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Composite Structures 89 (2009) 575–588

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Composite Structures

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Failure behaviour of honeycomb sandwich corner joints and inserts

Sebastian Heimbs a,*, Marc Pein b

a EADS Innovation Works, 81663 Munich, Germanyb Hamburg University of Technology, Denickestraße 17, 21073 Hamburg, Germany

a r t i c l e i n f o a b s t r a c t

Article history:Available online 11 December 2008

Keywords:Honeycomb sandwichCorner joint failure behaviourInsert failure behaviourFinite element modelling

0263-8223/$ - see front matter � 2008 Elsevier Ltd. Adoi:10.1016/j.compstruct.2008.11.013

* Corresponding author. Tel.: +49 89 607 25884; faE-mail address: [email protected] (S. Hei

In nearly all sandwich constructions certain types of joints have to be used for assembly, but little isknown about their failure behaviour. This paper deals with the investigation of the mechanical behaviourof three different corner joints as a right-angled connection of two sandwich panels and of two differentpotted inserts as a localised load introduction in Nomex� honeycomb sandwich structures with glassfibre-reinforced composite skins. For this purpose, experimental test series were conducted includingshear tests and bending tests of the corner joints and pull-out as well as shear-out tests of the threadedinserts. The failure mechanisms and sequences are described for each load case and the influence of thedifferent designs and of the loading rate is discussed. Based on these characteristics, finite element sim-ulation models were developed in LS-DYNA, which are able to represent the respective failure behaviours.

� 2008 Elsevier Ltd. All rights reserved.

1. Introduction

Sandwich structures with composite skins and a honeycombcore are widely used, especially in the aerospace industry, due totheir superior weight-specific bending stiffness and strength prop-erties. The failure behaviour of such sandwich panels is rathercomplex and has been investigated in numerous research studiesin the past. However, in virtually all technical sandwich construc-tions these panels have to be connected to subcomponents or pan-els have to be joined, and these joints are potential locations offailure as well, which have not been adequately treated in the tech-nical literature.

A number of different methods exist, how to introduce localisedloads into a sandwich structure, several of which are illustrated inFig. 1. Especially in aerospace design, threaded inserts, bonded intothe cellular core, are classically used for this purpose. The majortask of such an insert is to adequately transfer the load into thesandwich skins. In practice, tensile loads normal to the sandwichsurface and shear loads parallel to the surface are most relevant,since localised compression and bending loads are typicallyavoided due to large mounting surfaces, and torsion only occursduring the assembly of the construction and not in service. Becauseof usually very thin sandwich skins, the aim is to transfer the loadinto a preferably large area and into both skins. In case of honey-comb sandwich structures this may be achieved by filling the cellswith a potting compound in the insert installation area. Pull-outtests of inserts normal to the honeycomb sandwich structure are

ll rights reserved.

x: +49 89 607 23067.mbs).

very seldom documented and can be found for insert type aaccording to Fig. 1 in [1] and for type f in [2–4]. Raghu, Battleyand Southward [5] investigated the influence of potting diameteron the pull-out failure behaviour of insert type f. A recent study[6] investigates the influence of honeycomb core height, densityand skin thickness on the failure behaviour of insert type f underpull-out and shear-out loading. The fatigue behaviour of insertjoints under pull-out load was treated in [7–9]. Further papers dealwith pull-out tests of inserts in foam core sandwich, like in [10](types f, g and h), [11] (type f), [12] (type b) and [13] (partial metalinserts). Pull-out tests in balsa core sandwich structures are docu-mented in [14] (insert type f) and [15] (types b, f and h). The shear-out failure behaviour of a metallic bolt in a foam core sandwichwas analysed by Mares et al. [16].

In addition to these few quasi-static experimental studies,Thomsen [17,18] approached this topic analytically using a higherorder sandwich plate theory. As an example load case, he investi-gated the normal loading of a potted insert in an aluminium hon-eycomb sandwich structure. Numerical finite element analysesallow for the visualisation of stress distributions in the insert, coreand skins and have been performed for pull-out loads in honey-comb core sandwich in [1,19–21] and for foam core sandwich in[22–26].

