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The effect of foam core density at various slenderness ratios on axial strength of sandwich panels with glass-FRP skins
Composites Part B 106 (2016) 129e138
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Composites Part B journal homepage: www.elsevier.com/locate/compositesb
The effect of foam core density at various slenderness ratios on axial strength of sandwich panels with glass-FRP skins Luke CoDyre, Amir Fam* Department of Civil Engineering, Queen's University, Kingston, ON, K7L 3N6, Canada
a r t i c l e i n f o
a b s t r a c t
Article history: Received 24 June 2016 Received in revised form 5 September 2016 Accepted 5 September 2016 Available online 7 September 2016
In this study, the effect of the density of a polyisocyanurate (PIR) foam core on the behavior of axially loaded sandwich panels with glass fibre reinforced polymer (GFRP) skins are examined, for different panel heights. Thirty-six test specimens with a cross-sectional foam core area of 100 50 mm2 and symmetrical 100 mm wide GFRP skins were fabricated with heights of 750, 1000, 1250, and 1500 mm to examine various slenderness ratios (kLe/r). Three PIR foam core materials with densities of 32, 64 and 96 kg/m3 were used. An out-of-straightness assessment was first conducted to ensure that eccentricities were within acceptable limits. All samples were tested to failure in axial compression with pinned end conditions. It was shown that doubling and tripling the core density led to increases in peak load of 71 and 170%, respectively, due to the enhanced composite action and reduced shear deformations. On the other hand, slenderness had insignificant impact on peak loads, contrary to what one would expect from Euler's theory. The reason being the variety of governing failure modes. Columns with low (kLe/r) of 22 e25 failed largely due to localized single skin buckling, while those with high (kLe/r) of 51e61 failed primarily by global buckling followed by secondary core shear failure with skin debonding. Columns with intermediate (kLe/r) experienced a combination of localized and global failure modes. © 2016 Elsevier Ltd. All rights reserved.
Keywords: Sandwich panel GFRP skin Foam core Density Slenderness Axial
1. Introduction Sandwich panels have been widely studied and used for quite some time in aerospace and other industries, including marine applications (e.g. Ref. [12]. Studies have focused on their behavior under dynamic air-blast loading [15], behavior when exposed to under-water impulsive loading [10], their fatigue behavior [5], as well as acoustic behavior [14]. Traditionally, the applications of sandwich panels in civil engineering have been limited but in recent years they started gaining attention with the rise in popularity of modular and rapid construction techniques. These composite systems, comprised of a low density insulating core material between two stiff skins, boast high strength-to-weight ratios making them highly transportable and useful as pre-fabricated wall, roof or deck elements. A variety of closed-form foams can be used as the core material due to their water impermeability and insulation properties, such as polyisocyanurate (PIR), polyvinylchloride (PVC) and polyurethane (PUR) [4]. Among the many skin material alternatives, glass fibre reinforced polymer (GFRP) has
* Corresponding author. E-mail address: [email protected] (A. Fam). http://dx.doi.org/10.1016/j.compositesb.2016.09.016 1359-8368/© 2016 Elsevier Ltd. All rights reserved.
been extensively researched due to its high specific strength and corrosion-resistance. Allen [1] and Frostig [7] have shown that the shear stiffness of concentrically loaded slender soft-core sandwich panels is significant in the determination of the critical buckling load for the section. A reduction in the typical Euler buckling capacity is seen due to reduced core shear stiffness. Additional deflections due to core shear deformations are also seen in the flexural behavior of sandwich panels, however, increasing the density of the core material has been found to reduce this effect. Sharaf et al. [16] varied the density of PUR foam cores in GFRP skinned flexural panels and observed significantly reduced shear deformations and increased ultimate capacity and stiffness. Mathieson and Fam [11] conducted a study on the concentric axial loading of polyurethane foam core and GFRP skinned sandwich panels to determine the effects of slenderness on panel behavior. Core density was kept constant in this study. It was found that failure modes transitioned from local outward skin wrinkling to global buckling and subsequent foam shear failure with increases in slenderness ratios. This experimental study investigates axial strength of GFRPskinned sandwich panels with a variety of PIR foam core densities, namely 32, 64, and 96 kg/m3. Each core density was
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examined in a group of panels with various slenderness ratios, ranging from 22 to 61. 2. Material specifications and testing This section highlights the properties of all materials used in the fabrication of the test specimens. 2.1. GFRP skin Unidirectional E-glass fibre fabric was used with a reported tensile strength and modulus of 3.24 GPa and 72.4 GPa, respectively, and an ultimate elongation of 5%. The dry density of the fibres is 2.55 g/cm3 [8]. Tyfo S, a two-component high elongation epoxy was used as the saturating resin. The reported tensile strength and modulus are 72.4 MPa and 3.18 GPa, respectively, the compressive strength and modulus are 86.2 MPa and 3.2 GPa, respectively, and the maximum elongation is 5%. These reported material properties are given for 72 h post cure at 60 C [8]. The GFRP material stress-strain curves in both compression and tension were previously tested (Mak and Fam [17]) and can be seen in Fig. 1. 2.2. Foam core Three separate rigid, unfaced, closed cell PIR foams were used: ELOFOAM P200, P400, and P600 with respective densities of 32, 64, and 96 kg/m3. The P200 product has a reported parallel and perpendicular shear strength of 151 kPa and 110 kPa, parallel and perpendicular shear modulus of 1.52 and 1.22 MPa, and an R-value of 1.06 m2 C/W [6]. The reported density of the P200 foam is 32 kg/m3. The density of the foam used in this study was calculated to be 31.2 kg/m3. The P400 foam has a reported parallel and perpendicular shear strength of 379 kPa and 344 kPa, parallel and perpendicular shear modulus of 5.86 and 5.17 MPa, and an R-value of 1.04 m2 C/W [6]. The reported density of the P400 foam is 64 kg/m3. The density of the foam used in this study was calculated to be 62.4 kg/m3. The P600 product has a reported parallel and perpendicular shear strength of 585 kPa and 489 kPa, parallel and perpendicular shear modulus of 7.23 and 6.06 MPa, and an R-value of 0.97 m2 C/ W [6]. The reported density of the P600 foam is 96 kg/m3. The density of the foam used in this study was calculated to be 92.7 kg/ m3. The compressive behavior of each PIR foam density was tested
