strength evaluations of sinusoidal core for frp sandwich bridge deck panels

Composite Structures 92 (2010) 1561–1573

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Composite Structures

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Strength evaluations of sinusoidal core for FRP sandwich bridge deck panels

An Chen, Julio F. Davalos *

Department of Civil and Environmental Engineering, West Virginia University, Morgantown, WV 26506-6103, USA

a r t i c l e i n f o

Article history:Available online 17 November 2009

Keywords:Honeycomb FRP sandwich panelStrength evaluationExperimental investigationOut-of-plane compressionTransverse shear

0263-8223/$ - see front matter � 2009 Elsevier Ltd. Adoi:10.1016/j.compstruct.2009.10.039

* Corresponding author. Address: 611C Engineerinment of Civil and Environmental Engineering, Westtown, WV 26506-6103, USA. Tel.: +1 304 293 9928; f

E-mail address: [email protected] (J.F. D

a b s t r a c t

Fiber-reinforced polymer (FRP) sandwich deck panels with sinusoidal core geometry have shown to besuccessful both in new construction and the rehabilitation of existing bridge decks. This paper is focusedon an experimental study of the strength evaluations of a honeycomb sandwich core under out-of-planecompression and transverse shear. The sinusoidal core is made of E-glass Chopped Strand Mat (ChSM)and Polyester resin. The compressive, tensile and shear strengths were first obtained from coupon tests.The out-of-plane compression tests were performed on representative single-cell volume elements ofsandwich panels, and the tests included ‘‘stabilized” samples to induce compression failure, and ‘‘bare”samples to induce local buckling of the core. Finally, four-point bending tests were conducted to studythe structural strength behavior under transverse shear. Two types of beam samples were manufacturedby orienting the sinusoidal wave either along the length (longitudinal) or along the width (transverse).Both typical shear failure mode of the core material and delamination at the core–facesheet bondinginterface were observed for longitudinal samples. The failure for transverse samples was caused by corepanel separation. For both single-cell and beam-type specimen tests, the number of bonding layers, i.e.,the amount of ChSM contact layer and resin used to embed the core into the facesheet, and the core thick-ness are varied to study their influence. The experimental results described herein can be subsequentlyused to develop design guidelines.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

In recent years, the demands in civil infrastructure haveprovided opportunities for development and implementation of aFiber-Reinforced Polymer (FRP) sandwich panel with sinusoidalhoneycomb core both for new and rehabilitation projects. The con-cept of a lightweight and heavy duty FRP panel, with sinusoidalwave core configuration in the plane extending vertically betweenface laminates (Fig. 1) was developed for highway bridge decks andsuccessfully implemented in about sixteen bridge projects [35,24].

The development of standards and guidelines is needed in orderto promote wider acceptance of composite sandwich panels in con-struction. Characterizations of stiffness and strength properties arenecessary to facilitate the development of design guidelines. Mucheffort has been devoted to the stiffness modeling and optimizationof the present FRP sandwich panel. Davalos et al. [24] developedequivalent orthotropic properties representative of the complexhoneycomb geometry, and they presented a simplified analysisprocedure that can be used in design applications. Xu et al. [42] de-rived an analytical solution for the transverse shear stiffness of

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composite honeycomb with general configurations. However, onlylimited studies are available on strength evaluations. Davalos andChen [23] provided an analytical model for the buckling capacityof FRP panels with two loaded edges partially constrained. By con-sidering skin effect, Chen and Davalos [20] gave an accuratedescription of the transverse shear modulus and the interfacialstress distribution for honeycomb sandwich structures with sinu-soidal core. Therefore, in order to correlate with existing modelsand develop design guidelines for this sinusoidal honeycomb coregeometry, there is a need to further characterize the strength prop-erties of this product through experimental investigations, whichis the motivation of this study.

A previous study by DeTeresa et al. [25] indicated that honey-comb core materials for sandwich structures are primarily sub-jected to shear and through thickness compression. Therefore,the two topics of concern in this study are out-of-plane compres-sion and transverse shear.

Chopped Strand Mat (ChSM) is used for the core material, whichis composed of E-glass fibers and polyester resin. The facesheet ismade of several layers of ChSM, 0�/90� E-glass fiber and polyesterresin. The ChSM material is effectively used at the interface be-tween core and facesheet to serve as a bonding layer betweenthese two component parts which are joined by contact moldingprocess, where the precured core is ‘‘embedded” into the ‘‘wet”ChSM layer on the facesheet. Thus, as the first step, the material

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Longitudinal Direction

Transverse Direction

Fig. 1. FRP panel with sinusoidal core configuration.

1562 A. Chen, J.F. Davalos / Composite Structures 92 (2010) 1561–1573

strengths of ChSM were obtained from coupon samples tested incompression, shear (Iosipescu), and tension.

Next, compression tests of sandwich single-cell symmetric sam-ples (Fig. 1) were carried out to study the core behavior under out-of-plane compression. Two important parameters, the number ofChSM bonding layers (i.e., thickness of ChSM) and the thicknessof the core elements were considered for each specimen type.Two testing methods were used: one method consisted of bondingthe outer surfaces of the facesheets to steel plates, and the loadwas applied over the steel plate; this method is called stabilizedcompression test and is designed to induce compression failureof the core. The other method was to apply a distributed load di-rectly over the specimen facesheet using an elastomeric pad; thismethod is called bare compression test and it is used to induce pri-marily buckling of the core. It is shown that a typical compressionfailure occurs in the stabilized compression test, while local buck-ling occurs first in the bare compression test. The ultimate bucklingstrength in the bare compression test is sensitive to the number ofChSM interface bonding layers and the thicknesses of core ele-ments. It is shown that the actual case of core–facesheet rotationalconnectivity lies between simply-supported condition and fully-restrained condition.

Finally, four-point bending tests were carried out to evaluateflexural behavior primarily for shear and delamination failures.Two types of specimens were defined, according to the orientationof the sinusoidal wave of the core, namely longitudinal and trans-verse specimens (Fig. 1). The width of each sandwich specimenconsisted of single-cell and symmetric elements aligned in the lon-gitudinal or transverse directions (Fig. 1). For each type, the bond-ing layer thickness and core-element thickness were varied tostudy their influence on failure mode. The test results showed thatfor longitudinal specimens a shear failure of the core materials oc-curred for samples with sufficient bonding layers, while for lessernumber of bonding layers the failure was due primarily to interfacedelamination. For transverse specimens, core separation of single-cell contact surfaces was a typical failure mode.