Besides other general monographs on sandwich structures con-taining information on insert design [27,28], the most comprehen-sive collection of failure mode descriptions, test recommendationsor design guidelines can be found in the Insert Design Handbook[29] of the European Space Agency (ESA), which is also used as areference in most of the other papers listed here. Besides numerousstrength vs. core height diagrams for different load cases, insert

Fig. 1. Overview of different methods for local load introductions into honeycomb sandwich structures.

576 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

diameters and honeycomb core types, also analytical equationsfor the estimation of pull-out and shear-out strength values aregiven.

Compared to these few studies and information on inserts inhoneycomb sandwich structures, almost no literature is publishedon the failure behaviour of sandwich corner joints, also called L-joints. Some industrial relevant design options of corner joints,which are illustrated in Fig. 2, are shown in [27,28,30,31]. In thetextbook of Noakes [32] a comprehensive overview on the manu-facturing of folded corner joints is given, which is a state-of-the-art technique in modern sandwich design, often found in aircraftinterior components. Such folded corner joints of type f accordingto Fig. 2 in aluminium honeycomb sandwich specimens weretested by Joulia and Grove [33], providing the only informationon failure loads. However, the state of stress in their combinedbending-shear test is not clear and no comparison to other jointdesigns is given. Furthermore, no detailed description of the failurebehaviour is provided in this study. Carruthers [34] investigatedthe crash performance of foam core sandwich crash boxes withtwo different corner joints (type b and e) experimentally withoutcharacterising the single corner joints separately. The only finiteelement analysis of three different L-joint designs of type b in alu-minium honeycomb sandwich is documented in [35]. However,not the failure characteristics but only the dynamic behaviourwas investigated in a modal analysis.

Fig. 2. Overview of different methods of corner

This paper intends to gain insight into the failure behaviour ofinserts and corner joints of sandwich structures and to cover as-pects that have not been treated previously. The material used inthis study is the most relevant sandwich structure used in the air-craft industry with Nomex� aramid paper honeycomb cores and fi-bre-reinforced composite skins. Two different insert types andthree different corner joints are tested under various loading con-ditions including pull-out, shear and bending, and their failurebehaviour is characterised in detail. In addition to static testing,also the influence of the loading rate is addressed. Furthermore,analytical calculations of the failure loads are presented and com-pared to the experimental results. Subsequently, within a numeri-cal analysis with the explicit finite element code LS-DYNA,simulation models are generated and modelling methods to coverthe failure behaviour are derived. In this context, both detailedmeso-scale models and rather simple macro-models are covered.

2. Failure behaviour of inserts

2.1. Materials and specimen manufacturing

The following study is focused on honeycomb sandwich panelswith two different standard metallic, threaded inserts used in theaerospace industry. The 15 mm thick sandwich structure consists

joints in honeycomb sandwich structures.

S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588 577

of a Nomex� honeycomb core with a density of 48 kg/m3 and a cellsize of 3.2 mm (type Schütz Cormaster C1-3.2-48). The skins aremade of one single prepreg ply of woven (satin weave 1/7) E-glassfibre-reinforced phenolic resin (type Stesalit PHG 600-68-50) withan average cured ply thickness of 0.24 mm and a fibre volume frac-tion of 48%. The material properties are shown in Table 1. Beforecuring of the sandwich specimens, the honeycomb cells within acircle of approx. 38 mm diameter were filled with an epoxy-basedpotting material (type Cytec BR� 632 P4) for the later positioningof the inserts. In the co-curing process no additional adhesive filmbut only the resin of the skin prepregs was used for the skin-corebonding. The specimen plates were cured in an autoclave using aflat mold and a vacuum bag (2 bar, 125 �C, 90 min curing). Two dif-ferent self-locking steel inserts, specified by the US National Aero-space Standards NAS 1833-C3-370 [36] (diameter 14 mm, height9.4 mm) and NAS 1835-C3-430 [37] (diameter 17.4 mm, height11 mm), were used. The latter one is a so-called floating insert witha moveable nut inside the insert housing for compensating assem-bly inaccuracies. After curing of the sandwich panels a hole was

Table 1Mechanical properties of sandwich specimens in this study.

