In-situ characterization of interface delamination by a new miniature ...

Int J Fract (2009) 158:183–195DOI 10.1007/s10704-009-9356-1

ORIGINAL PAPER

In-situ characterization of interface delaminationby a new miniature mixed mode bending setup

M. Kolluri · M. H. L. Thissen ·J. P. M. Hoefnagels · J. A. W. van Dommelen ·M. G. D. Geers

Received: 14 November 2008 / Accepted: 20 April 2009 / Published online: 4 June 2009© The Author(s) 2009. This article is published with open access at Springerlink.com

Abstract A new miniature mixed mode bending(MMMB) setup for in-situ characterization of inter-face delamination in miniature multi-layer structureswas designed and realized. This setup consists of anovel test configuration to accomplish the full rangeof mode mixities and was specially designed with suf-ficiently small dimensions to fit in the chamber of ascanning electron microscope (SEM) or under an opti-cal microscope for detailed real-time fracture analysisduring delamination. Special care was taken to min-imize the effects of friction, the influence of gravity,and non-linearities due to the geometry of the setup.The performance of the setup was assessed using spe-cially-designed test samples supported by finite ele-ment analyses. Delamination experiments conductedon homogeneous bilayer samples in mode I and mixedmode loading were visualized with a scanning electronmicroscope and showed the formation of small microcracks ahead of the crack tip followed by crack bridgingand a full crack, thereby demonstrating the advantagesof in-situ testing to reveal the microscopic delaminationmechanism.

M. KolluriMaterials Innovation Institute (M2i), Mekelweg 2,P.O. Box 5008, 2600 GA Delft, The Netherlands

M. Kolluri · M. H. L. Thissen · J. P. M. Hoefnagels (B) ·J. A. W. van Dommelen · M. G. D. GeersDepartment of Mechanical Engineering, EindhovenUniversity of Technology, W.H. 4.13, P.O. Box 513,5600 MB Eindhoven, The Netherlandse-mail: [email protected]

Keywords Delamination · Interfacemechanics · Mode mixity · Miniaturization ·In-situ characterization

1 Introduction

The demand for the continuous miniaturization andmulti-functionality of electronic devices led the elec-tronic industry towards new packaging technologieslike “System in Package” (SiP). SiP technology primar-ily involves assembly of several components (activecomponents, passive components and MEMS) into asingle package to achieve different functionalities. Con-sequently, SIP-microsystems contain several interfacesformed between stacked, multiple, thin layers, manu-factured using different materials and processes. Inter-faces are often recognized as critical regions for thereliability of these products because the presence ofhigh stresses at these interfaces (e.g. due to a mismatchin coefficient of thermal expansion (CTE) and Pois-son’s ratio) and in-plane shear stresses often lead todelamination (Hutchinson and Suo 1992; Evans andHutchinson 1995; Dauskardt et al. 1998; Vlassak et al.2005). Much research is ongoing to improve the inter-face behavior by adopting suitable coatings or clean-ing techniques at the interface during manufacturingof these components (Srikanth et al. 2006). However,detailed quantitative characterization of the interfacesin these systems is required in order to improve design

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rules for manufacturing and improve reliability duringservice of these products.