2. Literature review

2.1. Strength properties of ChSM materials

ChSM is the material used for the core elements of the sand-wich; and the constituents are E-glass fiber and Polyester resin.In the aerospace industry, both modeling and testing of strength

properties of composite materials have been achieved throughthe industry’s relatively long history of applications. The modelingof ChSM can be dated back to the 1970s, mostly focused on com-pression properties. Hann [28] simulated the random compositeby an equivalent laminate consisting of in-plane unidirectionalplies in every direction of the laminate. Using maximum stress cri-terion, the strength of the random composite was given in terms ofuni-axial strength of the unidirectional composite through a sim-ple relation. Halpin and Kardos [27] modeled the random fibercomposites as a quasi-isotropic laminate consisting of (0�/90�/±45�)s plies. A maximum strain failure criterion was consideredto predict the ultimate strength. They provided several examplesillustrating the use of the model. Both of these studies treatedChSM as a balanced laminate layup, which is still in use nowadays[11].

Although various models to predict the strength were provided,for most cases, however, strength properties are obtained throughcoupon testing. ASTM specifies distinct methods for compression[10], tension [9], and shear tests [8], where compression test re-quires further consideration. It is noted that, unlike other materi-als, such as concrete, compressive strength of compositematerials is more difficult to measure due to brooming at bothends, causing premature failure, and thus, the result may not rep-resent the actual compressive strength. Therefore, a lot of efforthas been devoted to develop appropriate compression test fixturesin order to provide proper boundary conditions. ASTM specifiestwo methods for compression test [10]: for specimens thinner than3.2 mm, a support jig is recommended to prevent buckling; and forthose thicker than 3.2 mm, the specimen can be tested without anysupport, which applies to this study. Apparently, this method can-not avoid brooming. To solve this problem, Barbero et al. [12]developed a novel fixture. Using this fixture, splitting at the endof the sample was prevented while reducing stress concentrationat the ends, yielding compression failures at the center of speci-men. The custom made compression fixture has been used success-fully to determine compressive strength of composites [31,40] andis also adopted herein.

2.2. Out-of-plane compression for sandwich panels

One of the common failure modes for sandwich structures un-der out-of-plane compression is core crushing. Theotokoglou [38]offered an analytical determination of the ultimate strength ofsandwich beams considering the core failure in compression, ten-sion and shear using maximum failure strength method. He alsoperformed a pull-out test to verify his model. However, his studyonly gave an indication of the failure modes that took place in aT-joint under pull-out load and further research was required inorder to predict accurately the failure modes. Cvitkovich and Jack-son [22] studied the compression failure mechanisms in compositesandwich structures. The specimens they studied included thosewith no damage, specimens with 6.4 mm diameter hole, and alsospecimens with three levels of impact damage. Mouritz and Thom-son [32] investigated the compression, flexure and shear propertiesof a sandwich composite containing defects. They concluded thatdetermining the compressive properties of a large sandwich struc-ture was difficult because the strength and failure mechanismwere dependent on size effect, but core crushing under compres-sion was reported in all of their tests.

Another issue of concern for honeycomb cores is compressionbuckling. For FRP sandwich panels used for bridge deck applica-tions, the following distinct features characterize them from theircounterparts in other fields. They have relatively larger and spar-sely distributed honeycomb cells, and the core and facesheets aremanufactured separately and subsequently connected by contactbonding, using a ChSM layer and polymer resin at the interface.

A. Chen, J.F. Davalos / Composite Structures 92 (2010) 1561–1573 1563

Therefore, due to the relatively low material stiffness and thin-walled sectional geometries of sandwich components, two possibleinstability problems for sandwich panels may result under differ-ent compression loading conditions. Specifically, one is the wrin-kling of the facesheet under in-plane compression [33], and theother is the instability of the core due to out-of-plane compression[43]. The buckling of honeycomb core becomes more significantdue to the sparsely distributed thin-wall core panels. As reportedby Kumar et al. [29], local buckling of the thin walls precipitatedmost failure modes in their bending tests of tube bridge decks.Zhang and Ashby [43] also concluded that two possible failuremodes for out-of-plane compression were buckling and materialcrushing, or pure compression failure. In their study to predictbuckling strength, they assumed that the two edges of the corewall perpendicular to the loading direction were simply-sup-ported, while the other two loaded edges as rigidly constrained.Davalos and Chen [23] pointed out that the actual case at theinterface should be partially constrained, and this partial con-straint offered by the interface bond has a significant effect onthe behavior of FRP sandwich panels, that was accounted for byintroducing a coefficient of elastic restraint to quantify the bondinglayer effect. They also provided an analytical solution for compres-sion buckling capacity of a plate with two loaded edges partiallyconstrained.

Thus, it is necessary to investigate experimentally the afore-mentioned two failure modes of core elements. On that account,two compression tests were carried out, bare and stabilized com-pression tests. The bare compression test is more representativeof actual patch loading conditions. The stabilized compression testis intended to minimize buckling effect and induce primarily com-pression failure.

2.3. Transverse shear

It is commonly believed that two failure modes may occur fora sandwich panel under transverse shear: shear crushing [2,41]and shear buckling [36,34]. The concept of shear crushing isstraightforward. Allen and Feng [3] defined three categories ofsandwich panels: (1) composite beam theory (CBT), where thesandwich is treated as an ordinary composite beam and there isno shear deformation; (2) elementary sandwich theory (EST),where stresses and deflections are calculated by composite beamtheory, but there is an additional shear deflection associated withshear strains in the core; and (3) advanced sandwich theory (AST),where the faces must bend locally in order to follow the sheardeformation of the core. Most of the sandwich panels, includingthe FRP sandwich panel in this study, fall into the category ofEST. One basic assumption used for EST is that the core resiststhe shear force and the facesheet carries the membrane forcescaused by the bending moment [2,41]. It is shown [17] that ifthe Young’s modulus of the core is negligible with respect tothe facing elastic modulus, and the facing thickness is small com-pared to the height of the core, the shear stress field in the core ispractically uniform. Therefore, it is reasonable to assume thatonce this uniform shear stress exceeds the material shearstrength, the panel will fail due to shear crushing.