Skin properties Glass–fabric reinforced phenolicEskin 20 GPa mskin 0.059Gskin 1.8 GPa rskin 150 MPatskin 0.24 mm

Core properties Nomex� honeycombEL 0.58 MPa GLT 41.9 MPaEW 0.33 MPa GWT 25.5 MPaET 84.8 MPa GLW 0.31 MPasWT 0.9 MPa sLT 1.21 MPahcore (insert test) 14.6 mm hcore (corner test) 9.5 mm

Potting properties EpoxyEpotting 1.05 MPa mpotting 0.3rpotting 68.7 MPa rpotting 19 mm

Fig. 3. Cross-section of insert types N

drilled in the middle of the potted area, the insert was positionedinside this hole, and an epoxy resin (type Huntsman Araldite�

2011) was injected through an opening for the fixation of the insert(Fig. 3). The final test specimens were cut to a size of127 mm � 127 mm, each having one insert in the middle.

2.2. Experimental procedures

The pull-out and shear-out behaviour of these sandwich and in-sert configurations was to be investigated. However, no standard-ised test methods for this purpose exist. Some recommendationscan be found in the ESA Insert Design Handbook [29] and some air-craft manufacturers have developed their own test methods [38].Based on these data, two test rigs were designed and fabricatedwithin this study, shown in Fig. 4. The pull-out test rig features acircular hole of 100 mm diameter to reduce edge effects, underwhich the specimen is placed. The insert is pulled out vertically.For the fixation of the specimens in the shear-out test rig, the hon-eycomb core at the specimen sides had to be replaced by a solidmaterial with clearance holes. Once the specimen is bolted to thetest rig, the insert is loaded in the sandwich plane by pulling ashear plate. In this case, the weft direction of the skin’s woven fab-ric was oriented parallel to the loading direction. For both experi-ments, bolts of the type 0.190-32 UNJF-3A were used, a new onefor each test. The test rigs were mounted on an Instron 556610 kN universal testing machine. To investigate the influence ofthe loading rate on the failure behaviour, the two cross-headspeeds of 1 mm/min (static) and 500 mm/min (dynamic) weretested. This led to failure after 200 s and 0.4 s, respectively.

2.3. Experimental results and failure process

The pull-out test results of six specimens for each insert typeand load case showed a high level of reproducibility, representativecurves are shown in Fig. 5. The curves of both insert types are verysimilar. In general, two distinct points can be identified, character-

AS 1833 (a) and NAS 1835 (b).

Fig. 4. Test rigs for pull-out (a) and shear-out (b) testing.

Fig. 5. Force–displacement results of pull-out tests and post-failure cross-sections.

578 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

ising the failure process under vertical loads. In the beginning, theload increases due to elastic deformations. The first damage atpoint A, leading to a slight drop of the load level, is attributed toa transverse shear failure of the honeycomb core adjacent to thepotting mass. This can be confirmed by a view on the lower spec-imen surface at that time. The whole potted area exhibits a verticaldisplacement with respect to the rest of the panel, which is onlypossible under a large shear deformation of the core. Then the loadincreases again up to point B, where it drops because of a tensilerupture of the potted honeycomb cells combined with a shear fail-ure at the cell wall interfaces, which can be seen in Fig. 5. The aver-age static peak loads of both insert types at that point show almostequal values with 2220 N (standard deviation: 106.1 N) and2194 N (standard deviation: 105.7 N). However, the drop of theload level occurs at a larger displacement for insert type NAS

