Scholarly article on topic 'Improvement of the in-plane crushing response of CFRP sandwich panels by through-thickness reinforcements'

Improvement of the in-plane crushing response of CFRP sandwich panels by through-thickness reinforcements Academic research paper on "Materials engineering"

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Composite Structures
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{"Energy absorption" / "Impact behaviour" / "Through-thickness reinforcement" / "Composite sandwich panel"}

Abstract of research paper on Materials engineering, author of scientific article — L.G. Blok, J. Kratz, D. Lukaszewicz, S. Hesse, C. Ward, et al.

Abstract Fibre reinforced plastic (FRP) composite materials can provide superior specific energy absorption performance over conventional metallic structures if crush stability can be maintained during the impact event. The core in sandwich structures helps to stabilise the crush front by preventing global buckling, but delamination remains a barrier to optimal crushing performance. In this work, the in-plane crushing response of sandwich structures was improved by adding through-thickness reinforcement in the form of aramid fibre tufts. The effect of tufting different sandwich cores and facesheet orientations was investigated in both static and dynamic crushing modes. A drop-tower test rig was used to crush panels in realistic automotive crash conditions and a high-speed camera captured the crushing mechanisms. The through-thickness reinforcement improved the facesheet to core adhesion, resulting in a more localised and stable fracture of the facesheets. Tufting improved the specific energy absorption (SEA) from 11.5kJ/kg to 20.5kJ/kg and the crush force efficiency (CFE) from 0.22 to 0.55.

Academic research paper on topic "Improvement of the in-plane crushing response of CFRP sandwich panels by through-thickness reinforcements"



Accepted Manuscript

Improvement of the in-plane crushing response of CFRP sandwich panels by through-thickness reinforcements

L.G. Blok, J. Kratz, D. Lukaszewicz, S. Hesse, C. Ward, C. Kassapoglou

S0263-8223(16)31420-9 COST 7994

To appear in:

Composite Structures

Received Date: Revised Date: Accepted Date:

I August 2016

10 November 2016

II November 2016

Please cite this article as: Blok, L.G., Kratz, J., Lukaszewicz, D., Hesse, S., Ward, C., Kassapoglou, C., Improvement of the in-plane crushing response of CFRP sandwich panels by through-thickness reinforcements, Composite Structures (2016), doi:

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L.G. Blok12, J. Kratz3a, D. Lukaszewicz2, S. Hesse2, C. Ward3, C. Kassapoglou1

Faculty of Aerospace Engineering, Delft University of Technology, Kluyverweg 1, 2629 HS Delft, The Netherlands

BMW Group, Department, Research and Innovation Centre, Knorrstrasse 147, 80788 Munich, German

ACCIS, University of Bristol, Queen's Building, University Walk, BS8 1TR, Bristol, United Kingdo

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Fibre reinforced plastic (FRP) composite materials can provide superior specific energy absorption performance over conventional metallic structures if crush stability can be maintained during the impact event. The core in sandwich structures helps to stabilise the crush front by preventing global buckling, but delamination remains a barrier to optimal crushing performance. In this work, the inplane crushing response of sandwich structures was improved by adding through-thickness reinforcement in the form of aramid fibre tufts. The effect of tufting different sandwich cores and facesheet orientations was investigated in both static and dynamic crushing modes. A drop-tower test rig was used to crush panels in realistic automotive crash conditions and a high-speed camera captured the crushing mechanisms. The through-thickness reinforcement improved the facesheet to core adhesion, resulting in a more localised and stable fracture of the facesheets. Tufting improved the specific energy absorption (SEA) from 11.5 kJ/kg to 20.5 kJ/kg and the crush force efficiency (CFE) from 0.22 to 0.55.

