A low velocity impact study on press formed thermoplastic honeycomb sandwich panels

Composite Structures 225 (2019) 111061

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A low velocity impact study on press formed thermoplastic honeycomb sandwich panels

T

Manoja Rao Yellur , Holger Seidlitz, Felix Kuke, Kevin Wartig, Nikolas Tsombanis ⁎

Department of Lightweight Design with Structured Materials, Brandenburg University of Technology Cottbus-Senftenberg, Cottbus D-03046, Germany

ARTICLE INFO

ABSTRACT

Keywords: A. Low-velocity impact B. Sandwich panel C. Honeycomb core D. Finite element analysis (FEA)

At present plywood structures are used in the loading area of utility structures. Low velocity impact studies on these structures showed cracks on its lower surface. Hence, in the current study low-velocity impact of a lighter honeycomb sandwich structure is investigated to satisfy the needs of the utility vehicle segment. To meet this objective, facing sheets are manufactured using the polypropylene matrix and glass fibers. Polypropylene honeycombs are used in the study. Depending on the experimental boundary conditions, a cross-ply laminate set up is used for the facing sheets. An impact energy of 100 J is chosen in the study. This energy caused visible failure on the plywood sample. Hence a lighter sandwich construction which can resist 100 J impact is implemented in this study. Influence of top and bottom facing sheet thicknesses on the amount of damage inflicted on its surfaces are studied. Experimental histories of absorbed energy and contact force are recorded. A finite element analysis is performed using LS-DYNA and numerical results are compared with the experimental responses. A honeycomb sandwich panel [0/90/90/0/Core/0/90/90/0] meeting the objective of the study is seen as an optimum replacement for the existing plywood structures.

1. Introduction Automotive manufacturers are motivated by government regulations to produce environmentally friendly vehicles. Hence lighter FRP composites, which help in reducing carbon emission are widely used in vehicle sector [1,2]. Due to high stiffness/strength to weight ratios along with improved stability for buckling and torsional loads especially composite honeycomb sandwich structures have found greater importance in transportation sector [3,4]. Sandwich construction comprises of outer facing sheets with high stiffness and a lighter core having a lower modulus of elasticity. Despite its stability and flexural performance, various studies have showed honeycomb sandwich structures are sensitive to low velocity impact damage caused by various environmental factors. Foreign subjects such as debris, hailstones could be one among the several factors causing impact damage [5,6]. Damages caused by the impact can be of the form of matrix cracking, fiber failure, delamination of facing sheet and crushing failure of the core. Previous research work has shown internal damages are caused at lower energy levels in comparison to energy levels which are required to produce a visible indentation [7]. These damages can significantly lower the load carrying capacity and cause the honeycomb sandwich to lose its integrity. Experimental, numerical and theoretical are general approaches ⁎

used to investigate the low velocity impact response of a sandwich structure [8–10]. Contact force between impactor and honeycomb sandwich structure, energy absorption, deflection are various history variables which are recorded and compared [11–13]. Indentation and energy absorption capability of aluminum sandwich panels subjected to low velocity impact by a spherical impactor were studied by D. Zhang [14]. Refs. [3,4] highlighted various analytical solution methods such as spring-mass model, modal superposition and energy-balanced model used for various impact responses such as boundary controlled, small impactor mass, large mass impact respectively. Numerical techniques using finite element strategy is also one of the popular approach to investigate impact problems as it helps to save time and cost. C. Menna followed a numerical approach using FE-code LS-DYNA to assess the impact behavior of honeycomb sandwich structures and found good agreement between numerical and experimental results [10]. The facing sheets in the sandwich panels can be made of aluminum alloy or a fiber reinforced plastic. The different material choice for the sandwich core include balsa woods, corrugated sheets, honeycombs and cellular foams [15]. K.R. Ramakrishnan investigated medium velocity impact tests on sandwich panels with different cores and found polypropylene honeycomb as an optimum choice if minimum deformation on the bottom facing sheet is desired [16]. A. Dogan carried experiments to study low-velocity impact behavior of E-glass reinforced thermoset,

Corresponding author. E-mail address: [email protected] (M.R. Yellur).

https://doi.org/10.1016/j.compstruct.2019.111061 Received 26 September 2018; Received in revised form 6 November 2018; Accepted 27 May 2019 Available online 29 May 2019 0263-8223/ © 2019 Elsevier Ltd. All rights reserved.