1835. This may be ascribed to the lower surface of the insert beingbonded to the potting material, which is not the case for insert typeNAS 1833 (see Fig. 3), therefore, failure occurs earlier. After point Bthe post-damage behaviour is mainly characterised by a peeling/debonding of the upper skin and friction effects while pulling outthe insert. In case of the dynamic loading, the curves show thesame characteristics as in the static tests, and therefore the samefailure process, but a higher load level. Also the shear failure occursat a higher load. This may result from the rate dependency of theNomex� honeycomb core structure, which was described byFeichtinger [39] and Heimbs et al. [40,41]. Due to micro-inertial ef-fects, the shear strength under transient loading is significantlyhigher than in the static load case.

The results of the shear-out tests are shown in Fig. 6. As in thepull-out tests, both insert types show similar load curves. The aver-

Fig. 6. Force–displacement results of shear-out tests and post-failure images.

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age maximum force in the static tests is higher for type NAS 1835with 4360 N (standard deviation: 553.2 N), compared to 3921 N oftype NAS 1833 (standard deviation: 149.5 N). This can once morebe explained by the connection of the lower insert surface to thepotting material. Again, two characteristic points in the load–dis-placement diagrams are noticeable. After an elastic deformation,a tensile cohesive failure of the epoxy bond occurs at point A fol-lowed by a first drop of the load level. The final drop to almost zeroat point B happens because of a bearing failure with the uppercomposite skin and the potting compound below being shearedoff (Fig. 6). The whole block of potting material in front of the insertin load direction is separated from the surrounding material andthe lower skin, which remains intact. One important aspect is thedependency of the failure process on the number of filled honey-comb cells in front of the insert in load direction, or in other wordsthe insert position, which explains the relatively high values of themaximum forces’ standard deviations. In case of only a small quan-tity of potting material, the bearing failure occurs earlier – some-times even earlier than the cohesive failure – and therefore at alower load level. A number of different configurations with moreor less potting material were investigated in this framework. Forinsert type NAS 1835 the failure loads varied from 3649 N (justone filled cell ahead of insert) to 4920 N (five filled cells ahead ofinsert). The comparison of the static and dynamic tests againshows a rate dependency with higher load levels for the transientloading.

2.4. Analytical investigation

An analytical calculation of the insert pull-out force was con-ducted with the equations provided in the Insert Design Handbook[29]. This calculation assumes a core shear failure as the limitingload and is based on the radius rpotting of the potting mass, thethicknesses tskin and hcore of skins and core as well as their elasticmoduli and the core shear strength sTW. Confirming the experi-mental findings, the insert type, diameter or height have no influ-ence on the pull-out strength, which is only determined by thesandwich configuration and potting size. For the following calcula-tion the values given in Table 1 were used.

Fpull-out ¼2prpottinghcoresWT

CKmax¼ 1630 N ð1Þ

with c ¼ bbþ 1

; b ¼ hcore

tskinð2Þ

Kmax ¼rpotting

rs max1�

ffiffiffiffiffiffiffiffiffiffiffiffiffirsmax

rpotting

rev rpotting�rsmaxð Þ

� �ð3Þ

rs max ¼rpotting

1� ek vrpottingð Þn ð4Þ

v ¼ 1tskin

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiGskin

Eskin12 1� m2

skin

� � b2þ 1þ 2

3b

� �sð5Þ

k ¼ �0:931714n ¼ 0:262866

The calculated pull-out force at shear failure is 1630 N, which is 23%lower than the experimental results and therefore a rather conser-vative result.

The calculation of the maximum shear-out force according tothe Insert Design Handbook [29], based on the radius of pottingmass rpotting, the skin’s thickness tskin and compressive strengthrskin as well as the core’s shear strength sWT, leads to 3967 N.

Fshear-out ¼ 8r2pottingsWT þ 2tskinrpottingrskin ¼ 3967 N ð6Þ

The correlation to the experimental values of 3921 N and 4360 N forboth insert types is much better than in case of the pull-outstrength with a maximum deviation of 10%.