A number of experimental techniques have beendeveloped and reported in the literature (e.g. Evans et al.1990; Dreier et al. 1992; Gupta and Pronin 1995; Vo-linsky et al. 2003; Li and Siegmund 2004; Banks-Sillset al. 2005; Wang et al. 2006; Ocaña et al. 2006; Liuet al. 2008) to measure interfacial properties, most ofwhich use a fracture mechanics approach because ofits proven versatility. A fracture mechanics approachrequires an experimental setup capable of triggeringdelamination and uses the resulting force-displacementdata to obtain the interface fracture toughness or crit-ical energy release rate (CERR) as an important inter-face parameter to characterize the interface. It hasbeen shown by many authors (e.g. Hu and Evans 1989;Hutchinson and Suo 1992; Banks-Sills et al. 2000;Kuhl and Qu 2000; Shi et al. 2006; Tang et al. 2007)that the CERR for crack propagation varies signifi-cantly with mode mixity. Furthermore, the presenceof dissimilar materials forming the interface, whichis common in microsystems, increases the complex-ity of the CERR measurement. This indicates that theexperimental technique to characterize an interface sys-tem should be able to measure the CERR over theentire range of mode mixities, in order to identify validinput parameters in a design limit criterion. Techniquesreported in the literature to measure the CERR includethe well-known double cantilever beam (DCB) test forpure mode-I loading (ASTM D 5528-01, 2001) andend notch flexure (ENF) test for applying pure mode-IIloading (Carlsson et al. 1986), as well as mixed modebending (MMB) setups (e.g. Reeder and Crews 1990;Soboyejo et al. 1999; Merrill and Ho 2004; Blanco et al.2006; Thijsse et al. 2008), which cover a range of modemixities, see Fig. 1. Of all setups, only the MMB setupsyield the CERR over a full range of mode mixities witha single test configuration.

The prime difficulty in many of the existing delam-ination experiments is the identification of the cracktip location in order to trace the crack length, which isneeded to calculate the fracture toughness. To

circumvent this problem, critical loads deduced fromexperiments, where the delamination starts for a givenpre-crack length, were used to calculate the CERR withthe aid of analytical formulae (Davidson and Sundarar-aman 1996; Davies et al. 1998; Soboyejo et al. 1999).Stiffness lines generated with finite element (FE) mod-els of known crack length specimens were used to eval-uate the CERR by (Thijsse et al. 2008). Some authors(e.g. Vinciquerra and Davidson 2004; Gunderson et al.2007) reported that a crack length measurement with asimple optical magnification lens system may be erro-neous and also highlighted the need for techniqueswhich can measure crack lengths more accurately. Fur-thermore, detailed in-situ visualization of the delami-nation process is essential to accurately pin-point thecrack tip location, to measure quantitative delamina-tion characteristics such as the crack opening profileand the process zone size, and to obtain more insightin the fracture events along the interface.

In-situ characterization in a microscope, however,puts serious constraints on the overall size of the delam-ination setup. Evaluation of existing MMB setups elu-cidates the difficulties to use them for in-situ testing.In case of the setup of Reeder and Crews (1990), shownin Fig. 2a, restricted dimensions of the design space pre-vent the lever from having sufficient length to cover thecomplete range of mode mixities. Furthermore, this testis difficult to perform in the horizontal plane (i.e. keep-ing the direction of load application in the horizontalplane), which is necessary to enable the use of stan-dard microscopes to trace the crack tip movement dur-ing in-situ delamination testing, (e.g. the viewing axisof a scanning electron microscope (SEM) is always inthe vertical direction). Merrill and Ho’s setup (Merrilland Ho 2004), shown in Fig. 2b, was also constructedsuch that the loading direction is vertical, leading tosimilar limitations. Furthermore, gravity acts on bothof the above mentioned setups, resulting in additionalundesired forces on the sample. In the setup of Thijsseet al. (2008), a counter balance was added to the loadingconfiguration of Reeder and Crews (1990) to minimizethe influence of gravity. However, a persistent conse-

Fig. 1 Sketch of differentloading configurations forinterface delamination.a DCB; b ENF; c Mixedmode bending

P

P (a)

P

(b) (c)

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In-situ characterization of interface delamination 185

Fig. 2 Two differentexisting concepts for mixedmode delamination testing.a Reeder and Crews (1990);b Merrill and Ho (2004)

F1F2

F

(a) (b)

quence of gravity is a strong dependence of the modeangle on the crack length. Finally, all currently avail-able methods are hampered by frictional non-linearitiesin hinges that are attached to the sample and rotationpoints/joints in the loading frame.