Keywords: energy absorption, impact behaviour, through-thickness reinforcement, composite sandwich panel

1. Introduction

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Fibre reinforced plastic (FRP) composite materials are gaining interest from car manufacturers for many reasons. In addition to higher stiffness, fibre reinforced composites can provide superior energy absorption performance over conventional metallic structures, when compared by weight [1]. The use of composite materials as energy absorbers, however, is still limited as they crush in a more unstable manner than their metallic counterparts [2]. In a real-world crash situation, load scenarios are not well defined and this restricts the design of energy dissipating composite structures to beams and profiles at the moment [3]. Composite sandwich structures may provide a more stable crushing mechanism for fibre reinforced materials, enabling new vehicles with flat, shell-line structures capable of absorbing energy in multiple load paths in a crash event.

FRP composite materials typically crush in a brittle manner, which can result in unstable catastrophic failure, but if properly controlled, stable progressive end-crushing is possible. Progressive crushing is characterised by a localised zone of micro fracturing of the composite erial consisting of splaying/lamina bending and fragmentation/fibre fracture [1] [4] [5]. Several studies have investigated the change in energy absorption of FRP crush tubes using interlaminar toughening techniques [6], surface treatments [7], different matrix properties [8], [9] and different fibre orientations [10]. These studies showed that higher interlaminar properties lead to less splaying/lamina bending, resulting in higher energy absorption.

Most studies investigating impact events to sandwich structures have focused on the out-of-plane impact loading [10] [11], commonly related to bird strike in aerospace applications. The inplane (or edgewise) crushing of sandwich structures has received less attention in the literature, but

a Corresponding author email:, Tel.: +44 117 33 15775

becomes more relevant for automotive crashworthiness. The study by Mamalis et al. [12] reported three different collapse modes for edgewise compression of composite sandwich structures: unstable sandwich column buckling (Mode 1); unstable sandwich disintegration by facesheet/core disbonding (Mode 2); and progressive end crushing of the sandwich (Mode 3). They identified that the mechanical properties of the core have a large effect on the failure mode, as the core stabilises the facesheets during crushing and prevents sandwich disintegration [11], [12].

Introducing a foam core may stabilize the crushing of the fibre layers, but this adds another interface to the structure which can disbond, resulting in unstable collapse. Stapleton and Adams [13] investigated the effect of crush initiators to promote stable crushing of sandwich panels. They propose that a stable crushing mode depends on the strength and stiffness of the core, the composite facesheet, and the strength of the facesheet/core interface [13]. However, increasing the stiffness and strength of the facesheets does not necessarily improve the energy absorption as facesheet buckling failure may occur, reducing the energy absorbed during the impact event. More recently, they reported on structural enhancements (end-bevel, stitching and core webbing) to improve the energy absorption [14]. Their findings suggest that stable sandwich crushing may be realised by improving the facesheet/core interface.

To successfully implement the superior energy absorption of composites, a stable and predictable crushing platform is required. In this work, an improvement in crushing response is obtained by through-thickness reinforcements of the facesheets to the core using aramid fibre tufts. This effectively creates a three-dimensional composite structure which have been proposed to improve the impact resistance [15], [16]. The advantage of the tufting technique is that it only requires one-sided access to the preform. These tufts strengthen the interface between the facesheet and the core and prevent premature facesheet separation [17],[18]. At the same time, the tufts prevent the facesheet plies from delaminating from eachother, further increasing the energy absorbed during the impact event.

Static and dynamic edgewise compression tests were carried-out on eight different panel configurations to identify whether core type and facesheet fibre orientation influence the energy absorbing characteristics. A customised drop-tower test-rig was developed to allow facesheet/core failure, meanwhile a high speed camera was used to record the crushing front. The different crushing mechanisms offer detailed understanding of the failure progression in through-thickness reinforced sandwich structures subjected to edgewise loading.

2. Methodology 2.1 Materials and manufacturing

Sandwich specimens were manufactured using two common foam cores types: Rohacell 110 IG-F a polymethacrylimid (PMI) and Airex C70.90 a polyvinylchlorid (PVC), both 10 mm thick. The main difference between the two foam cores is the maximum strain to failure of 3% versus 23%, as shown in Table 1. The facesheets were made from uniaxial carbon fibre non-crimp fabric (NCF) from SGL utomotive with 50K tows, and a 310 g/m2 areal weight. The strength and modulus of the fibres according to manufacturer specifications is 4 GPa and 240 GPa, respectively [19].