The analytical calculation of the shear-out force can thereforebe used for an estimation of the insert strength, while the calcula-tion of the pull-out force is very conservative. But possible varia-tions and uncertainties of the parameters used in these equationshave to be kept in mind.

2.5. Numerical simulation

Meso-scale models of the insert and honeycomb sandwichstructure require a large amount of modelling work and long calcu-lation times and are not suitable for an implementation into a lar-ger structure’s model. However, they may be used to analyse stress

580 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

distributions and understand cell wall deformations and failuremechanisms of the insert under load, since this insight is not pos-sible during the experiment. Such a meso-scale model with quar-ter-symmetry was developed for dynamic pull-out simulationsusing the explicit finite element code LS-DYNA (Fig. 7). Both theskins and the honeycomb cell walls were modelled with 4-nodeshell elements and the orthotropic composite material modelMAT54. For the rest of the model, i.e. the potting mass, epoxy resin,steel insert and steel bolt, 6-node wedge or 8-node brick elementswith the isotropic material model MAT24 were used. This quarter

Fig. 7. Models of insert pull-out in honeycomb sa

model with an edge length of 50 mm had a total of 62000 ele-ments. The nodal boundary conditions with a circular support wereapplied corresponding to the test rig. The bolt was pulled using atime-dependent linear displacement function. This model couldbe used to prove the theory that a core shear failure under trans-verse pull-out loads occurs as the first failure mode, see Fig. 7.

In addition to this meso-scale model, an alternative modellingapproach on a macro-scale was developed, which is much morefeasible to be implemented into a larger sandwich structure modelwith a higher number of inserts. It should be as simple as possible,

ndwich: meso-scale (a) and macro-scale (b).

Fig. 8. Comparison of experiment and simulation for pull-out and shear-out loading.

S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588 581

while still being able to represent the failure behaviour correctly.For this model, the honeycomb core was homogenised with 8-nodecontinuum elements, and the orthotropic honeycomb materialmodel MAT126 in LS-DYNA was used. The library of LS-DYNA of-fers different options for joint modelling. In this case, the spotweldoption was used for the insert modelling. Such spotweld elementsare solid elements, which connect two surfaces and can be posi-tioned mesh-independently using a spotweld-contact formulation.Different failure criteria can be adopted for these spotweld ele-ments, making a pull-out or shear-out failure modelling possible.In this study, a quadratic failure criterion with respect to the nor-mal force FN and shear force FS in the element was chosen. Themaximum values Fpull-out and Fshear-out could directly be taken fromthe experimental results, since only one spotweld element wasused.

FN

Fpull-out

� �2

þ Fs

Fshear-out

� �2

� 1 ð7Þ

The final model as a quarter section is shown in Fig. 7b. With thismodelling approach, only 100 elements are necessary for the samespecimen size as for the meso-scale model before. Besides pull-outsimulations, the same model was also used for shear-out simula-tions by simply changing the boundary conditions.

Fig. 9. Overview of manufacturing methods of the thre

The simulation results of the simplified macro-model are shownin Fig. 8. For the pull-out loading, the curves match with a goodcorrelation. Especially the core shear failure prior to total ruptureis covered correctly thanks to the honeycomb material model.

Also in case of the shear-out loading the elastic behaviour andfailure in the simulation correlate to the experiment. However,the residual strength after first damage, which was observed inthe test results, can not be covered by the numerical model. Thereason for this is the spotweld model being no physical represen-tation of the real failure process with a bonding failure. It simplyuses the defined maximum shear force and deletes the elementwhen reaching it. No post-damage behaviour can be covered.

In both curves, the drop of the load level after reaching the de-fined maximum forces happens to some extent earlier than in theexperiment, which corresponds to a slightly lower energy absorp-tion at insert failure. To change this, either the maximum force lev-els need to be increased or the post-damage shear behaviour of thehoneycomb material model needs to be adjusted, since these twofactors influence the load curve.