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thermoplastic based sandwich composites and found thermoplastic facing sheets have better load carrying and deformation capacity [17]. Apart from the material used to manufacture sandwich panels, test parameters such as impactor diameter, velocity of impact and geometric properties such as thickness of facing sheets, core have an influence on the impact behavior [18–21]. E.J. Herup and A.N. Palazatto performed low velocity impact and static indentation tests to characterize damage initiation as a function of face sheet thickness and loading rate for 4–48 ply graphite/epoxy cross ply laminate face sheets and Nomex honeycomb cores [19]. Y. Shen carried low velocity impact experiments on sandwich panels made of cross-ply [0°/90°] glass fiber reinforced epoxy skins and aluminum honeycomb core to study the geometrical effects such as position of simple-supports and thickness of honeycomb core [20]. O. Ozdemir found the effect of core thickness as one of the deciding parameter to determine the behavior of sandwich structures. In his study he found an increasing trend in the energy absorption of the sandwich panels for increasing core thickness [21]. Plywood structures are widely used in loading area of the utility vehicles. During their operation in different environmental conditions, these structures are vulnerable to low velocity impact damages. Though there are substantial literature based on low velocity impact studies for sandwich panels, very few have concentrated on materials involving glass fiber reinforced polypropylene facing sheet in a cross-ply laminate set-up and a thermoplastic based polypropylene honeycomb. In the current study, focus is on using the above-mentioned materials for a lighter sandwich construction in comparison with the existing plywood structures and at the same time improved impact response. Impact and damage behavior are studied using a drop tower impact test.

as a starting point for the trial. A temperature of 190 °C and time of 40 s are chosen for the initial trial based on typical compression moulding cycle times for the facing sheet material. In the trials to determine the optimum parameters, a constant temperature of 190 °C is used throughout and only Fpress and t are chosen as variables. Fig. 2 show the honeycomb samples produced for varying Fpress and a constant cycle time of 40 s. From these trials an improvement in the quality of produced sample can be observed when Fpress was lowered. But even for a Fpress as low as 5 kN there were damages on the produced sample. This is because, the compression strength tests for the honeycomb cores were carried out at room temperature and hence the estimated press-force when used in environment higher than room temperature always damaged the core structure. Based on the observation from the previous three trials, further process trials were carried at zero press-force i.e. a position controlled pressing technique was adopted instead of force controlled pressing technique. In the Fig. 3 shown are the honeycomb samples produced for varying cycle times and zero press-force. Damages were found on sandwich samples produced for a cycle time of 40 s and 120 s. An optimum cycle time was found to be 50 s as seen from the quality of the sample in Fig. 3(c). From the trials, press-force of 0 kN, time of 50 s and temperature of 190 °C were identified as optimum parameters and these provided best results in terms of surface finish and bonding for the sandwich construction. During the process, the composite prepregs were thermally bonded to the honeycomb cores and an extra adhesive was not required. Sandwich panels of size 300 × 300 mm were used in the study. The 0-degree fiber direction and panel dimensions are illustrated in Fig. 4. Impact behavior for a sandwich set-up of (0°/90°/Core)S , (0°/90°/0°/ Core)S and (0°/90°/90°/0°/Core)S is investigated.

2. Materials and methods

3. Experimental Set-up

Birch plywood used in the study is obtained from a finnish manufacturer. Test samples of plywood measured 300 × 300 × 15.8 mm along its width, breadth and thickness respectively. Outermost veneer layers belonging to the test sample have grains oriented along 90° directions and inner veneer layers have simultaneous 0°/90° orientation pattern as shown in Fig. 1. The total thickness of the sample is 15.8 mm. Outermost veneer layers are 0.8 mm thick and thickness of each inner layer is around 1.56 mm. Unidirectional glass fiber polypropylene matrix with a fiber volume content of 35% is used for the sandwich skins. The thickness of each ply is 0.25 mm. By taking account of the type of loading and support boundary conditions, a cross-ply laminate [0/90] layup is examined. Thermoplastic polypropylene honeycomb cores obtained from ThermHex manufacturer are used. The density of the core is 80 kg/m3 and it has a cell width of 8 mm. As mentioned earlier a core height of 10 mm is chosen for the study. Press forming technique is used to manufacture the sandwich panels using the mentioned facing sheet and core material. Optimum process parameters such as press-force (Fpress), temperature (T) and time (t) are identified based on several trials. Based on honeycomb’s compressive strength of 1.28 MPa, the maximum force that a honeycomb of 300 × 300 mm size can withstand is evaluated as 115 kN. Hence the range for Fpress was set below 115 kN and a value of 80 kN was chosen