The research on shear buckling problem has a relatively longhistory. Bleich [15] firstly studied the shear buckling strength ofmetal structures. Timoshenko and Gere [39] refined this theoryand studied buckling of rectangular plates under action of shearstresses. Barbero and Raftoyiannis [13] used the first variation ofthe total potential energy equation to study the shear buckling ofFRP structures. Qiao et al. [36] further applied this theory to studythe local buckling of webs under shear loading. More recently,Papadopoulos and Kassapoglou [34] developed a method basedon a polynomial expansion of the out-of-plane displacement of a

plate and energy minimization, and they studied the shear buck-ling of rectangular composites plates with two concentric layups.

Recently Chen and Davalos [20] pointed out that the skin-effectcan significantly affect interfacial stress distribution, yielding acoupled stress state, where the normal stress may even be largerthan the shear stress [18,19]. They concluded that, unlike the com-mon belief that only shear stress occurs when the structure is un-der pure shear force, tensile force at the interface arises for thesinusoidal panel in the sandwich core as shown in Fig. 1, makingsuch locations critical for debonding. Thus, debonding may occurwell before shear crushing or buckling is achieved.

Therefore, it is advantageous to further experimentally investi-gate the behavior of the present sandwich core under transverseshear to correlate results with existing models. To study the shearbehavior of the sandwich core, ASTM [4] specifies a testing meth-od. However, this method cannot be directly applied to this studysince the core is very strong in shear. Trial tests using this methodillustrated that the failure is intra-laminar delamination, instead ofpure shear failure of the core. Another method based on four-pointbending test is also recommended by ASTM [7] to study shearstrength and shear stiffness of FRP sandwich cores, since relativelypure shear and bending regions will result from this loading condi-tion. Many researchers have performed bending tests on sandwichbeams. Lingaiah and Suryanarayana [30] carried out experimentalvs. analytical correlations of the mechanical properties of sand-wich-beam specimens. Four-point and three-point loading testswere conducted. It was observed that generally the failure loadwas higher for the case of the four-point bending test than forthe three-point bending test. The failure of most specimens wasdue to debonding between the core and the facing, and at loadswhich were less than estimated based either on core shear strengthor facing tensile/compressive strength, whichever was lowerdepending on the test condition. But they did not specify the loca-tion where the debonding initiated and did not examine rigorouslythe failure mechanism. Sotiropoulos et al. [37] analyzed the struc-tural efficiency of pultruded FRP components and systems in termsof joint efficiency, transverse load direction, composite action be-tween FRP components, and maximum deflections and strainsthrough three- and four-point bending tests. Brown et al. [16] per-formed three- and four-point bending tests on both laid up andpultruded triangular sections and used the results to improve de-signs and manufacturing approaches. Mouritz and Thomson [32]carried out four-point bending tests to study shear properties ofa sandwich composite containing interfacial cracks subjected toimpact load. GangaRao et al. [26] conducted static and fatiguebending tests on a 2.74 m, simply supported FRP deck specimenunder patch load. They compared the test results with the require-ments from Ohio DOT in terms of deflection, flexure and shear, andconcluded that deflection limit state governed the design of thedeck sample. Caprino and Langella [17] performed three-pointbending tests on a sandwich beam for shear characterization ofthe foam core, by inserting rigid blocks near the location of theconcentrated load. Kumar et al. [29] tested several coupon samplesconsisting of single, double, and a four-layered tube assemblies un-der static flexure loading up to failure. They observed that web-flange junction was the critical location of failure. Abbadia et al.[1] carried out four-point bending tests, analytical and numericalanalyses on a honeycomb sandwich panel, based on which theyproposed a general kinematic model. Belouettar et al. [14] usedfour-point bending test to investigate static and fatigue behaviorof honeycomb sandwich composites, where the fatigue resultswere presented in S–N diagrams. Global and local parameters wereconsidered to evaluate the fatigue life of the sandwich composites;and effects of core densities and cell orientation on the maximumload and the damage processes were discussed. Thus, the literatureindicates that the four-point bending test method is an effective

Table 1Properties of ChSM core material.

Type Nominal weight,g/m2 (oz/ft2)

Thickness (mm) Fiber volume fraction

M127 ChSM 3661.8 (12.0) 7.6 0.1877

Fig. 2. Compression fixture and failed specimen under compression.

Table 2Strength of ChSM core material.

Property Compression Shear Tension

Strength (MPa) 148.1 70.6 127.6Standard deviation (MPa) 4.0 2.8 –Range (MPa) – – 127.3–128.0

0

20

40

60

80

100

120

140

160

0 0.5 1 1.5 2Displacement (mm)

Stre

ss (M

Pa)

Fig. 3. Stress–displacement path under compression.

0

20

40

60

80

100

120

140

0 5000 10000 15000 20000

Micro Strain

Str

ess

(MP

a)

Fig. 4. Stress–strain path under compression.


approach to study the shear properties and strength of sandwichcore structures, and is also adopted in this study.

3. Strength evalution of CHSM

ChSM is the material used for the core of the present FRP sand-wich panels (Fig. 1), and basic properties of ChSM are given inTable 1.

3.1. Compressive strength

The custom made compression fixture was used to determinethe compressive strength (Fig. 2a). Each half of the fixture consistsof two stacked square plates (12.7 � 12.7 cm). Each of the innerplates, on top and bottom, has a centered rectangular cut-out toposition the specimen inside, supported around the specimenedges by adjustable shims having tapered ends in contact withthe base of the specimen. The shims are pressed against the spec-imen by side screws, to center the specimen, and more importantlyprovide relative restraint around the edges to avoid brooming atthe ends. The top plates are slided into position through four guid-ing posts, to maintain perfectly parallel surfaces over the specimenends, and avoid misalignment when the load is applied by the slid-ing stacked top plate assembly. Detailed description of the fixturecan be found in Makkapati [31] and Chen [18].

The specimens were cut into the dimension of 50.8 � 25.4 mm,and they were tested in a Baldwin Universal Testing machine.LVDTs were used to measure the movement of the top loadingplate assembly, and strain gages that were bonded the mid-heightof the specimen were used to measure the compressive strain.

