Design and manufacture of all-PP sandwich panels based on co-extruded polypropylene tapes

Composites: Part B 39 (2008) 1183–1195

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Design and manufacture of all-PP sandwich panels based on co-extruded polypropylene tapes N.O. Cabrera a,b, B. Alcock a,*, T. Peijs a,b a b

Queen Mary, University of London, Department of Materials, Mile End Road, E1 4NS London, UK Eindhoven University of Technology, Eindhoven Polymer Laboratories, P.O. Box 513, 5600 MB Eindhoven, The Netherlands

a r t i c l e

i n f o

Article history: Received 15 December 2005 Received in revised form 21 February 2008 Accepted 10 March 2008 Available online 7 July 2008 Keywords: A. Polymer (textile) fibre D. Mechanical testing A. Honeycomb A. Foam

a b s t r a c t This paper describes the creation of polypropylene sandwich panels, based on all-polypropylene (all-PP) composite laminates combined with a polypropylene based honeycomb or foam core. These all-PP composite laminates are based on high modulus polypropylene tape reinforcing a polypropylene matrix. Sandwich panels containing these all-PP composite laminate faces are compared with sandwich panels containing conventional glass fibre reinforced polypropylene laminate faces, and the mechanical properties, failure modes, and design requirements of these different materials are discussed. Ó 2008 Elsevier Ltd. All rights reserved.

1. Introduction Sandwich panels are relatively common composite structural products and are employed in areas such as sports equipment, and in automotive and aeronautical applications. A typical sandwich panel is composed of three layers, in which two thin sheets (faces) of a stiff and strong material are separated by a thick core of low-density material [1]. Composite laminates are often employed as the faces of a sandwich panel due to their high mechanical performance and low density. Sandwich structures have optimised specific flexural stiffness (flexural rigidity) because the separation of the two faces by a low density core increases the moment of inertia and hence contributes to the flexural stiffness of the panel, with only a small increase in panel mass. In addition to the high stiffness and low density of sandwich panels, an extra advantage of sandwich panels is that the designer can tailor the properties of the sandwich by adjusting the geometrical parameters of the panels (thickness of the core and the faces) independently of the material used. The different material characteristics required of the core and faces of a sandwich panel often demands the use of different materials. This complicates recycling, since recycling can necessitate costly material separation at the end of life of the product, if the constituent materials of the product have different recycling processes. Therefore, the concept of creating a sandwich panel in which the faces and the core are made from the same material is * Corresponding author. E-mail address: [email protected] (B. Alcock). 1359-8368/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesb.2008.03.010

attractive since it would remove this separation step in the recycling process. However, the desire to make a sandwich panel composed only of one material must be balanced with the need for adequate mechanical properties from the final sandwich panel. A candidate material for the construction of sandwich panels is polypropylene (PP). PP is commonly used for a range of applications due its low cost, environmental stability, ease of processing and because it is relatively easy to recycle [2–4]. However, due its modest mechanical properties, PP often needs to be reinforced to satisfy load-bearing applications. Glass fibres are popular reinforcing elements for PP due to their relatively low cost and high mechanical performance. Alternative composite systems reinforced with natural fibres such as flax, hemp and sisal have also been reported [5,6] and may appear to be more environmentally friendly [7]. However, although these natural fibres are renewable and can be incinerated, degradation and heterogeneity of fibres can cause performance issues. Clearly, the introduction of any foreign filler whether a glass or a natural fibre, complicates recycling [8]. However, composite materials based entirely on polypropylene have been developed and reported in literature [9–13]. These ‘‘all-PP” composites show potential for use as faces of sandwich panels. all-PP composites consist purely of a propylene copolymer matrix reinforced with a very high volume fraction of high-performance PP tapes. Since these composites do not contain any foreign reinforcements, recycling can be achieved by simple thermal processing at the end of life of the structure [14], and recycled in a subsequent generation of all-PP composites or simply used as a feedstock for less demanding applications. Because these all-PP composites are entirely thermoplastic, the performance of these

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composites can be expected to vary at different strain rates and temperatures, and this is the focus of a separate publication [12]. This paper describes the development of recyclable sandwich panels based on all-PP composite faces and PP based cores. Flexural testing is used to compare all-PP faced sandwich panels with sandwich panels having glass reinforced PP faces. The bond strength between the face plates and the sandwich core is also assessed. Sandwich panels constructed using the same PP based cores but commercial glass reinforced polypropylene face panels will be presented for comparison. Finally the minimum weight criterion of a sandwich panel with respect to a given flexural stiffness developed by Kuenzi [15] was found appropriate to compare the performance of the different face materials. 2. Designing with all-PP laminates