Low velocity impact is investigated using INSTRON CEAST 9350 drop tower impact system as shown in Fig. 5(a). A mass of 10 kg is dropped from a height of 1 m in the study. A spherical impactor with a diameter of 50 mm is chosen (Fig. 5(b)). A shaft is used to allow only a vertical movement for the impactor. After the first rebound the impactor is prevented from re-impacting the panel using a holding frame. Based on the dropping mass and the fall height, impacting energy of 100 J was evaluated by equating potential energy to kinetic energy. Hollow rectangular metal fixture of 40 mm height is used to support the sandwich panel (Fig. 5(c)). This support arrangement on all four sides of the sandwich panel helped to investigate the impact behaviour for a free edge boundary condition. 4. Results and discussion A numerical study is carried out to compare the results obtained through experiments. FE-code LS-DYNA is used to solve the low velocity impact problem. HyperMesh V14 and LS-PrePost are used as a preprocessor and post-processor respectively. Shell-Solid approach is incorporated to model the sandwich panel [22]. The facing sheets are modelled with full integrated shell elements and reduced integrated solid elements are used for the honeycomb core. The bonding between the facing sheets and the honeycomb core are assumed to be perfect and hence a tie-based contact formulation is used to hold them together.

Y (90°)

4.1. Glass-fiber polypropylene sheets (GF-PP)

X (0°) 15,8 mm

To model the energy balance between impacted panels and falling mass an accurate material model is required for both skin and the core. Different damage modes such as fiber failure, inter fiber failure, delamination between plies contribute to energy absorption. Extensive effort is required to characterize the damages modes for the composite skins and LS-DYNA offers variety of material models to model damage

Z

(90°/0°/ 90°/ 0°/90°/0°/ 90°/0°/90°/0°/90°) Fig. 1. Orientation of the veneer layers in the investigated plywood sample. 2

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(a)

(b)

(c)

Damage

Fig. 2. Quality of the sandwich sample produced for a cycle time of 40 s and varying press-force: a) Fpress = 80 kN b) Fpress = 40 kN c) Fpress = 5 kN.

a)

b)

c)

Fig. 3. Quality of the sandwich sample produced for zero press-force and varying cycle time: a) t = 40 s b) t = 120 s c) t = 50 s.

based on either progressive failure or continuum damage mechanics approach. However, in this study the unidirectional composite skins are modelled as shell elements and hence delamination aspect of damage is not considered. This macro approach will require minimum engineering effort and suffice the purpose in the current study. To model the composite skin MAT_ENHANCED_COMP OSITE_DAMAGE_054 is used. PART_COMPOSITE card is used to define the layup for the skins. Transverse isotropic behaviour of the composite skins is described by elastic constants EA, EB, GAB and υBA (minor Poisson’s ratio) respectively. The subscripts A and B corresponds to fiber and matrix direction respectively. The mechanical properties of

the composite skin are shown in Table 1. The tensile and compressive failure modes of fiber and matrix are based on Chang-Chang criteria. In the MAT_054 material model, the part can suffer failure in four different ways described below [23]:

• If maximum fiber strain to tension is zero in the material card, • 3

failure occurs when Chang-Chang failure criterion in met in the tensile fiber mode. If maximum strains to failure for fiber and matrix mode are defined in the material card, failure occurs when elemental strain exceed defined maximum strains.

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10 mm 300 mm

Z

Y (90°) X

300 mm X (0°)

Fig. 4. Honeycomb sandwich test coupon used in the study.

• If effective failure strain (EFS) is defined in the material card, failure occurs if effective strain exceed EFS. • If minimum time step size is defined for element deletion, failure

Table 1 Mechanical properties of GF-PP used in the study as given by the manufacturer.

occurs when the timestep is smaller than the defined value.