The average strength and standard deviation for five specimensis given in Table 2. It shows that the experimental results obtainedwere fairly consistent, and the standard deviation was about 4%.During the test, the specimens remained intact until the maximumload was reached, and the failure was characterized by a loudsound and a sudden drop of the load. Figs. 3 and 4 provide typicalgraphs of stress vs. displacement and stress vs. strain at mid-span,respectively, showing nearly linear relations up to failure. It isnoted that the displacement measurement corresponded to themovement of the loading plate, and the strain was recorded up

y = 0.0092x - 0.8707R2 = 0.9996

0

10

20

30

40

50

60

0 1000 2000 3000 4000 5000 6000 7000Micro Strain

Stre

ss (M

pa)

Fig. 5. Stress–strain curve to obtain shear stiffness.

L

L/2

90° (typ)

45° (typ)

r (typ)

w d1

d2

h

ediStnorF

Nominal Specimen Dimensions:

d1=19.1 mm d2=3.8 mm h=as required L=76.2 mm r=1.3 mm

Fig. 6. Shear specimen dimensions.

Upper Grip with Linear Bearing

Test Machine Adaptor

Baseplate

Lower Grip

Lower Grip Holder

Bearing Post

Adjacent Jaws Tightened by Thumbscrews

Specimen

P

Fig. 7. Iosipescu test setup.


to the capacity of the strain gage of 1.5% in this study. A typical fail-ure mode is shown in Fig. 2b illustrating the compressive failureachieved.

Based on the strain and stress data collected from the tests, thestiffness was obtained by fitting the data between the range of

Table 3Stiffness of ChSM core material.

Property Compression Shear Tension

Stiffness (MPa) 8604 3792 6471Range (MPa) 8052–9156 3718–3866 6303–6618

1000 and 6000 micro strain, as shown in Fig. 5, and the result is gi-ven in Table 3.

3.2. Shear strength

The Iosipescu V-notched test was used to evaluate the shearstrength and stiffness. The dimensions of the specimen are givenin Fig. 6. The sketch of the test setup is shown in Fig. 7, and a photois given in Fig. 8a. All tests were carried out in an MTS machine. Ashear strain gage was bonded between the two notches, and thedisplacement and load were recorded using the internal displace-ment transducer and load cell. The loading rate was controlled at1.3 mm/min.

Figs. 9 and 10 display stress vs. displacement and stress vs.strain curves, respectively, and typical shear failure mode is illus-trated in Fig. 8b. The same method as described in the previoussection was used to compute the strength and stiffness. The resultsof five specimens for strength and two specimens for stiffness aregiven in Tables 2 and 3, respectively.

3.3. Tensile strength

The dimensions of the specimen for tension test are given inFig. 11a [9]. Tabs were bonded at both ends to prevent crushingdamage caused by the grips. All tests were carried out in a Baldwinmachine. Strain gages were bonded at mid-span on both sides, andthe displacement and load were recorded respectively using LVDTsand a load cell.

Figs. 12 and 13 display stress vs. displacement and stress vs.strain curves, respectively, and typical tension failure mode is illus-trated in Fig. 11b. The strength and stiffness were computed usingthe same method as for other tests described above. Due to ex-pected close results and time constraints, the results for only twospecimens were obtained for strength and stiffness, are givenrespectively in Tables 2 and 3.

4. Out-of-plane compression test for single-cell samples

To study the out-of-plane compression behavior of sandwichpanels, two types of tests were performed on single-cell sandwichspecimens: the stabilized test to achieve compression failure, andthe bare test to induce compression buckling failure, as discussedabove.

4.1. Naming convention

In this study, the naming convention is defined in Fig. 14, wherethe letters B and C represent for ChSM, Bonding layer numbers andCore thickness, respectively. The integers i and j (i = j = 1, 2, 3) cor-respond to their respective nominal weights of ChSM layer used.

4.2. Test description

The specimen was a typical single-cell cut from the sandwichpanel, and it was representative of a symmetric volume elementof the structure when subjected to compression. The schematic isshown in Fig. 15a, with dimensions of 102 mm by 102 mm squareand 51 mm deep. To primarily evaluate the effect of bonding layersand minimize the influence of additional layers of the facesheet,only one external fabric was added, consisting of E-glass longitudi-nal fibers (0� roving) and continuous strand mat (ContSM), asshown in Fig. 15b. The thickness of ChSM bonding layers was var-ied from one layer to three layers (Bi, Fig. 14), and the core thick-ness was varied from one to two thicknesses for different type ofspecimens (Cj, Fig. 14). From a previous study by Davalos et al.

0

10

20

30

40

50

60

0 5000 10000 15000 20000Micro Strain

Stre

ss (M

Pa)

Fig. 10. Stress–strain path under shear.

(a) Tension specimen dimensions

(b) Failed specimen under tension

25.4 mm

R=31.8mm

12.7 mm

mm2.67mm2.67

254 mm

Fig. 11. Tension specimen dimensions and failed specimen.

0

20

40

60

80

100

120

140

0 0.5 1 1.5 2 2.5Displacement (in.)

Stre

ss (M

Pa)

Fig. 12. Stress–displacement path under tension.

Fig. 8. Photos of Iosipescu test setup and failed specimen under shear.

0

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40

50

60

70

80

0 0.05 0.1 0.15 0.2 0.25Displacement (mm)

Stre

ss (M

pa)

Fig. 9. Stress–displacement path under shear.


[24], the properties of the constituent materials are given in Table4, and the properties of each component material are provided inTable 5.

Two cases of compression tests were carried out. For the baretest, an elastomeric pad was placed between the loading blockand the single-cell specimen. This test is representative of activeloading condition and induces primarily buckling failure. For thestabilized test, the specimen was bonded to top and bottom steelplates, and the load was applied directly over the top steel plate.This test is intended to minimize buckling effect and induce pri-marily compression failure.

All tests were carried out according to ASTM standards [6]. Theywere performed in a universal testing machine with a 200,000 lbcapacity. An external load cell was used to record the load, andLVDTs were used to record the displacements. Four strain gageswere bonded at the mid-height of the core to obtain compressivestrains, two on the sinusoidal wave panel and two on the side flatpanel. The load was controlled at such a rate that the failure oc-curred within 3–6 min.