Table 1 Mechanical properties of materials described in this paper Material

Density (g cm3)

Tensile strength (MPa)

Compressive strength (MPa)

Young’s modulus (GPa)

Aluminium [69] Woven glass (60w%)/PP [70] GMT (30w%) [71] Isotropic PP [69] all-PP [13]

2.7 1.50

100–700 350

160

70 15

1.12 0.90 0.78

67 30 205

50 40

4.0 1.2 5.8

Fig. 1. Examples of materials indices for stiffness or strength limited design at a minimum mass.

100

Youngs Modulus (GPa)

The mechanical properties of all-PP laminates are presented in Table 1, compared to glass fibre reinforced polypropylene (GFRPP). The mechanical properties of GFRPP depend strongly on the architecture of the reinforcement and the volume of glass fibres present. Woven glass fibre reinforced PP (woven glass/PP), such as TwintexÒ, with a glass content of 60 wt% (percentage by weight) woven fabrics of continuous fibre, has considerably higher mechanical properties than glass mat thermoplastics (GMT) of lower glass content and based on random discontinuous glass fibre [16]. In general, all-PP composites have tensile properties which lie in between those of GMT and those of the woven glass/PP. However, the compressive strength of all-PP composites, which is only 20% of its tensile strength, is equivalent to that of isotropic PP. This is because PP fibres, like other polymer fibres such as aramid or polyethylene fibres, are weak in compression [17]. However, the criterion to select the right material for a particular structural component is not just the pure mechanical performance of the material but the optimisation of the performance to weight ratio. Here allPP composites have significant advantages as they have a lower density than isotropic PP due to the microstructure of the tape [13] so that the competitiveness of all-PP composites for structural application is enhanced. Well developed techniques exist to select a material for a given application [18,19]. The mechanical performance of materials for lightweight design can be evaluated and compared by using performance indices (Fig. 1). E=q, E1=2 =q and E1=3 =q where E is the Young’s modulus and q is the density, are the most common indices for stiffness limited design. Materials with maximal value of E=q, will be the best choice for e.g. a tensile strut or a cylinder with internal pressure. Similarly materials of greatest values of E1=2 =q would be suitable for a beam loaded in flexure or a column that fails in compression by elastic buckling. Materials with a high value of E1=3 =q minimise the weight of a plate loaded in flexure or failed by elastic buckling. The equivalent performance indices for strength limited design are r=q, r2=3 =q, and r1=2 =q where r is either the tensile or the compressive strength [20]. These performance indices can also be assessed graphically. Fig. 2 is a logarithmic plot of the Young’s modulus against the den-

10

a al l-PP GMT Woven Glass/PP PP Aluminium

b c 1 0.5

4

1

Density (g.cm-3) Fig. 2. Young’s modulus vs. density of all-PP and glass reinforced polypropylene. The lines represent the main mechanical performance indices for stiffness limited design [18]. Line ‘‘a” represents E=q corresponding to equivalence of Young’s modulus of all-PP composites in a strut application, line ‘‘b” represents E1=2 =q corresponding to equivalence of Young’s modulus of all-PP composites in a beam application and line ‘‘c” represents E1=3 =q corresponding to equivalence of Young’s modulus of all-PP composites in a panel application (see Fig. 1).

sity where all the materials from Table 1 are represented. Because the scales are logarithmic, contours of constant performance indices plot as straight guide lines. Any two materials on the same guide line have the same performance for that particular index while any material above that line has superior performance. Fig. 2 shows guide lines of equivalent performance to all-PP composites. It is obvious that all-PP plates can have potential weight saving over glass fibre reinforced PP for stiffness limited applications, in spite of a much higher modulus of the woven glass/PP laminates, according to the first performance index ðE=qÞ. A similar analysis can be conducted for strength limited design, as shown in Fig. 3. Both compressive and tensile strength of all-PP, isotropic PP and woven glass/PP are plotted whereas only the tensile strength of GMT is plotted. Only the first criterion corresponding to all-PP tensile strength is drawn, since buckling and flexural resistance of all-PP beams and panels are defined by compressive modulus. Although all-PP has a superior buckling performance (Fig. 2), its