Table 1: Mechanical properties polymer foams as supplied by manufacturer [20], [21]

Airex C70.90 Rohacell 110 IG-F

Density [kg/m3] 100 110

Compressive strength [MPa] 2.0 3.0

Compressive modulus [MPa] 130 160

Shear modulus [MPa] 40 50

Maximum strain [%] 23 3

Hand lay-up was used to preform the sandwich panels, after which a tkt 20 Kevlar® thread (from Somac threads UK) was tufted through the preformed sandwich in a 6 mm x 6 mm square pattern using a KSL tufting head mounted to a Kuka robotic arm at the UK National Composites Centre [17]. A vacuum infusion process was used to saturate the dry perform with RIMR 935/936 resin system from Momentive Speciality Chemicals. The plates were infused at 40 oC and then cured at 90 oC for 2 hours under -98.2 kPa (gauge) vaccum pressure. Specimens were waterjet cut from the cured plates and no further maching steps were carried out. Of note, the tufting needle created a void space in the foam core during thread insertion which was then filled during infusion resulting in a resin column around the Kevlar thread, as shown in Figure 1, and the tufted specimens were on average 28% heavier than the untufted control specimens.

Four different lay-ups were used for the facesheet skins; [0/90/0], [0/90/0]s, [-45/0/45] and [-45/0/45]s, yielding eight different sandwich configurations as summarised in Table 2. The thickness of the facesheets was 1.2 mm for the three ply configurations and 2.3 mm for the six ply configurations. A monolithic traveler laminate was tested for fibre volume fraction by acid digestion and a value of 54.5% was measured with a void volume fraction of 1.9%.

Figure 1. Example of tufted sandwich specimen showing resin columns around Kevlar tufts (inset).

Table 2. Sandwich configurations for static and dynamic testing Facesheet material Non-Crimp Fabric

Facesheet lay-ups [0/90/0], [0/90/0]s, [-45/0/45] and [-45/0/45L

Core materials Airex C70.90 and Rohacell 110 IG-F

Through-thickness reinforcement With and without Kevlar® tufts.

2.2 Test fixture

Most studies investigating the in-plane crushing response of sandwich panels use a fixed specimen and a moving impactor [11], [13], [14]. In this study, the specimen was mounted to the moving hammer in a drop-tower rig and crushed into a fixed plate, as shown in Figure 2. The specimen freely falls in between four alignment guides, as in shown in Figure 3. The advantage of the test set-up used here is that specimen rotation/bending is suppressed without constraining disbonding or delamination over its length. Interface failure between the core and facesheet was recorded at the crush front using a high speed camera.

Figure 2. Test set-up showing sandwich crush sp ecimen before and after crushing in test fixture

Constrain global bending and rotation Front view ✓ v Side

separation allowed

Figure 3. Detail of test fixture showing sandwich specimen and four alignment guide rods and impact plate

The test specimen dimensions for the static and dynamic tests are shown in Figure 4. Static tests were smaller as they were carried out for initial screening of the sandwich panels to determine the impact energy for the dynamic tests. Both specimens included a 14° bevel to promote stable progressive crushing as demonstrated by Velecela et al. [22], as shown in Figure 4 . For each sandwich configuration, three static tests and three dynamics tests were performed.

Figure 4. Dimensions of the test specimens for static (left) and dynamic (right)

2.3. Static testing

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The static tests were conducted using a hydraulic press as shown in Figure 5. The test was carried out at a constant displacement rate of 330 mm/min. The displacement and force were directly measured, which could then be used to determine relevant crushing properties such as maximum force, average force and energy absorption.

Figure 5: Sandw

mic testing

(a) (b)

ich specimens in a hydraulic press (a) before and (b) after static testing

R1.5a accc

2.4. Dynamic testi

Dynamic testing was carried out by mounting the sandwich specimen under a drop hammer and dropping the entire assembly onto the crush plate, as shown in Figure 2. The drop height was set to have an impact speed vimp of 8.6 m/s. A mass mimp was added to the drop hammer to change the impact energy Eimp. The impact energy was adjusted using the energy absorbed from the static test asurents in order to crush the specimen without colliding the drop hammer into the test fixture.