according to ASTM C365 [3]. Five coupons of each density were cut into 50 50 50 mm3 cubes and tested using an Instron 8802 testing machine. Each cube was placed between two steel plates and compressed at a rate of 1.0 mm/min until 80% strain was reached. The total displacement between the two plates was measured by a linear potentiometer (LP) which used to determine the strain in each sample. The initial behavior of the PIR foams were linear until reaching a plastic plateau after initial crushing, followed by strain hardening. The response of each density can be seen in Table 1 and Fig. 2. The behavior of each PIR foam density in tension was tested according to ASTM C297 [2]. Five coupons of each density were cut into 50 50 50 mm3 cubes and tested using an Instron 8802 testing machine. Each cube was bonded on opposite sides to two steel T-sections using an epoxy resin. After curing, each steel Tsection was gripped in the testing frame and loaded at a rate of 0.5 mm/min until failure. Strain in the sample was related to the displacement between the steel grips. The response of each density was linear until failure. The response of each density can be seen in Table 1 and Fig. 1. 3. Sandwich panel tests 3.1. Test parameters A total of 36 test specimens were fabricated and tested, including 12 columns for each foam core density, with four distinct slenderness ratios, as shown in Table 2. Each panel was comprised of a 50 mm thick and 100 mm wide foam core between two symmetrical composite skins, each comprised of one 1.8 mm thick GFRP layer. Three repetitions were fabricated for each combination of foam core density and slenderness ratio. A unique identification was given to each set of three repetitions (Table 2). Series A1-A4 had a nominal foam core density of 32 kg/m3, while series B1eB4 and C1eC4 had nominal foam core densities of 64 and 96 kg/m3, respectively. The number in each series title corresponds to the height of the columns (e.g. A1, A2, A3 and A4 refer to 750, 1000, 1250 and 1500 mm, respectively). As all the columns were of identical cross-section, the slenderness ratio (kLe/r) was varied only by changing the column height from 750 to 1500 mm. For this experimental study, (k) was taken as unity for the established freely rotating pinned-pinned connections in the testing frame. The effective length (Le) was calculated as the distance between the centre of each pin, equating to the length of each specimen plus 90 mm. The radius of gyration (r) was calculated as (I/A)0.5, where (I) is the moment of inertia and (A) is the area of the cross section of GFRP skins and transformed foam core, based on the modular ratio of foam and GFRP. Because the foam modulus is considerably lower than GFRP, the resulting (kLe/r) values did not vary much with core density and the nominal values for the four heights were 25.3, 41.6, 51.3 and 60.7, respectively. 3.2. Fabrication of test specimens
Fig. 1. Stress-strain plots for GFRP skins (Mak and Fam [17]). Compression is shown as negative and tension positive.
The panels were fabricated using the wet lay-up method for manufacturing composite materials. Rigid sheets of foam were set on plastic sheeting and dabbed with a damp towel to remove any loose particles that may impact the core-skin bond. The foam was evenly wetted with epoxy resin. A sheet of dry fibre was placed along the foam sheet and additional epoxy resin was added until the installed fibres were saturated. The excess epoxy resin and air was worked out of the skin. A sheet of plastic was placed over the completed side and the panel was flipped over to expose the opposite side of the panel. The process was repeated on the opposite face of the panel. Finally, a large plastic panel and two
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Table 1 PIR Foam core properties and test results. Foam type (nominal density)
Calculated density (kg/ Compression yield stress m3) (MPa)
32 64 96
31.2 62.4 92.7
0.209 ± 0.008 0.456 ± 0.002 0.869 ± 0.005
Compression elastic modulus (MPa)
Ultimate tensile strength (MPa)
Tensile elastic modulus (MPa)
4.9 ± 0.3 12.6 ± 0.1 35.1 ± 1.3
0.198 ± 0.029 0.317 ± 0.079 0.568 ± 0.197
10.1 ± 1.5 20.0 ± 0.41 59.3 ± 3.2
for steel columns, length (L)/1000 [9], was taken as a reference to determine if the specimens were acceptable for testing. The normalized OOS profiles for each sample in comparison to the acceptable tolerance is shown in Fig. 3. Generally, the manufactured specimens' profiles were within the strict tolerance.
3.4. Test setup and instrumentation Stub columns were setup in a pin-pin concentric axial loading configuration (Fig. 4). The frame consists of two steel channels with a steel cylinder welded symmetrically on the back of each, to form a free-rotating hinge around the weak axis of the sandwich panel.
Fig. 2. Stress-strain plots for PIR foam cores. Compression is shown as negative and tension positive.
Table 2 Test matrix. Specimen ID
Nominal core density (kg/ Repetitions Length m3) (mm)
Slenderness ratio
A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4
32
25.2 41.3 51.0 60.3 25.4 41.7 51.4 60.8 25.3 41.9 51.4 61.0
64
96
3 3 3 3 3 3 3 3 3 3 3 3
750 1000 1250 1500 750 1000 1250 1500 750 1000 1250 1500
Fig. 3. Out-of-Straightness profiles for all tested columns. The 0.001 limits represent typical milling tolerances of steel.
steel plates were placed on top of the panel during the curing process to ensure a smooth skin of even thickness. Panels were manufactured for each foam core density and length, and then cut into three repetitions of 100 mm equal width.
3.3. Out-of-straightness analysis An out-of-straightness (OOS) analysis of each test specimen was conducted in order to ensure that the experimental results were not affected by any eccentricities related to material or manufacturing imperfections. Due to the high sensitivity of axially compressed members to eccentricities, the OOS analysis was conducted with a high degree of precision. The profile of each skin along the length of each sample was determined and compared to a perfectly straight reference line between the corners of each skin. Lateral variations were measured at ten equal points along the reference line to develop an OOS profile. The average OOS profile of both skins was taken as representative of the sample and then normalized by the sample length. The acceptable milling tolerance
Fig. 4. Sandwich panel column axial compression test setup.
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Table 3 Summary of test results and predicted loads. ID
A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4
Max. Load (Pu) (kN)
Axial deflection @ Pu (mm)
Lateral deflection @ Pu (mm)
Longitudinal skin strain (micro)
Avg.
S.D.
Avg.
S.D.
Avg.
S.D.
Concave
Convex
9.08 10.7 10.5 8.53 16.3 16.2 15.5 19.5 30.4 29.3 18.9 26.9
1.23 0.502 0.472 2.26 2.06 0.935 2.29 2.69 3.54 0.418 1.62 1.07
0.387 0.867 4.58 2.46 0.437 2.35 1.47 7.45 0.273 2.48 2.36 2.96
0.168 0.638 2.83 3.11 0.461 0.841 0.884 3.72 0.395 0.926 2.31 0.872
2.09 2.43 2.53 2.45 2.74 2.91 3.19 3.74 3.76 3.86 2.42 4.92
0.449 1.33 0.165 0.914 0.543 0.092 1.27 0.694 0.457 0.120 2.13 0.230
1457 1368 1743 1488 2699 2605 2453 3614 4572 4784 3224 8697
1085 750 1073 964 1973 1973 1756 1915 3610 3519 1987 1565
Failure Mode (primary þ secondary)
SB, SB, SB SB, SB, LC GB þ SB, SB, SB GB þ SB, GB þ SB, SB þ S SB, SB, SB SB þ S þ D, SB, GB þ SB GB þ SB, SB, GB þ SB þ S þ D LC, GB þ S þ D, GB þ S þ D LC, SB þ SB, S þ D SB, SB, GB þ SB þ D SB, GB þ S þ D, GB þ FC þ D GB þ SB, GB þ S þ D, GB þ FC þ D
SB ¼ local buckling of a single skin, LC ¼ localized crushing/bending of the foam core and skins near a pin connection, GB ¼ global buckling, S ¼ core shear failure, FC ¼ flexural cracking of foam core, and D ¼ debonding of a skin.