The present work focuses on the design and realiza-tion of a miniaturized mixed mode bending setup whichenables in-situ delamination testing for the full range ofmode mixities. This setup is designed such that effectsof friction, gravity, and geometrical non-linearities areminimized. The performance of the setup is assessedusing specially-designed test samples supported by FEanalyses. In-situ delamination experiments performedon homogeneous bilayer samples are presented to provethe functionality of the setup and illustrate its capabil-ities for interfacial characterization.

2 Design of the miniaturized mixed mode bending(MMMB) apparatus

2.1 New configuration for mixed mode loading

A key constraint in the design of the new setup isits size, which should be small enough (i) to handlemulti-layer structures that are representative for stackedlayers present in SiPs and (ii) to be mounted in a com-mercially available micro tensile stage (Kammrath &Weiss GmbH, with an available design space of55×47×29 mm) which in turn fits in the chamber ofa SEM for in-situ delamination testing. Simple downscaling of existing MMB setups is not feasible becauseof the incompatible load frame configuration and thesample orientation that prevents in-situ microscopicobservation. Therefore, a new test configuration wasdeveloped that meets the above mentioned require-ments and realizes the MMB loading as shown inFig. 1c. The MMB configuration is preferred because

Fig. 3 Schematic representation of the new loading geometryfor mixed mode bending

it is standardized (ASTM D6671-01, 2001) andgenerally accepted for characterization of interfacialdelamination. A schematic representation of the newloading geometry of the MMMB apparatus is depictedin Fig. 3.

The setup consists of four rigid parts (A to D inFig. 3) connected with hinges. Advantages of the pres-ent design are: (i) its loading mechanism that allowsto access the maximum range of interfacial loadingmodes, from double cantilever bending (pure mode Idelamination), to pure mode II delamination, to endnotch flexure in a single setup; (ii) its compact geome-try (allowing it to be used in the chamber of an electronmicroscope); (iii) the insensitivity of its force measure-ments to its self weight, because the loading of thesample is done in the horizontal plane.

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Fig. 4 Schematicrepresentation of the loaddistribution in mixed modebending with the loads onthe sample (left) and thedecomposition of theseloads into mode I and modeII components (right)

The mixed mode bending loads are applied by a newlever mechanism. Part ‘C’ is pinned to the outside worldallowing it only to rotate in testing plane. By the appli-cation of a force PMMMB in a certain position on part‘D’, part ‘B’ moves downward while part ‘A’ movesupward, generating two oppositely directed forces PA

and PB on the sample. The ratio of the forces PA andPB depends on the position of the loading point on part‘D’, triggering different loading modes as discussed inmore detail below.

An analysis of the loads applied to the sample inthe new test setup shows that the loads transferred tothe sample, which are depicted in the left hand side ofFig. 4, can be written as:

PA = PMMMB(1 − ξ) (1a)

PB = PMMMBα

βξ (1b)

PC = PB

2(1c)

PD = −PA + PB

2, (1d)

where, ξ = Hγ

is a dimensionless shape parameterwhich represents the relative position of the appliedload, H is the corresponding absolute position andα, β, γ are the characteristic dimensions of the load-ing mechanism. Finally, PMMMB is the actual value ofthe applied mixed mode bending load.

The load on the sample can be decomposed in itsmode I and pure mode II components, PI and PI I ,respectively, which are defined by the load configura-tions depicted in the top and bottom right hand side ofFig. 4. The corresponding expressions for these modeI and mode II components are:

PI = PA + PD

2= PMMMB

(1 − ξ − α

4βξ

)(2)

PI I = PB = PMMMBα

βξ. (3)

Pure mode II loading is defined such that the twoloads PE and PF acting on the top and bottom armsof the specimen in the pre-cracked region (see bottomright of Fig. 4) are equal in order to have a zero modeI component on the interface. It is clear from the aboveanalysis that when the load is applied at the right handside of the upper lever (part ‘D’), i.e. when ξ = 0, theapplied loading corresponds to pure mode I loading:

PI = PA = PMMMB andPI I = 0. for ξ = 0. (4)

In the other extreme, when the load is applied at theleft end side of the upper lever (part ‘D’), i.e. whenξ = 1, the applied load resembles conventional ENFloading, i.e.