4.3. Test results and discussion

4.3.1. Bare compression testWhen the load was applied to the single-cell specimen, both of

the core side flat panels (see Fig. 16a) bent outwards, and this

0

20

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60

80

100

120

0 5000 10000 15000 20000

Micro Strain (in.)

Stre

ss (M

Pa)

Fig. 13. Stress–strain path under tension.

B i C j

Bonding layer Core thickness

1-915.5 g/m2

2-1373.2 g/m2

3-1830.9 g/m2j=

i= 1-915.5 g/m2

2-1830.9 g/m2

3-2746.4 g/m2

Fig. 14. Naming convention.

(a) Single-cell core specimen dimensions

(b) Layup of facesheet

t

Sinusoidal panel

Flat panel

t

t

t

102 mm

102 mm

UM –1810 (0 roving + ContSM)

Bonding layer (ChSM)

Fig. 15. In-plane single-cell core specimen dimensions and layup of facesheet.

Table 4Properties of constituent materials [24].

Material E (GPa) G (GPa) m q (g/cm3)

E-glass fiber 72.4 28.8 0.255 2.55Polyester resin 5.06 1.63 0.3 1.14

Table 5Layer properties of face laminate and core materials [24].

Ply name Ply type Nominal weight(g/m2)

Thickness (mm) Vf

Bonding layer ChSM 915.5 1.91 0.1877UM1810 0� 610.3 0.64 0.3774

ContSM 305.2 0.34 0.3582


deformation could be interpreted as a geometric imperfection. Asthe load increases, the specimens with distinct bonding layers dis-played different behaviors. For B1C2 (one bonding layer), the sidepanels buckled and delaminated from the specimen well before

ultimate failure occurred. While for the other types, the side panelsdid not delaminate. For all specimen types, upon sudden crushingof the side panel, the specimen did not fail immediately but contin-ued to carry load for several event failures, until collapse of thespecimen. A typical failure mode is shown in Fig. 16a.

The average value and standard deviation for six samples eachare given in Table 6, which shows that the magnitudes of failureloads are in the same order as the number of bonding layers andcore thickness; i.e., the specimen with three bonding layers ismuch stronger than that with one bonding layer, and the specimenwith two core thickness is stronger than that with one core thick-ness, clearly showing that the bonding layer and core thicknessplay an important role on the failure load. Fig. 17 shows theload–displacement curves. Fig. 18 shows the transverse strain vs.load curves for the core sinusoidal panel. As the elastomeric padwas placed between the loading block and the specimen, this dis-placement does not represent the actual deformation of the speci-men. However, from these figures we can conclude that thespecimen exhibits an approximate linear behavior up to failure.

4.3.2. Stabilized compression testIn this test, all three types of samples showed the same failure

mode. They all failed by crushing of the core panels. The sinusoidalwave panel failed first, followed by the crushing of the remainingcomponents of the core. The failure mode is shown in Fig. 16b.No apparent damage was observed prior to ultimate failure.

The failure loads for three specimens each are given in Table 7,which shows much higher values compared with what we ob-tained for the bare compression tests. Fig. 19 shows typical load–displacement curves for the inside sinusoidal wave panels, andFig. 20 displays the strain–load curves. Again we can see that thespecimens follow a nearly linear behavior until failure occurs.

4.4. Discussion of experimental results

Barbero et al. [12] reported in his study that the compressivestrength for ChSM was 153.06 MPa. From the present test resultson coupon samples, the compressive strength was found to be148.1 MPa. Halpin and Kardos [27] suggested a model to predictthe compressive strength for ChSM using a pseudo-isotropic lami-nation approach. Following his method, if the compression failurestrains for the equivalent unidirectional composite are estimatedto be e1c = 0.015 and e2c = 0.006, the stress–strain curve to failurecan be predicted as given in Fig. 21, with ultimate strength of about160 MPa. Thus, strength for ChSM of fc = 148.1 MPa in this studyseems reasonable. The total in-plane area of the core walls, eachwith two core thickness, is Ac = 13.74 cm2. Then, the nominal mate-rial failure load can be calculated as:

Fc ¼ fc � Ac ¼ 203:49 kN ð1Þ

The stabilized test provided failure loads ranging from 148.7 kNto 201.2 kN. If the unevenly distributed load effect is considered,we can conclude that the stabilized compression test results in atypical compression failure. For the bare test, the failure load wasmuch lower than the nominal ultimate compressive load. A previ-ous study indicated that the predicted core buckling load was

Fig. 16. Failed specimens under compression test.

Table 6Failure loads for bare compression tests.

B1C2 B2C2 B3C2 B3C1

Average value (kN) 74.60 93.46 101.86 31.74Standard deviation (kN) 3.89 8.47 9.43 3.45

0

20

40

60

80

100

120

0 2 4 6 8Displacement (mm)

Load

(kN

)

B1C2

B2C2

B3C2

B3C1

Fig. 17. Load–displacement curve for bare compression test.

0

2000

4000

6000

8000

10000

12000

0 20 40 60 80 100

Load (kN)

Mic

ro S

trai

n

B1C2

B2C2

B3C2

B3C1

Fig. 18. Strain–load curve for bare compression test.

Table 7Failure loads for stabilized compression tests.

B1C2 B2C2 B3C2

Average value (kN) 155.54 163.07 177.22Range (kN) 148.73–162.36 156.71–171.83 158.13–201.59

0

20

40

60

80

100

120

140

160

180

200

0 1 2 3 4 5 6 7 8Displacement (mm)

Load

(kN

) B1C2

B2C2

B3C2

Fig. 19. Load–displacement curve for stabilized compression test.

0

2000

4000

6000

8000

10000

12000

0 50 100 150 200Load (kN)

Mic

ro s

train

B1C2

B2C2

B3C2

Fig. 20. Strain–load curve for stabilized compression test.


145.6 kN for clamped condition and 50.0 kN for hinged conditionat the core–facesheet interface [23,18]. The failure loads obtained

for bare compression tests fall within this range, with values of102 kN for B3C2 to 75 kN for B1C2. These results indicate thatthe actual core–facesheet connection lies between simply sup-ported and fully-restrained conditions, and local buckling occursbefore maximum material compressive strength is reached. Once

0

20

40

60

80

100

120

140

160

180

0 5000 10000 15000 20000Micro strain

Stre

ss (M

Pa)

Fig. 21. Stress–strain curve for ChSM.