Fig. 5. Load-lateral displacement plots for each core foam density and slenderness ratio.
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Fig. 6. Load-axial displacement plots for each core foam density and slenderness ratio.
These cylinders are set inside the ears of steel bearing plates forming the top and bottom end connections. The top bearing plate was secured to the loading head of the test frame, while the bottom remained static throughout testing. Thin (3 mm thick) rubber pads were placed within each channel to ensure uniform distribution of the applied load across the entire cross section. Thin aluminum shims were installed symmetrically between the skins and steel channel, to ensure each column was tightly seated. Loading was applied using a Riehle Electro-Mechanical Testing Machine at a rate of 1 mm/min. Three linear potentiometers (LPs) were placed at quarter-lengths, along one of the sandwich skins to capture the lateral displacement of the columns. For axial displacement measurements, two additional LPs were placed against aluminum L-shaped extenders epoxied to the foam core just above the steel channel bracing. Axial displacement was measured by taking the difference between these two LPs to eliminate the effects of the rubber seating within the channels. Two electrical resistance strain gauges were adhered to each skin at
mid-height in order to measure longitudinal skin strains. 4. Experimental results and analysis Table 3 provides a summary of test results, namely the maximum loads, axial and lateral deflections at maximum loads, strains at maximum loads, and failure modes. The following sections discuss in detail the test results: 4.1. Load-lateral deflection response The load-lateral deflection response of each test repetition is shown in Fig. 5. The lateral deflections reported are measured at mid-height of each column, and were the maximum values along the height. Initially, very little lateral deflection is seen, then, the rate increases rapidly as the specimen approaches failure. Fig. 5 shows the trends for each series of density (A, B and C) (columns left to right) and for increasing slenderness ratio (1e4) (rows top to
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bottom). For all foam densities, there was a positive relationship between slenderness ratio and lateral deflection. The relative lateral deflections reached at the peak loads are significantly lower for series A1, B1 and C1 than the remaining slenderness ratios. It is also noticed that at a given load level, the lateral deflection reduces as core density increases from group A to C, for a given slenderness ratio.
engaging the skins, and (b) a stiffer response till failure once the rubber is compressed and GFRP skins are fully engaged. The axial displacement appears to be insensitive to slenderness ratio or core density, when comparing specimens at a given load level. At maximum load, the axial displacement increased with the density, simply due to the increase in maximum load. 4.3. Load-strain response
4.2. Load-axial displacement response The load-axial deflection response of each test specimen is illustrated in Fig. 6, showing the trends for each series of density (A, B and C) (columns left to right) and for increasing slenderness ratio (1e4) (rows top to bottom). The load-axial deflection curve consists of two distinct regions: (a) an initial low stiffness attributed to the initial compression of the core by the thin layer of rubber used at the end, which is actually stiffer than the foam, leading to some axial deformation of the core at low load levels prior to fully
The relationship between load and longitudinal skin strains at mid-height are shown in Figs. 7 and 8, for the concave and convex sides, respectively. Both skins behave similarly initially, under compressive strains, until rapidly diverging before failure where the concave skin reaches higher compressive strain while the convex skin curve reduces and in some cases reverses direction. Nonetheless, both skins remained in compression at maximum loads, even at high slenderness that were governed by global buckling. The degree of divergence in strains between both skins
Fig. 7. Load-longitudinal strain plots of the concave skin for each core foam density and slenderness ratio.
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Fig. 8. Load-longitudinal strain plots of the convex skin for each core foam density and slenderness ratio.
was dependent on both foam core density and column slenderness. The 32 kg/m3 density specimens generally showed less divergence between skins than higher density specimens. Shorter specimens showed little distinction between skin behavior while slenderer ones demonstrated greater differentiation between the two skins' strains. In general, the highest compressive strain achieved in the entire group was quite lower than the crushing strain based on coupon tests (Fig. 1). 4.4. Failure modes A number of failure modes were observed across the range of slenderness ratios tested. A sample of each failure mode is presented in Fig. 9. Series A1, B1 and C1 (the shortest) experienced localized stability failures, the most common being outward buckling of the convex side skin, followed by secondary buckling of the entire column. In contrast, series A4, B4 and C4 (the tallest) experienced global buckling followed by secondary core shear
failure and tertiary skin debonding. The two intermediate series (2 and 3) saw a combination of local and global stability failures, distinguishing slenderness ratios of 40e50 as an approximate transition range. Three repetitions failed due to localized inward skin wrinkling and foam crushing, however, each failure occurred at different foam core densities and slenderness ratios. Secondary core shear failures were only seen in columns with slenderness ratios of 51e61, and were more frequent in the B and C series than A, due to the higher achieved loads associated with higher core densities. One repetition in C3 and C4 saw a single horizontal flexure crack in the core material at mid-height following global buckling. 4.5. Effect of foam density on axial behavior Fig. 10 shows the variation of maximum loads with foam core density, for various slenderness ratios tested in this study (series 1 to 4). Additional series 5 show the results of another set of axially
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Fig. 9. Samples of all failure modes: single skin buckling (a), primary single skin buckling with secondary core shear and partial debonding of opposite skin (b and c), global buckling followed by secondary core shear and skin debonding (d), global buckling with core flexural cracks at mid-span (e), and localized end crushing of foam core and skins (f).
loaded specimens tested by CoDyre et al. [18]. This series involved specimens of identical cross-sectional properties, made of the same three core densities and GFRP skins as in this study, but were only 500 mm tall, which provides a (kLe/r) ratio of 22. There is a strong near-linear relationship between foam core density and peak load. This relationship is not significantly affected by slenderness ratio, statistically. Taking the lowest (32 kg/m3) density as a baseline (series A), increasing the foam core density to 64 and 96 kg/m3 results in average increases in peak load, for all slenderness ratios, of 71% and 170%, respectively. The increase in axial strength with foam density is quite remarkable, especially when its mechanical properties (Fig. 2) are significantly lower than the skins (Fig. 1). The reason being axial strength is governed, either by local skin wrinkling or global buckling, and not skin crushing. Higher core density affects both types of stability failures, by providing higher foam peeling (tensile) strength, which is associated with outward skin buckling, and
higher composite action as a result of the higher shear modulus, which is associated with global buckling.