PI I = PB, butPI = −PMMMBα

4β. for ξ = 1. (5)

Since in this case PI �= 0, this ENF test can be con-sidered as a combined mode II and compressive modeI test. This compressive mode I component may causefriction between the two contacting layers. Therefore,the End Notch Flexure (ENF) case, which is frequentlyused for mode II fracture analysis, does not representa pure mode II test. The resulting friction leads to anoverestimation of the interface toughness in the exper-iments. Several studies (Davies 1997; Fan et al. 2007;Bing and Sun 2007) also highlighted this problem dueto friction when the ENF test is used to determine themode II delamination resistance. However, a position

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for which the mode I component is zero and conse-quently a pure mode II loading is obtained can be identi-fied in the new loading geometry. This position is givenby:

ξ = 4β

α + 4β. (6)

2.2 Design of the test apparatus

The new configuration, shown in the previous section,consists of a mechanism with several moving partsand hinges. The use of conventional hinges commonlyintroduces hysteresis in the force measurementsbecause of the non-negligible frictional dissipation dur-ing rotation at these hinges, particularly in a miniatur-ized setup. To avoid this hysteresis in the test apparatus,specially designed elastic hinges were used. Figure 5shows an FE model of the elastic hinge used in thepresent setup in a deformed state. Parameters definingthe geometry are also shown in the figure, as well as asafety element that is used to limit the rotation angleφ of the hinge to the maximum achievable rotation,φm , to prevent permanent deformations in the centralbridge of the hinge. φm is determined by the materialproperties, for a given geometry of the hinge. A Ti-6Al-4V alloy was used for the current setup because of its

Fig. 5 Finite element model of the elastic hinge used in thecurrent setup in a deformed state. Parameters defining the geom-etry (r : radius, h: width of the bridge and g: width of the safetygroove) are also shown

favorable combination of yield strength and Young’smodulus. Further details about the design of such elas-tic hinges can be found in Smith et al. (1997). In total,the setup contains 8 elastic hinges, each with dimen-sions of r = 2.5 mm, h = 50µm, g = 0.3 mm andt = 2 mm and a maximum rotation φ = 5.2◦, allowingfor an axial load of 20 N, which was verified throughFE simulations.

The main parts of the device are depicted in Fig. 6.The ‘Main Loading Mechanism’, MLM, is connectedto the ‘Position bar’ with the so called ‘Mode selector’.To change the applied mode angle, the Mode selectorcan be placed at various discrete positions on the MLMusing dovetail connectors. The arrow on the ‘positionbar’ indicates the fixed direction and position of theexternally applied force. Two ‘dovetail connectors’attached on both sides of the sample are used to attachthe sample to the MLM and the support hinge. The sup-port hinge connects the sample to the real world leavingonly one rotational degree of freedom. Figure 7 shows apicture of the whole setup placed inside a micro tensilestage. The range of sample dimensions which can fit inthe MMMB device are (length × width × height): 35×1–7.5× 0.5–6 mm.

Other elements in the setup design include (i) anadjustment mechanism to adapt the setup to a certainsample height (part 9 and 11 in Fig. 7), (ii) a screwmechanism (part 10 in Fig. 7) to adjust the alignmentof the sample, and (iii) different details to overcomeout-of-plane deflections in the device and to avoid fric-tion between the moving parts and the rigid bottomplate. The latter is achieved by two flexible pins posi-tioned in vertical direction, which support the systemand increase its stability.