L

s

b

Strain gagesFacesheetFRP Core

d

P/2 P/2

Fig. 22. Bending test setup.

Fig. 23. Failed specimens under bending test.


local buckling occurs, the buckled panels of the core lose theirfunction and the compressive load is redistributed among theother parts. Finally the structure fails in compression or a combina-tion of bending and compression. As expected, the failure loads ofthe stabilized compression tests are much higher than those for thebare compression tests.

Thus, the two types of tests resulted in two distinct failuremodes. Buckling occurred for the bare compression test, whilethe stabilized compression test induced material compression fail-ure. As a matter of interest, these two failure modes were the sameas those reported by Zhang and Ashby [43] under out-of-planecompression.

5. Four-point bending tests for beam type samples

As discussed earlier, the standard method of [4] used for evalu-ating shear properties of sandwich core materials was not suitablefor this study. In trial tests, due to the high shear strength of thecore material, failure occurred as intra-laminar delamination with-in the facesheet, well before the shear failure of the core materialcan be achieved. Thus, the four-point bending test [7] was adoptedin this study to evaluate core shear strength and modulus.

5.1. Test description

The dimensions of the specimen were 71.1 cm long by 10.2 cmwide by 5.1 cm deep. There were seven single-cells (see Fig. 15a)either along the longitudinal or transverse directions. As men-tioned before, to minimize the influence of the facesheet layers,other than the effect of the bonding layers on the strength of thespecimen, only a combined 0�/ContSM layer was placed overthe ChSM bonding layer, as shown in Fig. 15b. The properties ofthe constituent materials of the facesheet (Fig. 15b) are providedin Table 4, and the properties of each component material aregiven in Table 5.

The core of the sandwich panels was ‘‘embedded” into the face-sheet ChSM/resin contact layer. The number of these bonding lay-ers and the core thickness were varied to study their effect onshear strength. Two types of beam samples were manufacturedby orienting the sinusoidal wave: (1) along the length (longitudi-nal), and (2) along the width (transverse). All tests were carriedout in accordance to ASTM standards [7]. Fig. 22 displays the testsetup, where L = 61.0 cm and s = 30.5 cm. An external load cellwas placed between the loading block and the specimen to recordthe load, and LVDTs were used to record the displacements. Twostrain gages on the top and two on the bottom facesheets were

bonded at the mid-span of the beam. The test was performed ata displacement rate of 1.5 mm/min.

5.2. Test results and discussion

5.2.1. Longitudinal testThe beams under static loadings showed nearly linear-elastic

behavior up to failure. The number of bonding layers affected theload capacity of the specimens. For specimens with one–threebonding layers, the failure was due to sudden debonding betweenthe facesheet and the core material, as shown in Fig. 23a, initiatingtypically at one end of the compression side. The energy stored inthe specimen was released in a relatively short time resulting in aloud failure. In order to avoid debonding and achieve primarilyshear failure of core materials, several B3C2 specimens were spe-cially treated during the manufacturing process, where rich resincontent was allowed to cure at the core–facesheet interface. As ex-pected, for the B3C2 specimens with excessive resin content, thefacesheet did not delaminate from the core, and instead a typicalshear failure of the core under the loading point occurred, asshown in Fig. 23b.

The average values of the maximum load for three specimenswith excessive resin and five specimens for each of the other types

Table 8Failure loads for longitudinal samples.

B1C2 B2C2 B3C2 B2C1 B2C3 B3C1 B3C3 B3C2 with excessive resign

Average value (kN) 16.68 24.18 30.16 17.64 23.51 23.62 41.41 70.46Standard deviation (kN) 0.89 2.18 3.74 1.67 3.56 1.82 3.29 17.73


are given in Table 8, which shows that the magnitudes of failureloads are in the same relation as the number of bonding layersand core thickness; i.e., the specimen with three bonding layersis much stronger than that with one bonding layer, and the speci-men with three core thickness is much stronger than that with onecore thickness, clearly showing that the effect of number of bond-ing layers and core thickness plays an important role on the failureload. As the number of bonding layers increases, the core embed-ment increases, and as the thickness increases the contact bondingarea increases. Fig. 24a and b shows the displacement at mid-spanvs. load curves, respectively, for constant two bonding layers andthree core thickness, and constant two core thickness. Fig. 25aand b shows the corresponding load–strain curves for the same

0

5

10

15

20

25

30

35

40

45

50

0 2 4 6 8 10 12 14Displacement (mm)

Load

(kN

)Lo

ad (k

N)

(a) Two bonding layers

0

10

20

30

40

50

60

151050Displacement (mm)

(b) Two core thickness

B2C2

B2C1

B2C3

B2C2

B3C2

B1C2

Fig. 24. Load–displacement curves for longitudinal tests.

Table 9Failure loads for transverse samples.

B1C2 B2C2 B3C2

Average value (kN) 5.25 6.98 11.19Standard deviation (kN) 0.71 1.27 2.67

two conditions as Fig. 24a and b. From these figures we can con-clude that specimens exhibited an approximate linear behaviorup to failure, and the specimens with higher number of bondinglayers and larger core thickness exhibited higher stiffness.

5.2.2. Transverse testAll types of specimens displayed the same failure mode. The

failure in the core was initiated by debonding at the contact areabetween the sinusoidal panel and flat panel, as shown in Fig. 26.The specimens continued to carry some load until the delamina-tion between the facesheet and core material occurred. Unlike lon-gitudinal specimens, the failure was not as sudden, and severalrises and drops of load were observed during the test.

B2C1 B2C3 B3C1 B3C3

5.29 7.16 6.78 11.030.33 0.76 0.87 1.82

0

5

10

15

20

2530

35

40

45

50

0 2000 4000 6000 8000

Micro strain

Load

(kN

)Lo

ad (k

N)


05

10

1520253035

404550

0 2000 4000 6000 8000 10000

Micro strain(b) Two core thickness

B2C2

B2C1

B2C3

B2C2B3C2

B1C2

Fig. 25. Load–strain curves for longitudinal tests.

Fig. 26. Core separation.