4.6. Effect of slenderness on axial behavior Fig. 11 shows the effect of slenderness ratio on peak load, for each core density. Results at (kLe/r) ratio of 22 from CoDyre et al. [18] are also included. Excluding the outlying C3 series at (kLe/r) ratio of 51, no significant variation in peak load can be seen across all slenderness ratios tested. Variances in average peak load are between 1.9 and 2.5 kN, representing 19, 15, and 8% of the total peak loads for the 32, 64 and 96 kg/m3 samples, respectively. C3 series is clearly an outlier, showing unjustifiably lower than expected strength. As this reduction in peak strength is seen in each C3 repetition, it is likely attributed to a manufacturing defect as all three repetitions were fabricated from the same batch.
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Fig. 12. Experimental and predicted peak loads.
Fig. 10. Effect of foam core density on peak load for varying slenderness ratios (slenderness # 5 is 22, from CoDyre et al. [18]).
Pcr ¼ 2bt scr
(3)
All test panels were assessed using Equation (2), and the values for the 32, 64 and 96 kN/m3 cores were 1183, 460, and 165, respectively, which are all larger than 100, confirming the validity of Equation (1). Fig. 12 shows the predicted critical buckling loads. The model appears to be generally on the conservative side with an average predicted-to-experimental load values of 0.86.
5. Conclusions
Fig. 11. Effect of slenderness on peak load for varying foam core densities (load values at slenderness 22 are from CoDyre et al. [18]).
4.7. Simplified analytical model For panels with soft core, Allen [1] suggested the following expression for the average critical stress in the skins (scr), which incorporates the core softness and shear deformation into the Euler buckling concept:
scr
8 2 9 > 2 > < = 3 1 þ ct t # ¼ 1þ" > 12 Le > : ; 2 E 1 þ p2 Gfc Ltc2
p2 Ef
(1)
e
To assess whether core is considered soft, the following expression is used:
6
Ef t c þ t 2 100 ðsoft coreÞ c Ec c
(2)
where Ef, Ec and Gc are Young's moduli of the skins in the longitudinal direction and the foam core and the shear modulus of the core, respectively. t, c, and Le are the skin thickness, core thickness and the length between the centers of the end pins, respectively. The critical buckling load is then calculated using Equation (3), where b is the width of the panel:
In this study, the effects of PIR foam core density and slenderness on the axial strength and behavior of sandwich panels with GFRP skins, were examined. The following conclusions were drawn from this experimental study: 1. Increasing the density of the PIR foam core resulted in significant increases in the peak compressive load, for all slenderness ratios (kLe/r) tested. Doubling and tripling the density to 64 and 96 kg/m3 resulted in average increases in peak load of 71 and 170%, respectively. 2. For each core density, the peak loads achieved were insensitive to (kLe/r), unlike the expected reducing trend of conventional Euler's theory. The maximum variance in peak loads seen across all (kLe/r) ratios were only 19, 15, and 8% for the 32, 64 and 96 kg/m3 foam cores, respectively. 3. Failure mode of axially loaded GFRP-skinned PIR foam core sandwich panels column is governed by (kLe/r) ratio. For all PIR foam core densities, specimens with low (kLe/r) of 22e25 experienced localized failures due to single skin buckling. Specimens with high (kLe/r) of 51e61 failed primarily due to global buckling followed by secondary core shear failure with skin debonding. Specimens of intermediate (kLe/r) of 41e51 experienced a combination of localized and global failure modes. 4. At failure, GFRP skin strains did not approach material compressive failure strain at any (kLe/r) ratio or core density, as local or global buckling always governed. 5. Lateral deflections reduced with core density, at a given load. This suggests an enhanced composite action as density increases due to the increased shear modulus and reduced shear deformations.
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Compliance with ethical standards Funding: This study was funded by Agriculture and Agri-Foods Canada (AAFC) and Bioindustrial Innovation Canada (BIC) (Grant: Growing Forward 2 Program). The authors also would like to acknowledge the in-kind contribution of the Elliot Company of Indianapolis and Fyfe Company LCC. Conflict of interest: The authors have no conflict of interest. References [1] Allen HG. Analysis and design of structural sandwich panels. Oxford, U.K: Pergamon Press; 1969. [2] ASTM. C297: standard test method for flatwise tensile strength of sandwich constructions. 2016 [West Conshohocken, PA]. [3] ASTM. C365: standard test method for flatwise compressive properties of sandwich cores. 2011 [West Conshohocken, PA]. [4] Carlsson LA. Structural and failure mechanics of sandwich composites. New York: Springer; 2011. [5] Chemami A, Bey K, Gilgert J, Azari Z. Behaviour of composite sandwich foamlaminated glass/epoxy under solicitation static and fatigue. Compos Part B Eng 2012;43(3):1178e84. [6] Elliott Company. ELFOAM technical data. ELFOAM; 2012, April. http://www. elliottfoam.com/tech.html. Accessed September 2015. [7] Frostig Y. Buckling of sandwich panels with a flexible core high-order theory. Int J Solids Struct 1997;35(3e4):183e204.
[8] Fyfe Co. LLC. Tyfo SEH-51A composite. Tyfo fibrwrap composite systems. 2012. http://www.fyfeco.com/Products/~/media/Files/Fyfe/2013-Products/Tyfo% 20SEH-51A%20Comp.ashx. Accessed September 2015. [9] Galambos TV. Guide to stability design criteria for metal Structures. fifth ed. Wiley; 1998. [10] Huang W, Zhang W, Ye N, Gao Y, Ren P. Dynamic response and failure of PVC foam core metallic sandwich subjected to underwater impulsive loading. Compos Part B Eng 2016;97:226e38. [11] Matheson H, Fam A. Axial loading tests and simplified modeling of sandwich panels with GFRP skins and soft core at various slenderness ratios. J Compos Constr 2014;19(2):13. [12] Rajapakse YDS, Hui D. Marine composites and sandwich structures. Compos Part B Eng 2008;39(1):1e4. [14] Yang Y, Li B, Chen Z, Sui N, Chen Z, Saeed MU, et al. Acoustic properties of glass fiber assembly-filled honeycomb sandwich panels. Compos Part B Eng 2016;96:281e6. [15] Zhang P, Cheng Y, Liu J, Li Y, Zhang C, Hou H, et al. Experimental study on the dynamic response of foam-filled corrugated core sandwich panels subjected to air blast loading. Compos Part B Eng 2016;105:67e81. [16] Sharaf T, Shawkat W, Fam A. Structural performance of sandwich wall panels with different foam core densities in one-way bending. J. Compos. Mater. 2010;44(19):2249e63. [17] Mak K, Fam A, MacDougall C. Flexural behavior of sandwich panels with bioFRP skins made of flax fibers and epoxidized pine oil resin. ASCE J. Compos. Construc. 2015. http://dx.doi.org/10.1061/(ASCE)CC.1943-5614.0000560. 04015005. [18] CoDyre, L, Fam, A. Axial strength of sandwich panels with natural flax fiber composite skins and different foam core densities at various slenderness. ASCE J. Compos. Construc. 2016 (under review).