2.3 Finite element analysis

The MMMB setup with a mounted homogeneousbilayer sample has been modeled in a finite elementprogram (MSC.Marc/Mentat), see Fig. 8. In total, thesample is described with 32,856 eight-node quadrilat-eral elements and the MLM is modeled with 52,196four-node quadrilateral elements. A predefined crackwas present in the bilayer sample, with a special cracktip mesh which has a rosette shape with transition ele-ments and quarter-point elements. Frictionless contactconditions were used between the two surfaces of thecracked region of the specimen. The material properties

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Fig. 6 a Design of the newMMMB device andb picture of the real device:1: Main LoadingMechanism (MLM), 2:Mode selector, 3: Sample,4: Support, 5: Supporthinge, 6: Dovetailconnector, 7: Position bar,8: Loading tip 9: Setscrewfor sample height, 10:Setscrew for alignment, 11:Sample height adjuster, 12:Protective plate

(a) (b)

Fig. 7 MMMB device,mounted in a micro tensilestage, with: 1: MainLoading Mechanism(MLM), 2: Mode selector, 3:Sample, 5: Support hinge, 7:Position bar, 9: Setscrew forsample height, 10: Setscrewfor alignment, 11: Sampleheight adjuster, 13: Microtensile stage, 14: Load cell

used for the sample and the MLM are given in Table 1.Simulations were performed by assuming plane strainconditions and linear elastic material behavior.

To assess the behavior of the setup over a completerange of mode mixities, simulations were performedby applying the load (PMMMB) at different load appli-cation positions (ξ ranging from 0 to 1) for a specimenwith various fixed crack lengths. Mode angles,ψ , werecalculated from the stress profiles ahead of the cracktip using,

ψ = arctan

(σ12

σ22

)δ

, (7)

whereσ22 is the normal stress,σ12 is the shear stress andδ is a characteristic distance from the crack tip alongthe interface. This distance needs to be tuned in caseof interfaces between dissimilar materials (Suo andHutchinson 1990), because of the existence of a com-plex stress state at those interfaces. In case of homo-geneous bilayer samples, δ can be any value withinthe K-dominance region, and therefore δ = 50µm was

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Fig. 8 a FE model of thesetup and sample withboundary conditions and bMagnified view of an elastichinge (indicated with arectangle in a)

(a) (b)

chosen in the present analysis. The mode angles, calcu-lated from the simulation results of a specimen with a6 mm pre-crack length at 400µm displacement appliedto the main loading mechanism (MLM displacement),are shown as a function of loading position ξ in Fig. 9a.This figure shows that the new MMMB setup permitsto access the complete range of mode mixities from 0◦to 90◦ mode angle by changing the loading position ξfrom 0 to 0.8, respectively. The mode angle is about 90◦for all positions in the range ξ = 0.8−1. As shown inthe previous section, pure mode II loading is obtainedwith a loading position ξ = 4β

α+4β , which, in the pres-ent setup, equals ξ = 0.8. When the load is applied atpositions ξ ≥ 0.8, an additional compressive mode Icomponent exists (Eq. 2) in addition to the pure modeII component acting at the crack tip. The presence of acompressive mode I component acting on the extrem-ity of the specimen does not significantly influencethe mode angle obtained from the stress field ahead ofthe crack tip. Note, however, that delamination experi-ments performed at those positions are not completelyrepresentative for pure mode II conditions, since thefriction induced at the fractured interface may influencethe delamination behaviour. This also applies to con-ventional ENF tests, which are widely used as a ModeII delamination test. Implications of these differencesare discussed in the results and discussion section.