Two accelerometers recorded the acceleration of the drop-hammer during specimen shing. The raw accelerometer data was filtered to remove any high frequency oscillation ccording to the SAE J211 guidelines for impact testing [23]. A CFC 180 filter has been used which is the recommended filter to be used for further integration of the acceleration data to obtain velocity and distance.

The crushing force was determined from the filtered acceleration data aacc using equation (1), where mimp is the mass of the drop hammer. The crushing distance was also determined from the filtered acceleration data, by integrating it twice according to equations (2) and (3), where vimp was the initial velocity, s was the distance travelled, and t was the crush time. A representative section of crushing was extracted from each force-displacement curve up to a displacement s = 75 mm to obtain a common basis for all sandwich panels.

^ imp Q

v{t) Vimp I &nrrdt s(f) =

(1) (2)

2.5. Crushing properties

lustrate th

Two force-displacement curves for the dynamic and static tests are shown in Figure 6 to illustrate the typical crushing response measured during testing. The force-displacement curve of the dynamic test (Figure 6a) shows less detail than that of the static test (Figure 6b). The dynamic data has been filtered to remove high frequency noise. The use of filtering prevented visual correlation between the local failure mechanisms and force recorded during crushing. The crushing performance of each sandwich configuration was compared using the global crushing properties indicated in Figure 6. The crushed mass mcrushed needed to determine the specific energy absorption was derived from the crushing distance and the density of the specimen, and the energy absorbed Eabsorbed was calculated by integrating the force-displacement graph. The following paramaters were identified from each test:

• Maximum recorded force during crushing Fmax

• Average crushing force during entire crushing event Favg

• Crush force efficiency (CFE = Favg/Fmax)

• Specific energy absorption (SEA = Eabsorbed/m crushed)

Figure 6: Force-displacement response of C70.90 [0/90/0] sandwich in (a) dynamic and (b) static crushing

3. Results and discussion 1 Maximum force and average crushing force

The maximum force supported by the sandwich specimens occurs at the early stages of -loading, as shown in Figure 6. The sandwich structure elastically compresses under the applied load until damage is initiated at the trigger. The maximum crushing force of the different configurations (three specimens per configuration) is shown in Figure 7. An overall increasing trend was observed in the maximum force by adding the through-thickness reinforcement. The tufts increase the load required to initiate failure of the facesheet-core interface. One exception was observed in the C70.90 [0/90/0]s sandwich system, where a small decrease in maximum force was recorded. This configuration has relatively stiff facesheets, due to the predominantly 0° ply orientation and number of plies, supported by a ductile foam core. In contrast, the 110 IG-F [0/90/0] sandwich panels

showed the highest increase in maximum force, due to the thinner facesheets and brittle foam core providing less support during crushing.

(a) Dynamic

(b) Static

Figure 7: Maximum crushing force during (a) dynamic and (b) static testing

Once damage was initiated in the sandwich specimens, the load supported decreases, as shown in Figure 6. At this point in the impact event is where the through-thickness reinforcement is truly advantageous. The average crushing force in all sandwich configurations was improved by tufting, as shown in Figure 8. In all facesheet lay-ups, the brittle 110 IG-F sandwiches structures shows greater improvement due to tufting compared to the more ductile Airex C70.90 sandwiches. The average crushing force does not take into account the increase in sandwich mass due to tufting, as a result, the SEA and CFE analysis in the following section offers a more direct comparison to the crashworthiness benefits of tufting sandwich structures.

Tuft and

(a) Dynamic (b) Static

Figure 8: Average crushing force during (a) dynamic and (b) static testing

3.2 Specific energy absorption and crush force efficiency

ufting clearly improved the specific energy absorption (SEA) of all sandwich panels under dynamic and static crushing, as shown in Figure 9. The average increase in SEA during dynamic crushing for the Rohacell 110 IG-F sandwiches was 117%, and for the Airex C70.90 sandwiches the average increase was 52%. Furthermore, the SEA during dynamic crushing is consistently 60% lower than in static test conditions. This difference may be attributed to the strain-rate dependency of the material. During dynamic impact, the strain-rate sensitivity of the material system becomes important, which is more dominant for the matrix response and friction forces [24]. In the dynamic tests, the predominant failure mode was delamination which is described in section 3.3.

