturethan the model with a Cu heat spreader. As Fig. 14shows, this phenomenon causes the model with the CuMoheat spreader to have an increasingly lower temperaturecompared to that with the Cu heat spreader. This is contraryto the traditional expectations based on the material proper-ties, where Cu, which has a higher Kt value, would beexpected to have a lower temperature. Although this isindeed the case once the chip has reached steady-state temp-eratures, it is not true for the transient, or pulsed transientoperation. In pulsed transient operation, the temperatureswithin the chip are always changing. Even when the assem-bly has reached a general steady-state, the chip is stillthermally transient. Therefore, the effect of the RhoSpproduct still has a significant effect on the thermal profileswithin the assembly.AlSiC was also considered for this study and was found

to have consistently and significantly higher temperaturesat both transient and steady-state compared to Cu andCuMo. The results of the simulations show that, althoughselecting heat spreader material can be based on the Kt

value in certain situations, this value alone cannot be usedas a definitive measure of a heat sink material’s suitabilityor efficiency. Furthermore, diffusivity alone is not awholly valid parameter by which to select materials foruse in transient applications. Since material combinationswithin the layered structure will give varied thermalresponses, an analysis of operational behaviour of thesecomponents, with attention given to the input frequencyas well as likely duty cycle would provide a guide to design-ing better and more suitable packaging assemblies andheat sinks.

6 Conclusion

Here, it has been shown that the temperatures developed, inthe chip and elsewhere, within IGBT structures are signifi-cantly dependent on the choice of heat spreader material.Furthermore, the heat flow situation is complex in pulsedoperation, and neither a high thermal conductivity nor a

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high thermal diffusivity can identify one material as yield-ing lower temperatures than another irrespective of operat-ing frequency and duty cycle.

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IET Circuits Devices Syst., Vol. 1, No. 2, April 2007