The FE model was also used to evaluate the influ-ence of the applied MLM displacement and the cracklength on the mode angle. In Fig. 9b, the mode angleis shown as a function of MLM displacement for three

Table 1 Material properties of loading mechanism and sample

Brass (sample)a Ti-6Al-4V (MLM)b

Young’s modulus (GPa) 112 113.8

Poisson’s ratio 0.346 0.342

Yield stress (MPa) 204 880

a Determined from uni-axial tensile experimentsb Boyer et al. (1994)

different loading positions for a specimen with a 6 mmlong pre-crack. It can be observed that the mode anglestays almost constant with a maximum difference of4◦ between the minimum and maximum values of thethree considered load cases for position ξ = 0.75.Figure 9c shows the mode angle as a function of thecrack length for the same three loading positions at400µm displacement. The mode angle stays nearlyconstant for different crack lengths with a maximumdifference of 2◦ between the extreme values for posi-tion ξ = 0.75. The simulations confirm that the newMMMB setup provides access to the complete range ofmode angles whereby the mode angle does not changesignificantly with the applied displacement and/orevolving crack length.

3 Validation

The performance of the MMMB device is assessedin order to determine the influence of potential inac-curacies that may result from the geometry, machine

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Fig. 9 Mode angleobtained by FEM analysisas a function of (a) relativeloading position (ξ ) for acrack length of 6 mm at400 µm MLMdisplacement, (b)displacement applied toMLM for three loadingpositions for a 6 mm cracklength specimen, and (c)crack length for threeloading positions at 400µmMLM displacement

Fig. 10 End parts of (a)mode I/ mixed mode testsample and (b) mode II testsample. (c) Detailed imageof the elastic beam of themode II test sample

compliance, or other possible factors like clearanceat connectors. Specially designed test samples (shownin Fig. 10) have been used to this purpose. These arehomogeneous, single layer brass samples (i.e. with-out an interface and no propagating crack), with a welldefined notch, having an opening width of 30µm, rep-resenting an existing crack of a fixed length. The thick-ness of these samples is 1 mm. Figure 10 shows endportions of two different types of these test samples.The specimen shown in Fig. 10a is designed for modeI and mixed mode loading, and the specimen shown inFig. 10b is designed for mode II loading with a verti-

cal elastic beam at the end of the notch (see Fig. 10c).This special design is used in order to prevent contactbetween two arms of the notched portion in mode IItests (ξ = 0.8−1) where compressive mode I compo-nent is present.

Test samples with 5 different notch lengths (3, 6, 9,12 and 15 mm) were used to check the performanceof the setup as a function of the notch length. Fig-ure 11 shows the results of these five samples loadedat ξ = 0.5. A small amount of hysteresis was iden-tified as the result of some clearance in the dovetailconnectors. The relative hysteresis, defined as the dis-

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0 0.05 0.1 0.150

2

4

6

8

10

Displacement [mm]

F[N

]

3 mm 6 mm

9 mm

12 mm15 mm

Fig. 11 Load-displacement results for loading–unloading cycleswith ξ = 0.5 for five crack lengths. The numbers indicate thecrack length

sipated energy during a loading–unloading cycle, rel-ative to the energy supplied during loading was calcu-lated for all crack lengths and different loading posi-tions. The maximum relative hysteresis was found to be4% with a mean value of 3% for all positions betweenξ = 0−0.8. Comparison with digital image correlation(DIC) results confirmed that the hysteresis indeed orig-inates from a minor clearance in the dovetail connec-tors. Even though the amount of hysteresis is limited,and not considered further in this paper, this will beimproved in a future new connector design.

4 Experimental results and discussion

In-situ experiments with the new setup were conductedin a scanning electron microscope and under an opticalstereo microscope at high magnifications. A batch ofbilayer samples, consisting of two 0.3 mm thick springsteel layers glued together with an Araldite 2020 gluewith a thickness of ∼5µm was used in the experiments.At the end of each sample, a pre-crack of 3 mm lengthwas introduced with electro discharge machining. Allthe samples were heat treated at 80◦C for 3 days tomake the glue brittle and to fine tune the resulting gluestrength. On the basis of these experiments, the inter-face properties of the custom made interface structurewere characterized over a complete range of mode mix-ities, thereby demonstrating the practical functionalityof the setup. Before the experiment, samples were finepolished on one side (perpendicular to the plane of the