In the present study, the maximum strains are chosen higher than the failure strains provided in data sheet so that the Chang-Chang failure criterion is not affected. According to this failure criterion, elastic properties EA, EB, GAB and υBA are degraded for the tensile fiber mode, which is one of the frequently observed mode of damage in low velocity impact investigations. 4.2. Thermoplastic honeycomb core

Property

Value

Density Youngs’s modulus, EA Youngs’s modulus, EB In plane shear modulus, GAB Minor poisson’s ratio, υBA Tensile strength fiber direction, XT Compressive strength fiber direction, Xc Tensile strength matrix direction, YT Shear strength, Sc

1.5 g/cm3 28,000 MPa 3720 MPa 1390 MPa 0.0505 720 MPa 366 MPa 11 MPa 19 MPa

possible to simulate the discussed deformation phases. The model behaves in an orthotropic manner until compaction, where the components of stress tensor are uncoupled (Poisson’s ratio ≈ 0). The elastic moduli vary from the initial values to fully compacted values linearly with relative volume. Nonlinear material responses for normal and shear loads can be defined separately based on the experimentally determined stress strain curves. Compression tests in L, W, T directions and shear tests in LT, WT, and LW plane must be carried out to characterize the discussed nonlinear response. However, in the current study only the compression test in T direction is performed and other five curves are assumed to be scaled version of test in T direction [25].

When a honeycomb core is compressed in transverse direction, it experiences deformation in three distinctive phases as shown in Fig. 6 [24]. The first phase is linear elastic phase, where modulus for the uncompressed phase is obtained. Once the crushing strength is reached a transformation to volumetric crushing occurs. The last phase is the hardening phase until full compaction of the honeycomb structure. Using the slope of the last deformation phase, the modulus of compacted honeycomb material is evaluated. As stated earlier, in the current study the honeycomb core is considered as a homogeneous continuum and modelled using MAT_HONEYCOMB_026 available in LS-DYNA. Using this model, it is

Support fixture

1000 mm 40 mm

Spherical impactor (Diameter = 50 mm)

a)

b)

c)

Fig. 5. Impact test arrangement: a) Drop tower impact system b) Spherical impactor c) Support fixture. 4

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Force/Stress

L Crush Strength

Efully compressed W T

Euncompressed Displacement/Strain Fig. 6. Deformation phases of honeycomb subjected to transverse loading [24].

The relative volume at full compaction is obtained from the test curve by looking for the sudden change in slope, where there is a consistent increase in force values for minimal or no increase in displacement or volume [26]. Effect of strain rate on deformation is not considered in the study. Since off-axis loading was not the topic of concern in the current study, MAT_MODIFIED_HONEYCOMB_126 was not used. The input parameters used in the material model are shown in table 2. The curves LCA, LCB, LCAB, LCBC and LCCA are scaled versions of curve LCC obtained using appropriate scaling factors as mentioned in table 2.

Input curve: MAT_026

Stress [MPa]

3 2.5 2 1.5 Average_Response

1

Input_Curve

0.5 -0.2

0

0

0.2

0.4

0.6

0.8

1

Volumetric Strain [%]

4.2.1. Quasistatic compression test – honeycomb core Flatwise compression tests are carried out as per DIN 53,291 to determine the compacted and uncompacted core properties for the MAT_026 material model in LS-DYNA. Test specimens of dimensions 60 × 60 × 28 mm are prepared and compressed with a velocity 0.5 mm/min. Five specimen are tested and an average response of five tests is chosen to evaluate the input load curve for the material card as shown in Fig. 7. The input stress-strain curves are based on volumetric strains and the curve begins with a positive stress for negative strain as recommended in [21]. All input curves for the material card contains same number of data points as suggested in [21]. The curve shown in Fig. 7 represents load curve LCC used in the material card. To validate the tests the experimental setup is replicated using brick elements of 4 mm size. Rigid material is used for load plate and for the bottom

Fig. 7. MAT_HONEYCOMB_026 Input curve.

Load plate

Honeycomb Z

28 mm Support plate

X Fig. 8. Compression test simulation set-up.