Nevertheless, this simple insert modelling approach with spot-weld elements is feasible of representing the normal and shearfailure load levels sufficiently. Even strain rate effects can becovered, corresponding to the experimental findings. In this case,a different failure criterion has to be chosen in the LS-DYNA

e different corner joints investigated in this study.

Fig. 10. Test methods for bending (a) and shear testing (b) of corner joints.

Fig. 11. Force–displacement results of bending tests and post-failure images.

582 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588 583

spotweld definition, requiring the input of the maximum forces asa function of the effective strain rate.

3. Failure behaviour of corner joints

3.1. Materials and specimen manufacturing

The experimental investigation of the corner joints was per-formed on 10 mm thick sandwich specimens with the same No-mex� honeycomb core and the same GFRP skin material as in theinsert study. A total of three different corner joints were analysedand compared.

The first one is the simplest variant, a bonded butt joint accord-ing to type a in Fig. 2. In one of the two pre-cured sandwich panelsthe upper skin and core were removed. The second panel wasbonded at this position with an epoxy-based 3M Scotch-Weld9323 B/A adhesive resulting in two contact surfaces (Fig. 9a).

The second variant is based on the ‘mortise and tenon’ methodaccording to Fig. 2c. At the edges of both pre-cured panels 60 mmwide pockets were milled out while keeping the outer skin intact,resulting in tenons of 50 mm width (Fig. 9b). Both panels werejoined and the same adhesive as before was used for the bonding.

The third corner joint features the ‘cut and fold’ technique as avery efficient and industrially relevant design. In contrast to theother two methods, just one pre-cured sandwich plate is necessary.In the middle of this plate a 16 mm wide and 3.5 mm deep groovewas milled out from the upper surface. The open honeycomb cellswere filled with a Mankiewicz Alexit FST compound. Then the platewas folded about 90�. Afterwards, a 60 mm wide reinforcement

Fig. 12. Force–displacement results of s

layer (glass fibre cloth with Bakelite EPR L 43 epoxy resin) waslaminated onto the surface to fixate the corner joint (Fig. 9c). Theadvantage of this method is that the outer skin remains intactand that even very complex constructions can be folded very effi-ciently from one pre-cured sandwich plate.

All specimens had a side length of 140 mm, whereas at bothedges the honeycomb core was replaced in a length of 25 mm bya solid block for load introduction purposes. The specimen widthwas 250 mm. The ribbon direction of the core cells was orientedperpendicular to the corner line.

3.2. Experimental procedures

Standard test methods for the evaluation of the strength ofsandwich corner joints do not exist. Besides some investigationson the failure behaviour of sandwich T-joints [42–46], the studyin [33] seems to be the only reference, where such tests on sand-wich L-joints are documented. However, in those tests the cornerjoints are exposed to both bending and shear loads, making a sys-tematic analysis of stress states and failure modes difficult. There-fore, two test procedures were developed for separatelyinvestigating the specimens under bending and shear loads(Fig. 10). In the bending test the sandwich edges were supportedby linear bearings while the load was applied directly onto the cor-ner, bending the two sides apart. A silicone strap was used underthe loading plate to assure a load introduction over the whole spec-imen length. In the shear test one specimen side was clamped ontoa rigid surface, also using a silicone strap, while the other one waspulled perpendicularly. This results in a shear loading of the corner

hear tests and post-failure images.

584 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

joint. For each experiment at least six specimens of all three cornerjoint types were tested. The loading rate in these tests was keptconstant at 50 mm/min.

3.3. Experimental results and failure process

Force–displacement curves of representative specimens in thebending test are illustrated in Fig. 11, showing the pre- and post-damage behaviour. The butt joint corners are characterised bythe lowest failure load. After a nearly linear elastic beginning, thebonding of the two panels fails abruptly on the inside of the joint.The post-failure residual strength is relatively low. The outer skinis bent at a low load level under increasing displacements.