Fig. 12 Load-displacement response of a bilayer steel sampletested at position ξ = 0. The stiffness lines are used to calculatethe CERR

interface) to visualize the interface at high magnifica-tions. Small markers with a regular spacing of approx-imately 500µm were introduced on the polished sidewith a sharp knife. These markers were used as a refer-ence for tracking the position of the crack tip, since thecrack tip moves out of the field of view of the micro-scope during the experiment. The dovetail connectorswere attached on both sides of the samples with a strongglue at the pre-cracked end, after which the sampleswere carefully mounted into the setup.

Before actual loading of the specimen, each sam-ple was loaded at position ξ = 0 until the initiationof a sharp pre-crack was observed with a microscope.Finally, the position selector was reinserted at theappropriate position to carry out a delamination exper-iment at the desired mode angle. Actual loading wasapplied at a displacement rate of 3µm/s and load andcross-head displacement values were recorded with thebuilt in load cell (20N) and linear variable differentialtransformer (LVDT) respectively. During the test, thesample was unloaded and reloaded at regular intervals.At each load reversal, images of the crack tip wererecorded at a magnification of 250×. These imageswere used to determine the crack length correspondingto the load reversals in the post processing analysis.

Figure 12 shows the load-displacement result of anin-situ experiment conducted under a stereo micro-scope, where the sample is loaded at position ξ =0. Initially, at the beginning of the experiment, the

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Fig. 13 Comparison of the raw load-displacement curves basedon LVDT and DIC displacements

load increases rapidly with the displacement until thepre-crack starts propagating further. During this stage,the load remains approximately constant with smalldiscontinuities with increasing displacement. Throughreal-time optical and scanning electron microscopicvisualization, it is observed that the interface is del-aminating in a discrete fashion. The observed serratedbehavior is attributed to the discrete crack growthcaused by non-uniformity of the adhesion strength overthe interface area. This is confirmed by the microscopicobservation which shows that, sometimes, the crackfront jumps from one interface to the other. Apart fromthat, Fig. 12 also shows a small amount of hysteresisduring each unloading–loading cycle.

In order to further verify that the observed (ser-rated and hysteresis) behavior originates from the inter-face and not from the setup, a separate displacementmeasurement is done using a digital image correlation(DIC) technique (Chu et al. 1985). Vertical opening dis-placements are measured by tracking the center pointsof the dovetail connectors on both sides of the sample,from a series of pictures taken during the delaminationtest, with the image correlation software Aramis. Com-parison of LVDT and DIC based displacements, shownin Fig. 13, reveals an adequate agreement between thesetwo measurements. An important observation is thatthe DIC measurement also shows the small hystere-sis observed during the unloading–loading cycle withthe LVDT measurement. This indicates that the hys-teresis observed originates from the interface behav-

Fig. 14 CERR plotted as a function of mode angle. At 90◦ modeangle the CERR measured for position ξ = 0.83 is much lowerthan for position ξ = 1

ior and that this effect is indeed captured by the newsetup. A relatively small clearance (15µm) in the dove-tail connectors explains the small deviation betweenboth measurements, particularly in the initial portionsof the curves. From the above discussion, it can beconcluded that, (i) the new MMMB setup accuratelycaptures the behavior of the interface and (ii) smalldeviations are present in the measured displacementdue to some of clearance in the dovetail connections,which will have only a minor influence on the measuredinterface toughness.