Table 2 Honeycomb material properties used in the study. Notation

Description

Value

RO E PR SIGY VF MU BULK LCA LCB LCC LCS LCAB LCBC LCCA LCSR EAAU EBBU ECCU GABU GBCU GCAU

Material density Young’s modulus of fully compacted honeycomb Poisson’s ratio Yield stress of fully compacted honeycomb Relative volume at which honeycomb is fully compacted Damping coefficient Bulk viscosity flag Load curve id for stress sigma-aa versus volumetric strain Load curve id for stress sigma-bb versus volumetric strain Load curve id for stress sigma-cc versus volumetric strain Load curve id for shear stress versus volumetric strain Load curve id for stress sigma-ab versus volumetric strain Load curve id for stress sigma-bc versus volumetric strain Load curve id for stress sigma-ca versus volumetric strain Load curve id for strain rate effects (optional) Elastic modulus Eaau in uncompressed configuration Elastic modulus Ebbu in uncompressed configuration Elastic modulus Eccu in uncompressed configuration Shear modulus Gabu in uncompressed configuration Shear modulus Gbcu in uncompressed configuration Shear modulus Gcau in uncompressed configuration

80 kg/m3 400 MPa ≈0 1.22 MPa 0.17 0.05 0.0 SF = 0.1 SF = 0.1 SF = 1 Not used in study SF = 0.53 SF = 0.53 SF = 0.53 Not used in study 4 MPa 4 MPa 40 MPa 7.5 MPa 15 MPa 15 MPa

5

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b) Undeformed

a) Force-Displacement curve 10500

Force [N]

9000 7500 6000 4500

EXP

3000

NUM

1500 0

c) Deformed 0

10

20

30

Displacement [mm] Fig. 9. Compression test: a) Comparison of force–displacement response b) Undeformed state c) Deformed state.

b) Energy vsTime Energy [J]

Force [N]

a) Force vs Time 12000 10000 8000 6000 4000 2000 0

EXP

0

20

120 100 80 60 40 20 0

EXP

0

Time [ms]

5

10

15

Time [ms]

c)

d)

Indentation

Crack

Fig. 10. a) Contact force history b) energy history from impact test on plywood sample c) Indentation on impacting side d) Cracks on non-impacting side.

Supporting fixture

Spherical impactor

Fig. 11. Impact simulation set-up. 6

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a) Force vs Time

b) Energy vs Time 150

4000 3000

EXP NUM

2000

Energy [J]

Force [N]

5000

100 EXP NUM

50

1000 0

0

5

0

10 15 20 25 30 35 40

0

10

20

30

40

Time [ms]

Time [ms]

Fig. 12. Comparison of experimental and numerical results for (0/90/Core)S lay-up: a) Force history b) Energy history.

a) Fiber damage plot: E-23287 1

Fiber Damage

0.8 0.6

Damage_IPT1[90°][HV1]

0.4 Damage_IPT2[0°][HV1]

0.2 0

0

5

10

15

20

25

30

35

40

Time [ms]

c)

b)

Fig. 13. (0/90/Core)S lay-up; (a) Fiber damage plot of element 23,287 (b) Fiber damage contour plot for the 0° layer from the simulation (c) Damaged test sample after the end of the test.

support. The load plate shown in Fig. 8 is constrained to move only in Z direction and the support plate is fixed in all direction. Stiffness form hourglass controls are used to avoid spurious nonphysical deformation modes. The simulation is displacement controlled and the cross-section forces are extracted using appropriate keywords in LS-DYNA. The numerical force-displacement response shown in Fig. 9(a) is in good agreement with the experiment and all the three deformation phases discussed earlier are well captured. 4.3. Impact investigations

3 ms. Then there is a sudden drop in the contact force, which indicates a crack initiation at the non-impacting side as shown in Fig. 10(d). The ball attains positive velocity after the absorbed elastic energy is recovered by the plywood samples in the next few milli seconds and rebounds back. The impacting side suffers a plastic dent as shown in Fig. 10(c) and there are no visible cracks in this side. Impact test on two specimens showed identical damage response and hence no further samples were impacted. Hence in the further impact investigations involving sandwich samples, this behavior is set as a constraint, which lets the sample to deform plastically on the impacting side, but without any visible cracks on either side of the test sample.