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Composites Part B journal homepage: www.elsevier.com/locate/compositesb
The effect of foam core density at various slenderness ratios on axial strength of sandwich panels with glass-FRP skins Luke CoDyre, Amir Fam* Department of Civil Engineering, Queen's University, Kingston, ON, K7L 3N6, Canada
a r t i c l e i n f o
a b s t r a c t
Article history: Received 24 June 2016 Received in revised form 5 September 2016 Accepted 5 September 2016 Available online 7 September 2016
In this study, the effect of the density of a polyisocyanurate (PIR) foam core on the behavior of axially loaded sandwich panels with glass fibre reinforced polymer (GFRP) skins are examined, for different panel heights. Thirty-six test specimens with a cross-sectional foam core area of 100 50 mm2 and symmetrical 100 mm wide GFRP skins were fabricated with heights of 750, 1000, 1250, and 1500 mm to examine various slenderness ratios (kLe/r). Three PIR foam core materials with densities of 32, 64 and 96 kg/m3 were used. An out-of-straightness assessment was first conducted to ensure that eccentricities were within acceptable limits. All samples were tested to failure in axial compression with pinned end conditions. It was shown that doubling and tripling the core density led to increases in peak load of 71 and 170%, respectively, due to the enhanced composite action and reduced shear deformations. On the other hand, slenderness had insignificant impact on peak loads, contrary to what one would expect from Euler's theory. The reason being the variety of governing failure modes. Columns with low (kLe/r) of 22 e25 failed largely due to localized single skin buckling, while those with high (kLe/r) of 51e61 failed primarily by global buckling followed by secondary core shear failure with skin debonding. Columns with intermediate (kLe/r) experienced a combination of localized and global failure modes. © 2016 Elsevier Ltd. All rights reserved.
Keywords: Sandwich panel GFRP skin Foam core Density Slenderness Axial
1. Introduction Sandwich panels have been widely studied and used for quite some time in aerospace and other industries, including marine applications (e.g. Ref. [12]. Studies have focused on their behavior under dynamic air-blast loading [15], behavior when exposed to under-water impulsive loading [10], their fatigue behavior [5], as well as acoustic behavior [14]. Traditionally, the applications of sandwich panels in civil engineering have been limited but in recent years they started gaining attention with the rise in popularity of modular and rapid construction techniques. These composite systems, comprised of a low density insulating core material between two stiff skins, boast high strength-to-weight ratios making them highly transportable and useful as pre-fabricated wall, roof or deck elements. A variety of closed-form foams can be used as the core material due to their water impermeability and insulation properties, such as polyisocyanurate (PIR), polyvinylchloride (PVC) and polyurethane (PUR) [4]. Among the many skin material alternatives, glass fibre reinforced polymer (GFRP) has
* Corresponding author. E-mail address: [email protected] (A. Fam). http://dx.doi.org/10.1016/j.compositesb.2016.09.016 1359-8368/© 2016 Elsevier Ltd. All rights reserved.
been extensively researched due to its high specific strength and corrosion-resistance. Allen [1] and Frostig [7] have shown that the shear stiffness of concentrically loaded slender soft-core sandwich panels is significant in the determination of the critical buckling load for the section. A reduction in the typical Euler buckling capacity is seen due to reduced core shear stiffness. Additional deflections due to core shear deformations are also seen in the flexural behavior of sandwich panels, however, increasing the density of the core material has been found to reduce this effect. Sharaf et al. [16] varied the density of PUR foam cores in GFRP skinned flexural panels and observed significantly reduced shear deformations and increased ultimate capacity and stiffness. Mathieson and Fam [11] conducted a study on the concentric axial loading of polyurethane foam core and GFRP skinned sandwich panels to determine the effects of slenderness on panel behavior. Core density was kept constant in this study. It was found that failure modes transitioned from local outward skin wrinkling to global buckling and subsequent foam shear failure with increases in slenderness ratios. This experimental study investigates axial strength of GFRPskinned sandwich panels with a variety of PIR foam core densities, namely 32, 64, and 96 kg/m3. Each core density was
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examined in a group of panels with various slenderness ratios, ranging from 22 to 61. 2. Material specifications and testing This section highlights the properties of all materials used in the fabrication of the test specimens. 2.1. GFRP skin Unidirectional E-glass fibre fabric was used with a reported tensile strength and modulus of 3.24 GPa and 72.4 GPa, respectively, and an ultimate elongation of 5%. The dry density of the fibres is 2.55 g/cm3 [8]. Tyfo S, a two-component high elongation epoxy was used as the saturating resin. The reported tensile strength and modulus are 72.4 MPa and 3.18 GPa, respectively, the compressive strength and modulus are 86.2 MPa and 3.2 GPa, respectively, and the maximum elongation is 5%. These reported material properties are given for 72 h post cure at 60 C [8]. The GFRP material stress-strain curves in both compression and tension were previously tested (Mak and Fam [17]) and can be seen in Fig. 1. 2.2. Foam core Three separate rigid, unfaced, closed cell PIR foams were used: ELOFOAM P200, P400, and P600 with respective densities of 32, 64, and 96 kg/m3. The P200 product has a reported parallel and perpendicular shear strength of 151 kPa and 110 kPa, parallel and perpendicular shear modulus of 1.52 and 1.22 MPa, and an R-value of 1.06 m2 C/W [6]. The reported density of the P200 foam is 32 kg/m3. The density of the foam used in this study was calculated to be 31.2 kg/m3. The P400 foam has a reported parallel and perpendicular shear strength of 379 kPa and 344 kPa, parallel and perpendicular shear modulus of 5.86 and 5.17 MPa, and an R-value of 1.04 m2 C/W [6]. The reported density of the P400 foam is 64 kg/m3. The density of the foam used in this study was calculated to be 62.4 kg/m3. The P600 product has a reported parallel and perpendicular shear strength of 585 kPa and 489 kPa, parallel and perpendicular shear modulus of 7.23 and 6.06 MPa, and an R-value of 0.97 m2 C/ W [6]. The reported density of the P600 foam is 96 kg/m3. The density of the foam used in this study was calculated to be 92.7 kg/ m3. The compressive behavior of each PIR foam density was tested
according to ASTM C365 [3]. Five coupons of each density were cut into 50 50 50 mm3 cubes and tested using an Instron 8802 testing machine. Each cube was placed between two steel plates and compressed at a rate of 1.0 mm/min until 80% strain was reached. The total displacement between the two plates was measured by a linear potentiometer (LP) which used to determine the strain in each sample. The initial behavior of the PIR foams were linear until reaching a plastic plateau after initial crushing, followed by strain hardening. The response of each density can be seen in Table 1 and Fig. 2. The behavior of each PIR foam density in tension was tested according to ASTM C297 [2]. Five coupons of each density were cut into 50 50 50 mm3 cubes and tested using an Instron 8802 testing machine. Each cube was bonded on opposite sides to two steel T-sections using an epoxy resin. After curing, each steel Tsection was gripped in the testing frame and loaded at a rate of 0.5 mm/min until failure. Strain in the sample was related to the displacement between the steel grips. The response of each density was linear until failure. The response of each density can be seen in Table 1 and Fig. 1. 3. Sandwich panel tests 3.1. Test parameters A total of 36 test specimens were fabricated and tested, including 12 columns for each foam core density, with four distinct slenderness ratios, as shown in Table 2. Each panel was comprised of a 50 mm thick and 100 mm wide foam core between two symmetrical composite skins, each comprised of one 1.8 mm thick GFRP layer. Three repetitions were fabricated for each combination of foam core density and slenderness ratio. A unique identification was given to each set of three repetitions (Table 2). Series A1-A4 had a nominal foam core density of 32 kg/m3, while series B1eB4 and C1eC4 had nominal foam core densities of 64 and 96 kg/m3, respectively. The number in each series title corresponds to the height of the columns (e.g. A1, A2, A3 and A4 refer to 750, 1000, 1250 and 1500 mm, respectively). As all the columns were of identical cross-section, the slenderness ratio (kLe/r) was varied only by changing the column height from 750 to 1500 mm. For this experimental study, (k) was taken as unity for the established freely rotating pinned-pinned connections in the testing frame. The effective length (Le) was calculated as the distance between the centre of each pin, equating to the length of each specimen plus 90 mm. The radius of gyration (r) was calculated as (I/A)0.5, where (I) is the moment of inertia and (A) is the area of the cross section of GFRP skins and transformed foam core, based on the modular ratio of foam and GFRP. Because the foam modulus is considerably lower than GFRP, the resulting (kLe/r) values did not vary much with core density and the nominal values for the four heights were 25.3, 41.6, 51.3 and 60.7, respectively. 3.2. Fabrication of test specimens
Fig. 1. Stress-strain plots for GFRP skins (Mak and Fam [17]). Compression is shown as negative and tension positive.