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Strength (MPa)

1000

all-PP (T ension) all-PP (Compression) GMT (Tension) Woven Glass/PP (Tension) Woven Glass/PP (Compression) PP (Tension) PP (Compression) Aluminium (Tension)

100 a

b c 10

d 1

10

Density (g.cm-3) Fig. 3. Tensile and compressive strength vs. density of all-PP and glass reinforced polypropylene. The lines represent the main mechanical performance indices for strength limited design [18]. Line ‘‘a” represents r=q corresponding to equivalence of tensile strength of all-PP composites in a strut application, line ‘‘b” represents r=q corresponding to equivalence of compressive strength of all-PP composites in a strut application, line ‘‘c” represents r2=3 =q corresponding to equivalence of compressive strength of all-PP composites in a beam application and line ‘‘d” represents r1=2 =q corresponding to equivalence of compressive strength of all-PP composites in a panel application (see Fig. 1).

low compressive strength can be a limiting factor for its use in structural applications. An analogy can be made with other oriented polymer fibres such as UHMW-PE fibres which are mainly used in high-performance yarns and ballistic protective equipment because their compressive strength is barely 2% of their tensile strength [17,20]. This limited compressive performance must be considered when designing all-PP components to resist flexural loading. 3. Materials 3.1. Sandwich face panels 3.1.1. all-PP sandwich face panels It has been widely reported in literature that it is possible to create highly oriented PP fibres or tapes, with high tensile strength and stiffness, by molecular orientation achieved during solid state drawing [21–31], and thus it is conceivable to use such tapes as a reinforcement for a composite material. Much has been published on alternative processing routes to achieve single polymer composites such as combining high performance fibres with isotropic films [32–34], impregnating fibres with a particulate matrix in slurry or dry powder form [35–38], or simply by sintering fibres together [39–43]. These existing technologies have some inherent limitations which reduce their viability, such as small temperature processing windows or low volume fractions of reinforcement limiting the ultimate mechanical properties of the composites. Highly oriented, high modulus mono-extruded fibres or tapes can be effectively welded together by melting the surface of the tapes and applying pressure to achieve a good bonding and fill any voids between the tapes [44–46]. These mono-extruded tape or fibre have the advantage of only containing a single grade of material. However, these mono-extruded systems are highly sensitive to compaction temperature during consolidation, since the matrix phase is achieved by the sacrifice of a proportion of the reinforcement phase and so there is a risk of excessive molecular relaxation during consolidation of tapes or fibre bundles into composites.

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The research reported in this paper focuses on the creation of polypropylene sandwich panels, consisting of all-PP composite laminate faces and PP honeycomb or foam cores. The all-PP composite laminate faces are created by using co-extruded tape technology to allow a large temperature processing window for consolidation and high volume fraction of reinforcement. In previous publications [9–13,47–52] and a series of theses [14,53,54], these composites have been described in greater detail, as has a novel processing route which shall be summarised here. A polymer tape with a skin:core:skin structure is produced by co-extrusion of a ethylene–propylene copolymer skin surrounding a polypropylene homopolymer core. The relative proportions of this tape are 1:20:1 (skin:core:skin) and the tape is co-extruded at a rate of 6 m.min1. This tape is subsequently drawn in a continuous twostage drawing process through hot air ovens [30] and leaves the final oven at 102 m min1. Thus when drawn, the tape has experienced a draw ratio of 17 and has approximate dimensions of 2.15 mm wide and 0.65 lm thick. This drawing process results in a high degree of molecular orientation and the drawn tapes possess a high tensile strength (450 MPa) and stiffness (15 GPa). While the mechanical properties of the PP tapes are clearly much less than conventional composite reinforcements such as glass fibres (tensile strength = 3.5–5 GPa, tensile modulus = 70–90 GPa [55]), the high volume fraction of reinforcement achievable in all-PP composites (Vf > 90%) allows all-PP composites to have competitive mechanical properties with conventional PP matrix composites [13]. This highly drawn tape is then woven into a plain weave fabric with an areal density of 100 g m2, at a rate of approximately 600 m2 h1 using commercial polyolefin tape geotextile weaving apparatus. This plain weave fabric is consolidated into all-PP laminates by pressing in a double belt press (Institut für Verbundwerkstoffe GmbH, Germany) at a pressure and temperature of 2.5 MPa and 145 °C, at a linear output rate of 1 m min1. In this consolidation process, the copolymer skin layer of the tapes is melted, while the highly oriented homopolymer core structure retains the high degree of molecular alignment. The application of pressure forces adjacent fabric plies together, causing the copolymer skin layers to weld, and upon cooling, this results in a rigid laminate structure. Therefore, there is no matrix impregnation process as may be expected in some composite processing routes, instead the process described in this paper uses a process akin to welding to bond together polymeric tapes which are already coated with the matrix phase. The high molecular orientation achieved by solid state drawing of the constituent tapes is retained during consolidation and hence the laminate has high mechanical properties. The reinforcement to matrix ratio is extremely high compared to traditional glass fibre reinforced composite materials. The all-PP composite laminates reported in this paper have a volume fraction of reinforcement (highly oriented PP phase) of 89%, with the remaining 11% of the composite being composed of the propylene copolymer skin serving as a matrix phase. The effects of temperature and pressure on the mechanical properties of similar laminates, created in a non-continuous processing route, have been presented previously, together with images of the all-PP fabric prior to consolidation and consolidated laminates [13,50]. Two thicknesses of all-PP laminates were investigated, consisting of either 3 or 9 plies of fabric, which resulted in face panels of 0.43 and 1.19 mm thickness, respectively. These laminates are subsequently used as faces for production of sandwich panels. 3.1.2. Glass fibre reinforced PP sandwich face panels To compare the performance of sandwich panels containing all-PP face panels to sandwich panels containing conventional glass fibre reinforced PP face panels, two types of glass fibre