4.3.1. Plywood sample A 100 J impact test is performed on plywood samples as explained in Section 2 of this paper. The impact performance of this lay-up is investigated to compare it with similarly laid plywood structures used in the loading area of the utility vehicles. Contact force and energy history curves from the experiments are shown in Fig. 10. The Initial peak of 10,000 N is observed at around

4.3.2. Thermoplastic honeycomb sandwich sample Low velocity impact investigations are carried out on thermoplastic honeycomb sandwich panels as illustrated in section 3. To reproduce the experimental set-up, the spherical impactor is modelled as a ball with a diameter of 50 mm. A hollow rectangular solid is used to replicate the supporting fixture used in experiment. The 7

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a) Force vs Time

b) Energy vs Time

8000

4000

EXP

2000 0

Energy [J]

Force [N]

6000

NUM 0

5

10

15

20

120 100 80 60 40 20 0

EXP NUM 0

5

Time [ms]

10

15

20

Time [ms] c)

Crack Fig. 14. (0/90/0/Core)S lay-up: (a) Force history (b) Energy history (c) Crack observed on top facing sheet.

a) Force vs Time

b) Energy vs Time 150

10000

Energy [J]

Force [N]

8000 6000 4000

EXP

2000

NUM

0

0

5

10

15

100

0

20

Time [ms]

EXP

50

NUM 0

5

10

15

20

Time [ms]

Fig. 15. (0/90/90/0/Core)S lay-up: (a) Force history (b) Energy history.

a)

b)

c)

d)

Impact simulation arrangement is shown in Fig. 11. In the conducted experiments, no significant deformation was observed on the impactor and supporting fixture. Hence, they are modelled as rigid solid elements. ELEMENT_MASS_PART card available in LS-DYNA was used to add mass to the impactor to match the drop mass of 10 kg used in experiment. The impactor was assigned with an initial velocity of 4.42 m/ s, which corresponded to a kinetic energy of 100 J. The motion of the impactor was constrained in all direction except in Z-Axis. The movement of the rigid supporting fixture was fixed in all the directions. FEmodel used in the study comprised of 47,547 solid elements and 11,250 shell elements. Uniform mesh was used due to reasonably shorter simulation time. With the available material data for the study, a compromise between the numerical and experimental results was possible for this mesh size. AUTOMATIC_SURFACE_TO_SURFACE contact behavior was used for interfaces between impactor-sandwich panel and supporting fixture-sandwich panel. Force histories and energy histories from both numerical and experimental study are presented. In Fig. 12 the experimental and numerical histories for (0/90/Core)S lay-up configuration is shown. The recorded values highlight a reasonable agreement in terms of peak load as shown in Fig. 12(a). Peak load predicted in the numerical study is 4700 N, which is slightly higher than 4190 N as observed in experiment. The main mode of damage is determined as fiber failure which is

No visible damage Small Dent Fig. 16. (0/90/90/0/Core)S lay-up: (a) fiber failure contour-impact side (b) fiber failure contour- non-impact side (c) Indentation on impact side (d) Nonimpact side after impact.

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b)

a)

[0/90/Core]S

[0/90/0/Core]S

c)

[0/90/90/0/Core]S

Fig. 17. Effective strain contour plots of honeycomb core for the investigated sandwich panels.