0

2

4

6

8

10

12

14

0 500 1000 1500 2000 2500

Micro strain

Lo

ad (

kN)


B2C2

B2C1

B2C3

10

15

20

25

Lo

ad (

kN)

B3C2


The failure loads for five specimens each are given in Table 9,which shows much lower values compared with what we obtainedfor the longitudinal samples. Fig. 27a and b shows typical load–dis-placement curves, respectively, for constant two bonding layersand constant two core thickness; and the corresponding typicalload–strain curves are shown in Fig. 28a and b. We can observe

0

2

4

6

8

10

12

14

0 5 10 15 20

Displacement (in.)

Load

(kN

)


B2C2

B2C1

B2C3

0

5

10

15

20

25

0 5 10 15 20 25 30 35Displacement (in)

Load

(kN

)


B2C2

B3C2

B1C2

Fig. 27. Load–displacement curves for transverse tests.

0

5

0 500 1000 1500 2000 2500 3000Micro strain


B2C2

B1C2

Fig. 28. Load–strain curves for transverse tests.

that the specimens followed a nearly linear behavior until first fail-ure occurred.

5.2.3. DiscussionFor longitudinal specimens with excessive resin, if we assume

that the shear force is carried by the core only and the shear stressis uniformly distributed, we have

c ¼ P2Gxzbðdþ tf Þ

ð2Þ

where c is the shear strain, b and d, are the width and height of thebeam, tf is the thickness of the facesheet, as given in Table 10, andGxz is the equivalent shear modulus, which is given by a previousstudy [20,18] as:

Gxz ¼ 318:06 MPa ð3Þ

Since the shear strain of the core flat panel should be the sameas the global shear strain, the shear stress in the flat panel can becalculated as:

s ¼ cG12 ð4Þ

where G12 is the material shear modulus given in Table 3, which is3.79 GPa. We can substitute s = 70.6 MPa from Table 2 into Eq. (4)

Table 10Parameters for sandwich beam specimen with excessive resin.

b (mm) d (mm) tf (mm)

102 51 7.2


to back calculate c, and then use Eq. (2) to estimate the failure loadas P = 69.81 kN. This value is in good correlation with the averageexperimental load of P = 70.46 kN.

Chen and Davalos [20] pointed out that tensile force arises atthe interface for the sinusoidal core panel when under pure shearforce, causing delamination. It was observed in this study thatdelamination of the top facesheet occurred first within shear load-ing region in bending. Thus, using a combination of Flatwise Ten-sion Test [5] of interfaces and the failure loads obtained in thisstudy, practical formulas can be provided to predict the onset ofthe delamination in a follow up study.

6. Conclusions

An experimental investigation on strength properties of sinu-soidal honeycomb core FRP specimens under out-of-plane com-pression and bending is presented. In particular, the influence offacesheet–core interface bonding effect and core thickness isexamined by varying the interface bonding layers of the specimenand core thicknesses. Two cases of compression tests are carriedout, the stabilized compression test and the bare compression test.Two cases of bending tests are carried out: longitudinal and trans-verse bending tests. From this study, the following conclusions canbe drawn:

(1) The material strength of the ChSM can be obtained throughthe coupon tests as shown in this study.

(2) For compression tests of single-cell sandwich specimens, allresponses followed an approximate linear behavior prior tofailure in compression. The failure load for the stabilizedcompression test was much higher than that of the barecompression test. In the stabilized compression test, thespecimens failed in compression, and in the bare compres-sion test, local buckling occurred before the maximum com-pression strength was reached.

(3) In the bare compression test, the failure load was sensitiveto the bonding layer effect. Specimens with three bondinglayers failed at a higher load than those with one bondinglayer. While in the stabilized compression test, the failureloads were not affected much by the number of bonding lay-ers. In both tests, specimens with larger core thickness failedat higher loads than those with lesser core thickness.

(4) All specimens followed an approximate linear behavior priorto failure in bending. The failure load for the longitudinalspecimens was much higher than that for the transversespecimens. For longitudinal samples, those with excessiveresin within the bonding layers failed in shear, and the otherspecimen types failed by debonding. All of the transversespecimens failed by core panel separation. Thus, as it is thecase in practice, transverse-type beams should be avoidedwhen high shear stresses are expected.

(5) The longitudinal samples were much stronger in shear thanthe transverse samples. The number of bonding layers andcore thickness corresponded clearly to the maximumstrengths achieved. However, there is variability in resultseven for specimens with the same number of bonding layers,especially for the type with excessive resin within the bond-ing layer. One of the factors that may contribute to this var-iability is the bonding quality. For some specimens, theresin–fillets are not well formed at the core–facesheet inter-face, resulting in minor cracks. This indicates the importanceof quality control during the manufacturing process.

All of the test data given in this study will be subsequently usedtogether with analytical models from the authors [18,23,20] to

develop design guidelines [21] for the light-weight honeycombsandwich panel as shown in Fig. 1, intended primarily for highwaybridge applications. The methods described in this paper can be ex-tended to study other types of sandwich panels.

Acknowledgement

We gratefully acknowledge financial support from the NationalScience Foundation (NSF) Partnerships for Innovation program andthe West Virginia University Research Corporation. The samples forthis study and technical advice were generously provided by Dr.Jerry Plunkett of Kansas Structural Composites, Inc., Russell, Kan-sas, USA.

References

[1] Abbadia A, Koutsawaa Y, Carmasolb A, Belouettara S, Azarib Z. Experimentaland numerical characterization of honeycomb sandwich composite panels.Simulat Model Pract Theory 2009;17(10):1533–47.

[2] Allen HG. Analysis and design of structural sandwichpanels. Oxford: Pergamon Press; 1969.

[3] Allen HG, Feng Zhengnong. Mechanics of sandwichstructures. Netherlands: Kluwer Academic Publisher; 1998. p. 1–12.

[4] ASTM designation: C273-00, Standard test method for shear properties ofsandwich core materials, ASTM, PA, USA; 2000.

[5] ASTM designation: C297-94, Standard test method for flatwise tensile strengthof sandwich constructions, ASTM, PA, USA; 2000.

[6] ASTM designation: C365-00, Standard test method for flatwise compressionproperties of sandwich cores, ASTM, PA, USA; 2000.

[7] ASTM designation: C393-00, Standard test method for flexure properties ofsandwich constructions, ASTM, PA, USA; 2000.