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reinforced PP were obtained. These were a plain woven glass fibre reinforced PP and a short random glass fibre reinforced PP. The plain woven glass fibre reinforced PP (woven glass/PP) was supplied as laminates of Twintex material (Saint Gobain-Vetrotex, France). This material has a 60% weight fraction of glass fibres (35% volume fraction), and was supplied in sheets of 0.5 mm thickness. The short random glass fibre reinforced was supplied as laminates of GMT (Quadrant Composites, Switzerland). This material has a 30% weight fraction of glass fibres (23% volume fraction), and was supplied in sheets of 1.12 mm thickness. The mechanical properties of the face panel materials used are presented in Table 1. 3.2. Sandwich core materials Two PP tubular honeycombs and one PP foam have been used as core materials in this study. The honeycombs were supplied by Tubus Bauer GmbH, Germany, and were type PP 8-80 and PP 8-120, having densities of 80 kg m3 and 120 kg m3, respectively. Both honeycombs have a thickness of 20 mm, a cell diameter of 8 mm and consist of hexagonally packed tubes. These honeycombs are

e

A

B

L or Ribbon direction

d

Y

manufactured by first producing co-extruded tubes, which are aligned and stacked in a parallel fashion before being bonded into a honeycomb structure in a subsequent step. The honeycomb is shown in Fig. 4. The foam core of type RPF was supplied by Sekisui, Japan [56], which has a density of 100 kg m3 and a thickness of 9 mm. It is a rigid PP-foam with pores of ellipsoidal shape oriented in the thickness direction. The foam core is shown in Fig. 5. 3.3. Adhesive materials Traditional thermoset adhesives cannot be used to bond the core to the face of polypropylene based sandwich panels because polyolefins have a very low surface tension [57]. Although oxidative treatment can increase the surface tension of polyolefins, welding is a preferred bonding technique, where the polyolefin surface is melted and brought in contact with the core [58]. An alternative to welding is to use hot-melt PP adhesive film to bond the sandwich faces to the core [59], and this process is applied in this paper. To assess the use of melt adhesive films to bond the all-PP face panels to the sandwich cores, two types of adhesive films were investigated, as shown in Table 2. These consisted of two types of propylene based copolymers, and are termed copolymer HS (high sealing temperature) and copolymer LS (low sealing temperature). These are random copolymers in which ethylene groups are inserted at intervals along the propylene chain. The ethylene units disrupt polymer regularity and so makes the chain more flexible, which in turn reduces crystallinity, modulus, melting point, and the sharpness of melting point. Thin films were produced in a film blowing extruder, with a thickness of 40 lm. Both copolymers have a ‘‘melting” temperature at approximately 135 °C but the copolymer LS has a lower sealing temperature due to a broader melting peak and a higher content of ethylene groups. The melting behaviour of these materials was characterised by differential scanning calorimetry.