confirmed from the fiber damage plot (History variable1-HV1) for the element 23287, which is first element to meet the failure criteria as shown in Fig. 13(a). Element is deleted when all the Integration point (IPT) associated with it meet the defined failure criteria in the material card. In the PART_COMPOSITE card lay-up is defined starting from the bottom most integration point. Hence IPT-1 corresponds to 90° and IPT2 to 0°. As observed in Fig. 13(a) failure is initiated in the 0° layer at around 6.3 ms which can also be confirmed from the force histories, where the curve tends to drop down due to the suffered failure. Fiber damage contour plot for the first layer on the impacting side i.e the 0° layer is shown in Fig. 13(b) which is comparable with the visible cracks observed on the top facing sheet in the experiment (Fig. 13(c)). In the contour plot shown, value of 1 signifies that the element is still elastic where as a zero value indicate a failed status for the element. Investigation of the numerical traces in the latter phase of the impact i.e after the linear peak shows slight discrepancy with the recorded experimental data. The strain energy in the simulation is also slightly lesser compared to the test impact energy of 100 J as shown in Fig. 12(b). This could be due to the accuracy of the damage model used in the study, which was set-up with the minimum available material data from the manufacturer. Fig. 14 show the force and energy histories for the (0/90/0/Core)S lay-up used in the study. Linear peak of 6500 N was observed in simulation in comparison to 7400 N observed in experiment. Agreement between simulated and experimental energy curve is better in comparison to the first variant. Amount of damage or failure inflicted on to the facing sheets in this case is lesser as compared to first case, which is evident from the shorter crests and troughs observed after 6 ms in the simulation. An observation of experimental force curve does not signify a failure as it was difficult to spot any force drop and the contact force slowly reduces as the ball recedes back after impact. However, a visual inspection showed crack on the top facing sheet which supported the results from the numerical study. For the lay-up configuration (0/90/ Core)S and (0/90/0/Core)S only one sample was tested as for both the lay-up’s the test sample could not resist an impact of 100 J. Shown in Fig. 15 are the force and energy histories for the final layup [0/90/90/0/Core]S investigated in the study. Force history follow a near sinusoidal pattern as observed in experiment. However, an initial spike in the force of around 400 N captured in the very initial phases of the experiment [0–0.1 ms] could not be exactly replicated in the simulation. Energy history also exhibit similar trend except that the maximum strain energy is 91 J in comparison to an input kinetic energy of 100 J. There are no visible cracks on the tested sample except a very small residual dent on the impacting side as observed in the post impact pictures shown in Fig. 16(c) and (d). The fiber damage contours for the first layer in the impacting side (Fig. 16(a)) and outermost layer in the non-impacting side (Fig. 16(b)) also show very minimum levels of damage on the face sheets. Three samples were tested for this lay-up to ensure that the responses were repetitive. For the first two lay-up configuration of the sandwich panel, the

deformation of honeycomb is higher, which causes larger deflection on the facing sheet thermally bonded to the honeycomb, subsequently leading to its failure. In Fig. 17, the effective strain contours of the honeycomb core for all the three lay-up’s investigated are shown and it can be clearly observed that larger strains are present for the first two variants. 5. Conclusions Low-velocity impact performance of a thermoplastic based honeycomb sandwich panel for three different facing sheet thickness is compared with the existing plywood material used in the loading area of utility vehicles. Numerical investigation was carried out using LSDYNA. Contact force and energy histories were compared with test results. When impacted, plywood samples could not resist an energy of 100 J and cracks were observed on the non-impacting side. Three variants of sandwich panel, [0/90/Core]S, [0/90/0/Core]S and [0/90/90/ 0/Core]S were prepared and tested for 100 J impact. Numerical investigation carried out to validate the experimental observation confirmed the failure on the top facing sheet for the first two variants. Final variant [0/90/90/0/Core]S , in a symmetrical cross-ply layup provided the desired impact performance in comparison with its impacted sandwich counterparts and the existing plywood sample. Apart from improved impact performance, close to 60% advantage in weight was achieved for the sandwich construction in comparison to plywood. Numerical study agreed well in terms of peak load. But for the first two variants, the latter phase of impact after the peak load had satisfactory agreement with the test data. This can be due to accuracy of the material model in the vicinity of failure. Fiber failure was identified to the main failure mechanism for the variants studied. The level of agreement between experimental and numerical observation illustrated the approach followed in the study as satisfactory and highlights the need of extensive testing to characterize damage in the composite skins. Appendix A. Supplementary data Supplementary data to this article can be found online at https:// doi.org/10.1016/j.compstruct.2019.111061. References [1] Gerstenberger C, Osiecki T, Kroll L, Scholz P, Seidlitz H. Processing and characterization of cathodic dip coated metal/composite-laminates. Arch Civil Mech Eng 2016;16:467–72. [2] Santhanakrishnan Balakrishnan V, Seidlitz H. Potential repair techniques for automotive composites: a Review. Compos Part B: Eng 2018;145:28–38. [3] Mackerle J. Finite element analyses of sandwich structures: a bibliography (1998–2001). Eng Comput 2002;19:206–45. [4] Abrate S. Impact on composite structures. Cambridge University Press; 1998. [5] Chai GB, Zhu S. A review of low-velocity impact on sandwich structures. Proc Inst Mech Eng, Part L: J Mater: Des Appl 2011;225:207–30. [6] Ramakrishnan KR, Guérard S, Viot P, Shankar K. Effect of block copolymer nano-

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