The panels were fabricated using the wet lay-up method for manufacturing composite materials. Rigid sheets of foam were set on plastic sheeting and dabbed with a damp towel to remove any loose particles that may impact the core-skin bond. The foam was evenly wetted with epoxy resin. A sheet of dry fibre was placed along the foam sheet and additional epoxy resin was added until the installed fibres were saturated. The excess epoxy resin and air was worked out of the skin. A sheet of plastic was placed over the completed side and the panel was flipped over to expose the opposite side of the panel. The process was repeated on the opposite face of the panel. Finally, a large plastic panel and two
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Table 1 PIR Foam core properties and test results. Foam type (nominal density)
Calculated density (kg/ Compression yield stress m3) (MPa)
32 64 96
31.2 62.4 92.7
0.209 ± 0.008 0.456 ± 0.002 0.869 ± 0.005
Compression elastic modulus (MPa)
Ultimate tensile strength (MPa)
Tensile elastic modulus (MPa)
4.9 ± 0.3 12.6 ± 0.1 35.1 ± 1.3
0.198 ± 0.029 0.317 ± 0.079 0.568 ± 0.197
10.1 ± 1.5 20.0 ± 0.41 59.3 ± 3.2
for steel columns, length (L)/1000 [9], was taken as a reference to determine if the specimens were acceptable for testing. The normalized OOS profiles for each sample in comparison to the acceptable tolerance is shown in Fig. 3. Generally, the manufactured specimens' profiles were within the strict tolerance.
3.4. Test setup and instrumentation Stub columns were setup in a pin-pin concentric axial loading configuration (Fig. 4). The frame consists of two steel channels with a steel cylinder welded symmetrically on the back of each, to form a free-rotating hinge around the weak axis of the sandwich panel.
Fig. 2. Stress-strain plots for PIR foam cores. Compression is shown as negative and tension positive.
Table 2 Test matrix. Specimen ID
Nominal core density (kg/ Repetitions Length m3) (mm)
Slenderness ratio
A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4
32
25.2 41.3 51.0 60.3 25.4 41.7 51.4 60.8 25.3 41.9 51.4 61.0
64
96
3 3 3 3 3 3 3 3 3 3 3 3
750 1000 1250 1500 750 1000 1250 1500 750 1000 1250 1500
Fig. 3. Out-of-Straightness profiles for all tested columns. The 0.001 limits represent typical milling tolerances of steel.
steel plates were placed on top of the panel during the curing process to ensure a smooth skin of even thickness. Panels were manufactured for each foam core density and length, and then cut into three repetitions of 100 mm equal width.
3.3. Out-of-straightness analysis An out-of-straightness (OOS) analysis of each test specimen was conducted in order to ensure that the experimental results were not affected by any eccentricities related to material or manufacturing imperfections. Due to the high sensitivity of axially compressed members to eccentricities, the OOS analysis was conducted with a high degree of precision. The profile of each skin along the length of each sample was determined and compared to a perfectly straight reference line between the corners of each skin. Lateral variations were measured at ten equal points along the reference line to develop an OOS profile. The average OOS profile of both skins was taken as representative of the sample and then normalized by the sample length. The acceptable milling tolerance
Fig. 4. Sandwich panel column axial compression test setup.
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Table 3 Summary of test results and predicted loads. ID
A1 A2 A3 A4 B1 B2 B3 B4 C1 C2 C3 C4
Max. Load (Pu) (kN)
Axial deflection @ Pu (mm)
Lateral deflection @ Pu (mm)
Longitudinal skin strain (micro)
Avg.
S.D.
Avg.
S.D.
Avg.
S.D.
Concave
Convex
9.08 10.7 10.5 8.53 16.3 16.2 15.5 19.5 30.4 29.3 18.9 26.9
1.23 0.502 0.472 2.26 2.06 0.935 2.29 2.69 3.54 0.418 1.62 1.07
0.387 0.867 4.58 2.46 0.437 2.35 1.47 7.45 0.273 2.48 2.36 2.96
0.168 0.638 2.83 3.11 0.461 0.841 0.884 3.72 0.395 0.926 2.31 0.872
2.09 2.43 2.53 2.45 2.74 2.91 3.19 3.74 3.76 3.86 2.42 4.92
0.449 1.33 0.165 0.914 0.543 0.092 1.27 0.694 0.457 0.120 2.13 0.230
1457 1368 1743 1488 2699 2605 2453 3614 4572 4784 3224 8697
1085 750 1073 964 1973 1973 1756 1915 3610 3519 1987 1565
Failure Mode (primary þ secondary)
SB, SB, SB SB, SB, LC GB þ SB, SB, SB GB þ SB, GB þ SB, SB þ S SB, SB, SB SB þ S þ D, SB, GB þ SB GB þ SB, SB, GB þ SB þ S þ D LC, GB þ S þ D, GB þ S þ D LC, SB þ SB, S þ D SB, SB, GB þ SB þ D SB, GB þ S þ D, GB þ FC þ D GB þ SB, GB þ S þ D, GB þ FC þ D
SB ¼ local buckling of a single skin, LC ¼ localized crushing/bending of the foam core and skins near a pin connection, GB ¼ global buckling, S ¼ core shear failure, FC ¼ flexural cracking of foam core, and D ¼ debonding of a skin.