The ‘mortise and tenon’ specimens exhibit a similar bendingstiffness with a slightly higher failure load. At the peak load thebonding of the tenons on the inside of the corner joint fails, whilethe bonding on the outside remains intact. The residual strength ishigher than for the butt joint specimens and is characterised by thebending of the outer sandwich skins and the pull-out of the tenons.

The ‘cut and fold’ specimens show the highest failure loads andalso the highest bending stiffness, which is the result of the solidfilling compound in the corner. The failure mode is a rupture ofthe filled and neighbouring unfilled honeycomb cells on one sideof the corner joint, which cannot be avoided by the reinforcementlayer. Afterwards, a bending of the outer skin occurs at the failedside, while on the opposite intact side the reinforcement layer iscontinuously delaminated from the inner sandwich skin. Thesemechanisms lead to the highest post-failure load level compared

Fig. 13. Comparison of experiment and simulation for co

to the other corner joints. The damage tolerance of this folded cor-ner is superior, because as long as the delamination has notreached a critical extent, loads can still be transmitted throughthe intact outer sandwich skin.

The force–displacement curves of the corner joint shear testsare almost congruent for the elastic regime until first failure occurs(Fig. 12). The reason for this behaviour is a simple transverse shearloading of the honeycomb core in all three cases, followed by a coreshear failure, independent of the corner joint type. The averagemaximum force Fmax for the butt joint specimens has a value of2839 N (standard deviation: 90 N). With respect to the shearedsurface A, this results in a shear strength of 1.2 MPa:

sLT ¼Fmax

A¼ 2839 N

250 mm � 9:5 mm¼ 1:2 MPa: ð8Þ

This value corresponds exactly to the data sheet value of theNomex� honeycomb core’s shear strength of 1.21 MPa. For the‘cut and fold’ specimens the peak load is slightly higher with3310 N (standard deviation: 171 N). This effect may be attributedto the filling compound, which has also partly been filled into theshear-loaded honeycomb cells, leading to an increase of shearstrength. After this core shear failure the real loading and character-isation of the corner joint takes place. For the butt joint specimens adebonding of the two sandwich panels occurs under a decliningload curve. The mortised corner first shows a slight increase ofthe load level due to the larger number of bonded surfaces, but oncethe debonding is initiated, the load level also declines. The foldedcorners show a different behaviour. Because of the reinforcement

rner bending and shear test of butt joint specimens.

S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588 585

layer being clamped in the specimen fixation and being loaded intension, no complete failure of the corner joint occurs. A crack be-tween the upper connection of filled and unfilled honeycomb cellsas well as a delamination of the reinforcement layer develop. Thishappens at a comparably high and increasing load level. After largedeformations a crack occurs in the middle of the reinforcementlayer, leading to a declining load level.

3.4. Numerical simulation

Just as for the inserts, a modelling method in LS-DYNA wasdeveloped, which on the one hand is able to represent the cornerjoint’s failure behaviour correctly, and on the other hand is as sim-

Fig. 14. Comparison of failure process in experiment and si

ple as possible to be implemented into large-scale models of sand-wich constructions with a number of such joints. The same macro-scale modelling approach with shell elements for the compositeskins (MAT54) and brick elements for the homogenised honey-comb core (MAT126) was adopted here. The material models andmechanical properties are identical to the insert model.

The experiments have shown that the primary driver of the cor-ner joint failure is debonding or the disconnection of bonded sur-faces. Therefore, the modelling of the corner joint failure is basedon contact formulations with a failure option. In LS-DYNA suchcontact definitions are called tiebreak contacts, since first theytie two surfaces together, and once the interface stresses meet adefined failure criterion, the contact breaks open and the surfaces

mulation for corner shear test of butt joint specimens.

586 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

are separated. In this case maximum normal stresses NFLS andmaximum shear stresses SFLS in the interface are combined in aquadratic failure criterion

jrjNFLS

� �2

þ jsjSFLS

� �2

� 1: ð9Þ

Those maximum stress values for the bonded contact interfaceswere unknown and had to be adjusted within parameter studiesin correlation to the experimental results.