Pictures taken during the load reversals were used todetermine the precise crack lengths at each load rever-sal. Stiffness lines were fitted to the loading curves asshown in Fig. 12. The critical energy release rate of theinterface was calculated from the area between the suc-cessive stiffness lines (the hashed region in Fig. 12),divided by the delaminated area corresponding to anincrease of the crack length from position a1 to a2.The resulting CERR values are shown as a function ofthe mode angle, as determined from FE simulations,in Fig. 14. From this figure, it is clear that the CERRincreases with increasing mode angle (towards posi-tion ξ = 1). It is noted that the CERR reported inthis study represents the macroscopic interfacial frac-ture energy, which may include significant contribu-tions from other dissipative mechanisms like plasticityin the layers apart from the intrinsic fracture energy.This is particularly the case when a ductile layer is pres-ent in the composite stack (Lane et al. 2000; Strohband

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Fig. 15 SEM micrographsshowing the mechanism ofdelamination: bridging ofthe small cracks into a maincrack. a Before crackevolution; b After crackevolution

(a) (b)

Fig. 16 SEM micrograph showing the precise crack tip location

and Dauskardt 2003). However, in the present study,the samples were heat treated for the specific reasonto make the glue layer brittle which was confirmedfrom in-situ electron microscopy observation. There-fore, the plastic contributions to the CERR were min-imized and hence neglected in this analysis. Anotherimportant observation from this figure is that a largedifference exists between the CERR values measuredat ξ = 0.83 and at ξ = 1. The position ξ = 0.83 isclose to pure mode II loading, and has a negligible com-pressive loading at the interface, whereas for the ξ = 1position (which resembles an ENF test), a compressiveload is applied at the cracked portion of the two lay-ers. The presence of this compressive load between thecracked layers leads to frictional dissipation causing anoverestimation of the CERR values (Davies 1997). Thisresult also indicates that to measure mode II delamina-tion CERR, one needs to apply a pure mode II loadinginstead of a conventional ENF loading.

In-situ measurements performed in the SEM allowedto (i) visualize the delamination mechanism at the inter-face and (ii) identify the precise crack tip location. InFig. 15, an interface with the crack growing from theright, is shown. Small pre-cracks of 50–100µm appearin front of the crack tip before the actual crack evolu-tion (Fig. 15a). These small pre-cracks grow and inter-connect, resulting in the propagation of the main crack(Fig. 15b).

From the in-situ measurements, the position of the“crack tip” was determined with a n accuracy of∼5µm,see Fig. 16. In principle, this allows to extract theCERR more accurately than with conventional MMBsetups. One remaining difficulty is tracking the cracktip as the delamination proceeds. This can be overcomeby using special markers and interrupting the test tobring the crack tip location back into focus.

5 Conclusion

A new miniature setup capable of applying a mixedmode bending load to a bilayer delamination samplewith a pre-crack was successfully introduced. The newsetup can be used for in-situ delamination tests inadvanced microscopic systems (e.g. SEM). The setupwas designed with elastic hinges to overcome hystere-sis due to friction. Comparison of the load-displace-ment diagrams from LVDT and DIC measurementsshowed that the present setup can accurately capturethe interface behavior. An analysis of the setup provedthe capability of the new setup to achieve a full range ofmode mixities, with a nearly constant mode angle as afunction of the crack length and crack opening displace-ment. The analysis also revealed that the conventionalENF test is not a true representation of a mode II delam-

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ination. The present setup can also apply a pure modeII loading (different from ENF loading) to obtain rep-resentative CERR values. The CERR measured for anENF test is considerably higher than for a pure mode IIloading, which was attributed to frictional dissipationbetween the two cracked surfaces. Finally, an in-situtest done in a SEM showed a high-resolution observa-tion of the delamination mechanism. In addition, thesein-situ tests allow for a more precise determination ofthe location of the crack tip.

Acknowledgements This research was carried out under theproject number MC2.05235 in the framework of the ResearchProgram of the Materials innovation institute M2i (http://www.m2i.nl), the former Netherlands Institute for Metals Research.Authors are grateful to M.P.F.H.L. van Maris for his assistancein conducting the experiments reported in this article.

Open Access This article is distributed under the terms of theCreative Commons Attribution Noncommercial License whichpermits any noncommercial use, distribution, and reproductionin any medium, provided the original author(s) and source arecredited.

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