Fig. 5. Load-lateral displacement plots for each core foam density and slenderness ratio.
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Fig. 6. Load-axial displacement plots for each core foam density and slenderness ratio.
These cylinders are set inside the ears of steel bearing plates forming the top and bottom end connections. The top bearing plate was secured to the loading head of the test frame, while the bottom remained static throughout testing. Thin (3 mm thick) rubber pads were placed within each channel to ensure uniform distribution of the applied load across the entire cross section. Thin aluminum shims were installed symmetrically between the skins and steel channel, to ensure each column was tightly seated. Loading was applied using a Riehle Electro-Mechanical Testing Machine at a rate of 1 mm/min. Three linear potentiometers (LPs) were placed at quarter-lengths, along one of the sandwich skins to capture the lateral displacement of the columns. For axial displacement measurements, two additional LPs were placed against aluminum L-shaped extenders epoxied to the foam core just above the steel channel bracing. Axial displacement was measured by taking the difference between these two LPs to eliminate the effects of the rubber seating within the channels. Two electrical resistance strain gauges were adhered to each skin at
mid-height in order to measure longitudinal skin strains. 4. Experimental results and analysis Table 3 provides a summary of test results, namely the maximum loads, axial and lateral deflections at maximum loads, strains at maximum loads, and failure modes. The following sections discuss in detail the test results: 4.1. Load-lateral deflection response The load-lateral deflection response of each test repetition is shown in Fig. 5. The lateral deflections reported are measured at mid-height of each column, and were the maximum values along the height. Initially, very little lateral deflection is seen, then, the rate increases rapidly as the specimen approaches failure. Fig. 5 shows the trends for each series of density (A, B and C) (columns left to right) and for increasing slenderness ratio (1e4) (rows top to
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bottom). For all foam densities, there was a positive relationship between slenderness ratio and lateral deflection. The relative lateral deflections reached at the peak loads are significantly lower for series A1, B1 and C1 than the remaining slenderness ratios. It is also noticed that at a given load level, the lateral deflection reduces as core density increases from group A to C, for a given slenderness ratio.
engaging the skins, and (b) a stiffer response till failure once the rubber is compressed and GFRP skins are fully engaged. The axial displacement appears to be insensitive to slenderness ratio or core density, when comparing specimens at a given load level. At maximum load, the axial displacement increased with the density, simply due to the increase in maximum load. 4.3. Load-strain response
4.2. Load-axial displacement response The load-axial deflection response of each test specimen is illustrated in Fig. 6, showing the trends for each series of density (A, B and C) (columns left to right) and for increasing slenderness ratio (1e4) (rows top to bottom). The load-axial deflection curve consists of two distinct regions: (a) an initial low stiffness attributed to the initial compression of the core by the thin layer of rubber used at the end, which is actually stiffer than the foam, leading to some axial deformation of the core at low load levels prior to fully
The relationship between load and longitudinal skin strains at mid-height are shown in Figs. 7 and 8, for the concave and convex sides, respectively. Both skins behave similarly initially, under compressive strains, until rapidly diverging before failure where the concave skin reaches higher compressive strain while the convex skin curve reduces and in some cases reverses direction. Nonetheless, both skins remained in compression at maximum loads, even at high slenderness that were governed by global buckling. The degree of divergence in strains between both skins
Fig. 7. Load-longitudinal strain plots of the concave skin for each core foam density and slenderness ratio.
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Fig. 8. Load-longitudinal strain plots of the convex skin for each core foam density and slenderness ratio.
was dependent on both foam core density and column slenderness. The 32 kg/m3 density specimens generally showed less divergence between skins than higher density specimens. Shorter specimens showed little distinction between skin behavior while slenderer ones demonstrated greater differentiation between the two skins' strains. In general, the highest compressive strain achieved in the entire group was quite lower than the crushing strain based on coupon tests (Fig. 1). 4.4. Failure modes A number of failure modes were observed across the range of slenderness ratios tested. A sample of each failure mode is presented in Fig. 9. Series A1, B1 and C1 (the shortest) experienced localized stability failures, the most common being outward buckling of the convex side skin, followed by secondary buckling of the entire column. In contrast, series A4, B4 and C4 (the tallest) experienced global buckling followed by secondary core shear
failure and tertiary skin debonding. The two intermediate series (2 and 3) saw a combination of local and global stability failures, distinguishing slenderness ratios of 40e50 as an approximate transition range. Three repetitions failed due to localized inward skin wrinkling and foam crushing, however, each failure occurred at different foam core densities and slenderness ratios. Secondary core shear failures were only seen in columns with slenderness ratios of 51e61, and were more frequent in the B and C series than A, due to the higher achieved loads associated with higher core densities. One repetition in C3 and C4 saw a single horizontal flexure crack in the core material at mid-height following global buckling. 4.5. Effect of foam density on axial behavior Fig. 10 shows the variation of maximum loads with foam core density, for various slenderness ratios tested in this study (series 1 to 4). Additional series 5 show the results of another set of axially
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Fig. 9. Samples of all failure modes: single skin buckling (a), primary single skin buckling with secondary core shear and partial debonding of opposite skin (b and c), global buckling followed by secondary core shear and skin debonding (d), global buckling with core flexural cracks at mid-span (e), and localized end crushing of foam core and skins (f).
loaded specimens tested by CoDyre et al. [18]. This series involved specimens of identical cross-sectional properties, made of the same three core densities and GFRP skins as in this study, but were only 500 mm tall, which provides a (kLe/r) ratio of 22. There is a strong near-linear relationship between foam core density and peak load. This relationship is not significantly affected by slenderness ratio, statistically. Taking the lowest (32 kg/m3) density as a baseline (series A), increasing the foam core density to 64 and 96 kg/m3 results in average increases in peak load, for all slenderness ratios, of 71% and 170%, respectively. The increase in axial strength with foam density is quite remarkable, especially when its mechanical properties (Fig. 2) are significantly lower than the skins (Fig. 1). The reason being axial strength is governed, either by local skin wrinkling or global buckling, and not skin crushing. Higher core density affects both types of stability failures, by providing higher foam peeling (tensile) strength, which is associated with outward skin buckling, and
higher composite action as a result of the higher shear modulus, which is associated with global buckling.