[8] ASTM designation: D5379-98, Standard test method for shear properties ofcomposite materials by the V-notched beam method, ASTM, PA, USA; 2000.

[9] ASTM designation: D638-99, Standard test method for tensile properties ofplastics, ASTM, PA, USA; 2000.

[10] ASTM designation: D695-96, Standard test method for compressive propertiesof rigid plastics, ASTM, PA, USA; 2000.

[11] Barbero EJ. Introduction to composite materials design. Philadelphia(USA): Taylor & Francis; 1999.

[12] Barbero EJ, Makkapati S, Tomblin JS. Experimental determination ofcompressive strength of pultruded structural shapes. Compos Sci Technol1999;59:2047–54.

[13] Barbero EJ, Raftoyiannis IG. Local buckling of FRP beams and columns. J MaterCivil Eng ASCE 1993;5(3):339–55.

[14] Belouettar S, Abbadi A, Azari Z, Belouettar R, Freres P. Experimentalinvestigation of static and fatigue behaviour of composites honeycombmaterials using four point bending tests. Compos Struct 2009;87(3):265–73.

[15] Bleich F. Buckling strength of metal structures. New York (USA): McGraw-Hill;1952.

[16] Brown Richard T, Zureick Abdul-Hamid. Truss section fiberglass reinforcedcomposite causeway deck. In: Proceedings of 30th international SAMPEtechnical conference; October 20–24 1998. p. 303–15.

[17] Caprino G, Langella A. Study of a three-point bending specimen for shearcharacterisation of sandwich cores. J Compos Mater 2000;34(9):791–814.

[18] Chen A. Strength evaluation of honeycomb FRP sandwich panels withsinusoidal core geometry, Ph.D. Dissertation, Department of Civil andEnvironmental Engineering, West Virginia University, Morgantown, WV,USA; 2004.

[19] Chen A, Davalos JF. A solution including skin effect for stiffness and stress fieldof sandwich honeycomb core. Int J Solids Struct 2005;42(9–10):2711–39.

[20] Chen A, Davalos JF. Transverse shear with skin effect for composite sandwichwith honeycomb sinusoidal core. J Eng Mech-ASCE 2007;133(3):247–56.

[21] Chen A, Davalos JF. Design guidelines for honeycomb FRP sandwich bridgepanel. Compos Struct, submitted for publication.

[22] Cvitkovich MK, Jackson WD. Compression failure mechanisms in compositesandwich structures. In: American helicopter society 54th annual forum,Washington, DC; May 20–22, 1998.

[23] Davalos JF, Chen A. Buckling behavior of honeycomb FRP core with partiallyrestrained loaded edges under out-of-plane compression. J Compos Mater2005;39(16):1465–85.

[24] Davalos JF, Qiao P, Xu XF, Robinson J, Barth KE. Modeling and characterizationof fiber-reinforced plastic honeycomb sandwich panels for highway bridgeapplications. Compos Struct 2001;52:441–52.

[25] DeTeresa SJ, Freeman DC, Groves SE, Sanchez RE. Failure under multiaxialstresses of component materials for fiber composite sandwich construction.In: Proceedings of twelfth international conferences on composite materials(ICCM-12), Paris, France; July 5–9 1999. p. 198.

[26] GangaRao HVS, Thippeswamy HK, Shekar V, Craigo C. Development of glassfiber reinforced polymer composite bridge decks. SAMPE J 1999;35(4):12–24.


[27] Halpin JC, Kardos JL. Strength of discontinuous reinforced composites: I. Fiberreinforced composites. Polym Eng Sci 1978;18(6):496–504.

[28] Hann HT. On approximation for strength of random fiber composites. J ComposMater 1975;9:316–26.

[29] Kumar P, Chandrashekhara K, Nanni A. Testing and evaluation of componentsfor a composite bridge deck. J Reinf Plast Compos 2003;22(5):441–61.

[30] Lingaiah K, Suryanarayana BG. Strength and stiffness of sandwich beams inbending. Exp Mech 1991;3:1–7.

[31] Makkapati S. Compressive strength of pultruded structural shapes, MasterThesis, Department of Mechanical and Aerospace Engineering, West VirginiaUniversity, Morgantown, WV, USA; 1994.

[32] Mouritz AP, Thomson RS. Compression, flexure and shear properties of asandwich composite containing defects. Compos Struct 1999;44:263–78.

[33] Niu K, Talreja R. Modeling of wrinkling in sandwich panels under compression.J Eng Mech 1998;125(8):875–83.

[34] Papadopoulos L, Kassapoglou C. Shear buckling of rectangular compositeplates with two concentric layups. J Reinf Plast Compos 2004;23(1):5–16.

[35] Plunkett JD. Fiber-reinforcement polymer honeycomb short span bridge forrapid installation, IDEA Project Report; 1997.

[36] Qiao P, Davalos JF, Wang J. Local buckling of composite FRP shapes by discreteplate analysis. J Struct Eng 2001;127(3):245–55.

[37] Sotiropoulos SN, GangaRao HVS, Allison RW. Structural efficiency of pultrudedFRP bolted and adhesive connections. In: Proceedings of 49th annualconference, composite institute. Cincinnati, Ohio: The Society of PlasticsIndustry, Inc.; 1994 [SPT/Composite Institute, New York].

[38] Theotokoglou EE. Analytical determination of the ultimate strength ofsandwich beams. Appl Compos Mater 1996;3:345–53.

[39] Timoshenko SP, Gere JM. Theory of elastic stability. NY (USA): McGraw-HillBook Company, Inc.; 1961.

[40] Tomblin JS. Compressive strength models for pultruded glass fiber reinforcedcomposites, Ph.D. Dissertation, Department of Mechanical and AerospaceEngineering, West Virginia University, Morgantown, WV, USA; 1994.

[41] Vinson JR. The behavior of sandwich structures of isotropic and compositematerial. Technomic Publishing Company; 1999.

[42] Xu XF, Qiao P, Davalos JF. Transverse shear stiffness of composite honeycombcore with general configuration. J Eng Mech 2001;127(11):1144–51.

[43] Zhang J, Ashby MF. Out-of-plane properties of honeycombs. Int J Mech Sci1992;34(6):475–89.

strength evaluations of sinusoidal core for frp sandwich bridge deck panels

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