Figure 9 also shows that the SEA of untufted C70.90 sandwich panels generally exceeded that of the untufted 110 IG-F sandwiches. When the tufted sandwiches are compared, the 110 IG-F

sandwiches show a higher SEA than the C70.90 sandwiches. The 110 IG-F sandwich panels showed a larger relative improvement by tufting than similar C70.90 sandwiches. The crushing response of the untufted sandwich panels tested in this paper are in agreement to those reported by Mamalis et al. [11], where foam cores with higher ductility better stabilise the facesheets during crushing.

The different facesheet lay-ups show how the SEA changes with thicker and/or stiffer facesheets. For untufted sandwiches under dynamic crushing, thicker facesheets ([0/90/0]s and [-45/0/45]s) showed a higher SEA than similar thinner facesheets ([0/90/0] and [-45/0/45]). For the tufted sandwiches, however, the thickness of the facesheets seem to have little effect on the SEA during dynamic crushing.

For each sandwich configuration, three specimens were tested. The coefficient of variation in SEA was 9% on average over all dynamic test specimens and 12.9% for the static test specimens. During crushing, different failures occur progressively over a representative distance, which may reduce the global error as it is the average of all fracture.

(a) Dynamic (b) Static

Figure 9: Specific energy absorption of sandwich specimens in (a) dynamic and (b) static crushing

Figure 10 shows the crush force efficiency (CFE = Favg/Fmax) during static and dynamic crushing for the different sandwich panels. The average improvement by tufting in dynamic crushing of the 110 IG-F sandwiches was +100% (+225% in static crushing) and for the C70.90 sandwiches +58% (+214% in static crushing). This indicates that tufting increased the energy absorption by having a higher sustained crushing force compared to the initial peak load.

(a) Dynamic

(b) Static

Figure 10: Crush force efficiency (Favg/Fmax) during (a) dynamic and (b) static testing

3.3 Crushing mechanisms

A comparison of the crushing mechanisms for the different sandwich panels is shown in Table 3. No unstable sandwich column buckling was observed in these tests, confirming that the trigger was an effective initiator of facesheet/core disbonding and progressive end crushing of the sandwich. A difference in crushing response was observed in the untufted panels, where the core ductility influenced facesheet core separation as reported by Mamalis et al. [11]. The crushing response was improved for both ductile and brittle cores by tufting, which led to more localised progressive end crushing in both instances.

Table 3. Observed Crushing Mechanisms of Sandwich Panels

Non Tufted Tufted

Brittle Rohacell 110IG-F Dominated by separation of the facesheet from the core, and bending of the separated facesheet Localized splaying and fracture of facesheet (Mode 3 progressive crushing)

Ductile Airex C70.90 Dominated by delaminations in the facesheet, core build-up and bending of the delaminated facesheet Localized splaying and fracture of facesheet (Mode 3 progressive crushing)

The main crushing mechanisms observed during edgewise loading and the corresponding force-displacement graphs of the 110 IG-F [0/90/0] and C70.90 [0/90/0] sandwiches are shown in Figure 11 and 12. Facesheet/core interfacial failure was the dominant failure mechanism in the untufted 110 IG-F sandwich panels, and is clearly visible from the start of crushing, as shown in Figure 10 (see arrows A,B and C). The facesheet/core disbonding remained the only failure mode present throughout crushing. When tufting was introduced to the sandwich, the facesheets remained attached to the core and bent through a small radius (arrows D and E) before the Mode 3 fracture processes started.