4.6. Effect of slenderness on axial behavior Fig. 11 shows the effect of slenderness ratio on peak load, for each core density. Results at (kLe/r) ratio of 22 from CoDyre et al. [18] are also included. Excluding the outlying C3 series at (kLe/r) ratio of 51, no significant variation in peak load can be seen across all slenderness ratios tested. Variances in average peak load are between 1.9 and 2.5 kN, representing 19, 15, and 8% of the total peak loads for the 32, 64 and 96 kg/m3 samples, respectively. C3 series is clearly an outlier, showing unjustifiably lower than expected strength. As this reduction in peak strength is seen in each C3 repetition, it is likely attributed to a manufacturing defect as all three repetitions were fabricated from the same batch.
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Fig. 12. Experimental and predicted peak loads.
Fig. 10. Effect of foam core density on peak load for varying slenderness ratios (slenderness # 5 is 22, from CoDyre et al. [18]).
Pcr ¼ 2bt scr
(3)
All test panels were assessed using Equation (2), and the values for the 32, 64 and 96 kN/m3 cores were 1183, 460, and 165, respectively, which are all larger than 100, confirming the validity of Equation (1). Fig. 12 shows the predicted critical buckling loads. The model appears to be generally on the conservative side with an average predicted-to-experimental load values of 0.86.
5. Conclusions
Fig. 11. Effect of slenderness on peak load for varying foam core densities (load values at slenderness 22 are from CoDyre et al. [18]).
4.7. Simplified analytical model For panels with soft core, Allen [1] suggested the following expression for the average critical stress in the skins (scr), which incorporates the core softness and shear deformation into the Euler buckling concept:
scr
8 2 9 > 2 > < = 3 1 þ ct t # ¼ 1þ" > 12 Le > : ; 2 E 1 þ p2 Gfc Ltc2
p2 Ef
(1)
e
To assess whether core is considered soft, the following expression is used:
6
Ef t c þ t 2 100 ðsoft coreÞ c Ec c
(2)
where Ef, Ec and Gc are Young's moduli of the skins in the longitudinal direction and the foam core and the shear modulus of the core, respectively. t, c, and Le are the skin thickness, core thickness and the length between the centers of the end pins, respectively. The critical buckling load is then calculated using Equation (3), where b is the width of the panel:
In this study, the effects of PIR foam core density and slenderness on the axial strength and behavior of sandwich panels with GFRP skins, were examined. The following conclusions were drawn from this experimental study: 1. Increasing the density of the PIR foam core resulted in significant increases in the peak compressive load, for all slenderness ratios (kLe/r) tested. Doubling and tripling the density to 64 and 96 kg/m3 resulted in average increases in peak load of 71 and 170%, respectively. 2. For each core density, the peak loads achieved were insensitive to (kLe/r), unlike the expected reducing trend of conventional Euler's theory. The maximum variance in peak loads seen across all (kLe/r) ratios were only 19, 15, and 8% for the 32, 64 and 96 kg/m3 foam cores, respectively. 3. Failure mode of axially loaded GFRP-skinned PIR foam core sandwich panels column is governed by (kLe/r) ratio. For all PIR foam core densities, specimens with low (kLe/r) of 22e25 experienced localized failures due to single skin buckling. Specimens with high (kLe/r) of 51e61 failed primarily due to global buckling followed by secondary core shear failure with skin debonding. Specimens of intermediate (kLe/r) of 41e51 experienced a combination of localized and global failure modes. 4. At failure, GFRP skin strains did not approach material compressive failure strain at any (kLe/r) ratio or core density, as local or global buckling always governed. 5. Lateral deflections reduced with core density, at a given load. This suggests an enhanced composite action as density increases due to the increased shear modulus and reduced shear deformations.
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Compliance with ethical standards Funding: This study was funded by Agriculture and Agri-Foods Canada (AAFC) and Bioindustrial Innovation Canada (BIC) (Grant: Growing Forward 2 Program). The authors also would like to acknowledge the in-kind contribution of the Elliot Company of Indianapolis and Fyfe Company LCC. Conflict of interest: The authors have no conflict of interest. References [1] Allen HG. Analysis and design of structural sandwich panels. Oxford, U.K: Pergamon Press; 1969. [2] ASTM. C297: standard test method for flatwise tensile strength of sandwich constructions. 2016 [West Conshohocken, PA]. [3] ASTM. C365: standard test method for flatwise compressive properties of sandwich cores. 2011 [West Conshohocken, PA]. [4] Carlsson LA. Structural and failure mechanics of sandwich composites. New York: Springer; 2011. [5] Chemami A, Bey K, Gilgert J, Azari Z. Behaviour of composite sandwich foamlaminated glass/epoxy under solicitation static and fatigue. Compos Part B Eng 2012;43(3):1178e84. [6] Elliott Company. ELFOAM technical data. ELFOAM; 2012, April. http://www. elliottfoam.com/tech.html. Accessed September 2015. [7] Frostig Y. Buckling of sandwich panels with a flexible core high-order theory. Int J Solids Struct 1997;35(3e4):183e204.
[8] Fyfe Co. LLC. Tyfo SEH-51A composite. Tyfo fibrwrap composite systems. 2012. http://www.fyfeco.com/Products/~/media/Files/Fyfe/2013-Products/Tyfo% 20SEH-51A%20Comp.ashx. Accessed September 2015. [9] Galambos TV. Guide to stability design criteria for metal Structures. fifth ed. Wiley; 1998. [10] Huang W, Zhang W, Ye N, Gao Y, Ren P. Dynamic response and failure of PVC foam core metallic sandwich subjected to underwater impulsive loading. Compos Part B Eng 2016;97:226e38. [11] Matheson H, Fam A. Axial loading tests and simplified modeling of sandwich panels with GFRP skins and soft core at various slenderness ratios. J Compos Constr 2014;19(2):13. [12] Rajapakse YDS, Hui D. Marine composites and sandwich structures. Compos Part B Eng 2008;39(1):1e4. [14] Yang Y, Li B, Chen Z, Sui N, Chen Z, Saeed MU, et al. Acoustic properties of glass fiber assembly-filled honeycomb sandwich panels. Compos Part B Eng 2016;96:281e6. [15] Zhang P, Cheng Y, Liu J, Li Y, Zhang C, Hou H, et al. Experimental study on the dynamic response of foam-filled corrugated core sandwich panels subjected to air blast loading. Compos Part B Eng 2016;105:67e81. [16] Sharaf T, Shawkat W, Fam A. Structural performance of sandwich wall panels with different foam core densities in one-way bending. J. Compos. Mater. 2010;44(19):2249e63. [17] Mak K, Fam A, MacDougall C. Flexural behavior of sandwich panels with bioFRP skins made of flax fibers and epoxidized pine oil resin. ASCE J. Compos. Construc. 2015. http://dx.doi.org/10.1061/(ASCE)CC.1943-5614.0000560. 04015005. [18] CoDyre, L, Fam, A. Axial strength of sandwich panels with natural flax fiber composite skins and different foam core densities at various slenderness. ASCE J. Compos. Construc. 2016 (under review).