In case of the butt joint corner only two contact formulationshad to be defined in the model. The interface failure stresses couldbe adjusted so that a satisfactory representation of both the failuremodes and the load curves could be achieved (Fig. 13). However, anexact compliance of the curves could not be achieved, which maybe ascribed to imperfections in the hand-made sandwich speci-mens that are not fully covered by the rather ideal model. Further-more, the silicone strap, used in both tests, was not included in thesimulation. One requirement for a good compliance is that the fail-ure process is reproduced correctly. This could be achieved with ahigh degree of accuracy, which can be seen in the illustration of thecorner shear test simulation in Fig. 14. At first, core shear failureoccurs. Then the contact at the inner side of the joint opens, leadingto a crack. This crack grows until also the outer contact fails. Sim-ilar results could be achieved for the ‘mortise and tenon’ speci-mens. The difference here lies in the higher number of contactsurfaces.

Fig. 15. Comparison of experiment and simulation for corn

In the model of the ‘cut and fold’ specimens an isotropic materialmodel was used for the filled corner and the curvature of the edgewas approximated in the mesh. Since in the experiment a discon-nection of the filled and unfilled cells could be observed, tiebreakcontact formulations between the isotropic material and the regularhoneycomb elements were generated in the finite element model.Further contact definitions with different failure parameters wereintroduced between the sandwich skin and the reinforcement layer,which was modelled as a separate layer of shell elements. Despite itssimplicity, this simulation model allows for a good representation ofthe failure process. However, the difference in the load curves isslightly higher (Fig. 15). This may be attributed to the same factorsas described before, i.e. possible imperfections and the neglect of thesilicone strap. In addition, no exact material data were available forthe filling compound, which strongly influences the corner’s proper-ties. The failure process in the corner shear test simulation is shownin Fig. 16 and it correlates well with the experimental results. Afterthe core shear failure, a crack develops between the isotropic cornerelements and the honeycomb core, which is responsible for the firstload drop-off. Afterwards, the contact of the reinforcement layerfails successively as in the experiment. The simulation results andthe contact interface strength values are without doubt mesh sizedependent. However, in this study only one mesh size was used.Although the curves do not match exactly, this modelling approachshows the potential to be a simple and efficient technique to imple-ment a corner joint failure possibility into a larger finite elementmodel of a sandwich construction.

er bending and shear test of ‘cut and fold’ specimens.

Fig. 16. Comparison of failure process in experiment and simulation for corner shear test of ‘cut and fold’ specimens.

S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588 587

4. Conclusions

The failure behaviour of different types of potted inserts andcorner joints in Nomex� honeycomb sandwich structures wasinvestigated experimentally. Insert pull-out tests showed a coreshear failure to occur first, before the potted cells fail under tensilerupture. Under shear-out loading, the potted cells together withthe upper skin fail in shear with the insert position within the pot-ted area having a significant influence on the results. However, theinsert type had no influence, only the potting diameter. Thefailure stresses in both experiments were affected by the loadingrate.

The failure behaviour of the corner joints under bending orshear loads was primarily driven by the debonding of the respec-

tive connection surfaces. The ‘cut and fold’ corner showed thehighest failure loads and a superior post-damage behaviour.

For both inserts and corner joints, simple and efficient model-ling methods in LS-DYNA based on spotweld elements or tiebreakcontact formulations have been developed, which are able to coverthe failure behaviour with an acceptable degree of accuracy andcan be implemented into large-scale models of sandwichstructures.

Acknowledgement

This work was performed within the project INTECK, with par-tial funding by the Free and Hanseatic City of Hamburg, Germany.Sincere thanks are given to COMTAS Composite, Hamburg for

588 S. Heimbs, M. Pein / Composite Structures 89 (2009) 575–588

manufacturing the specimens and Dipl.-Ing. Manfred Heimbs formanufacturing the test rigs.

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