The failure of the untufted C70.90 [0/90/0] specimen was initiated by delaminations in the facesheets (arrows F and G), with subsequent bending of the facesheets (arrow H). In contrast untufted 110 IG-F sandwich panels, little fracture of the facesheet or disbonding of the facesheet/core was observed. Instead, core debris built up on the crush plate. For the tufted C70.90 specimens, no delaminations within the facesheet were observed. Although some facesheet/core disbonding failure was observed (arrow I), clearly more facesheet fracture (arrow J) was present together with crushing of the core material.

Overall, the untufted C70.90 [0/90/0] specimens had a higher energy absorption than the untufted 110 IG-F [0/90/0] specimens. This was attributed to the extra energy required to delaminate the plies within the C70.90 specimen facesheets, whereas less energy was absorbed by the untufted 110 IG-F specimens because the facesheets simply disbonded from the core and subsequently bent.

Adding the tufts to the sandwich panels introduced a facesheet fracture mechanism to both the C70.90 and the 110 IG-F specimens. The tufted 110 IG-F specimens showed more fracture and fragmenting of the foam core (arrow K) compared to tufted C70.90 specimens, which may have led to more energy absorption. The effect of tufting may therefore be twofold: strengthen the facesheet/core interface and preventing premature delaminations in-between the facesheet plies. This is similar to the finding of Stapleton and Adams [14], who used biaxial Kevlar® stitching to improve the SEA crushing response of woven CFRP/polyurethane sandwiches by 224.5 %. In addition, the failure mechanisms within the tufts are the topic of on-going research [25]. Identifying the individual contributions of the core, facesheets and fracture of the tufting thread may help better understand the crushing process.

30 40 50

Distance [mm]

Figure 11. Comparison of force-displacement curves of Rohacell 110 IG-F [0/90/0] sandwiches (3x tufted & 3x

non-tufted), with typical failure sequence

40 35 30 Z* 25

o 15 LL

10 5 0

.........Si,r—^ Untufted specimens

nV.......... / / \ \ ✓ * \\ A""^ \ N " Tufted specimens i ■ ]■.......................................... ..............

Distance [mm]

Figure 12. Comparison of force-displacement curves of Airex C70.90 [0/90/0] sandwiches (3x tufted and 3x

non-tufted) with typical failure sequence

4. Conclusions

The edgewise crushing response of foam core sandwich panels was investigated in this study by considering brittle and ductile core, facesheet lay-up, and through-thickness reinforcement by tufting. The most significant improvement in edgewise crushing response was obtained by tufting the sandwich panels with aramid fibre threads. Both specific energy absorption (SEA) and crush force efficiency (CFE) increased by an average of 78 %, and 129 %, respectively. The tufted specimens showed more localized and stable fracture of the facesheets, leading to a higher sustained crushing load and as a result better absorbing more energy than their untfuted counterparts.

The non-tufted sandwich structure failure was dominated by facesheet/core disbonding and delaminations within the facesheet plies. The failure mechanism of the untufted structures was governed by the foam core and skin thickness. A ductile core stabilised the facesheets, allowing energy absorption through ply/ply delaimination, whereas a brittle core encouraged facesheet/core disbonding. Accordingly, thicker facesheets with more ply interfaces absorbed more energy, however weak facesheet/core interfaces inhibited fracture of the actual carbon fibres, which is the main energy absorption mechanism activated by tufting the facesheet to the core.

Overall, through-thickness reinforcement by tufting showed the ability to stabilise and tailor the crushing response of fibre reinforced sandwich structres. This may be used to exploit new load-paths in crash structures, and as a result, increase the crashworthiness of composite structures in a larger variety of load cases.


roup for their f

The authors would like to acknowledge BMW Group for their financial support towards this project, providing the testing facilities and the materials. This work is further supported with the help and expertise provided by the Delft University of Technology and by the UK Engineering and Physical Sciences Research Council (EPSRC) under the Impact Acceleration Award scheme at the University of Bristol (EP/K503824/1). The UK National Composites Centre is acknowledged for granting access to the tufting equipment. Finally, we would like to thank Mr. Greissl and Mr. Palmer for help conducting the tests.

Access to supporting data may be requested from the corresponding author, which due to commercial contracts in place, will be subject to consent being granted from the original project participants.

5. Refere

data n n place

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