W direction X

Table 2 Properties of melt adhesive copolymer films

Fig. 4. The structure of the tubular PP honeycomb core material used in this study. The right hand side shows a photograph of the cross-section of the honeycomb, while the left hand side shows a schematic of the tubular honeycomb at the same scale for clarity. The thickness ðtc Þ of the honeycomb was 20 mm. The cells have a diameter (d) of 8 mm and a wall thickness (e) of 0.2 mm (80 kg m3) or 0.3 mm (120 kg m3). The section A–B shown is the section with highest density across the L direction.

Tensile modulus (GPa) Tensile yield Stress (MPa) Melt flow index (230 °C, 2.16 kg) (g 10 min1) Initial sealing temperature (30 lm film) (°C)

Fig. 5. Photograph of the PP foam sandwich core material used in this study.

Copolymer HS

Copolymer LS

0.89 23 5.5 106

0.54 17 5.5 80

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Face Panel

Hot melt adhesive PP

Core Face Panel Fig. 6. Composition of a polypropylene sandwich structure. In the case of all-PP laminates being used for the sandwich face panels, together with PP based hot melt adhesives and a PP based core, the resultant sandwich panels clearly contain no ‘‘foreign” reinforcements.

4. Specimen production Sandwich panels were produced consisting of all-PP, GMT or woven glass/PP face panels, and a core composed of either a PP honeycomb or a PP foam. The stacking sequence used to create the sandwich panels was: face panel-adhesive film-core-adhesive film-face panel, as illustrated in Fig. 6. The stacked collection of layers is then subjected to heat and pressure to bond the sandwich faces to the core material, by melting the adhesive film. The method used to apply this heat and pressure depends on the dimensions required for the test specimen. Specimens produced for peel testing used a hot press to apply heat and pressure, since this is the most well controlled route to reproduce small sandwich specimens. Specimens produced for flexural testing used a vacuum bagging technique which will be described later. 4.1. Production of sandwich panels for peel testing Sandwich panels were made from all-PP face panels and honeycomb or PP foam cores to investigate the effect of a film adhesive layer on the interfacial strength between the all-PP face panels and the core. Sandwich panels were produced with an unsymmetrical lay-up consisting of a 3 mm GMT sheet (Quadrant Composites, Switzerland), a melt adhesive film, the core material, a melt adhesive film and a single ply of the all-PP fabric. This unsymmetrical lay-up was necessary to provide a combination of sandwich panel stiffness achieved by the thick GMT sheet, with a highly flexible all-PP face to apply peeling in order to determine the interfacial bond between the all-PP fabric face and the core. Specimens were produced at temperatures between 125 °C and 145 °C, using either one of the copolymer melt adhesive films or no melt adhesive film at all. These sandwich panels were produced as follows in a hot press fitted with mould stops to prevent crushing of the sandwich core. The samples were placed between the cold platens of a hot press. The press was closed to reach the mould stops, and the platen temperature was raised to the processing temperature. The sandwich panels were held at temperature for 15 min, as determined by thermocouples mounted on the panel surface, after which time, the platens were cooled by an internal water cooling mechanism. When the sandwich panels had decreased to 40 °C, the platens of the press were opened and the sandwich panels removed. 4.2. Sandwich panels for flexural testing To assess sandwich panels by four-point bending, the specimens require a high length to thickness ratio so that shear failure of the core or the adhesive is avoided. As the honeycombs are 20 mm thick, panels of 500 mm long are needed. Consolidation in a hot press used for smaller composite laminates presented elsewhere [13] could not guarantee a uniform temperature distribu-

Fig. 7. Cross-section of a vacuum bagging lay-up used to produce large honeycomb sandwich panels. Showing 1: sandwich lay-up, 2: thin metal plate, 3: hollow metallic frame to avoid lateral pressure, 4: nylon vacuum bag, 5: bleeder cloth, 6: sealing tape, 7: connection to vacuum pump, 8: connection to vacuum pressure gauge, 9: 1 mm metallic plate to provide planar pressure distribution and avoid surface defects (only used with 0.5 mm thick face panels).

tion over a specimen of this length. For this reason, a vacuum bagging technique was used. The vacuum bagging technique is shown schematically in Fig. 7, and applies a uniform pressure over the sandwich panel surface, by partially evacuating air from the bag containing the sandwich panel. The entire system is placed in an oven to provide the elevated temperature required to bond the sandwich face panels to the core. The maximum pressure applicable by this method is atmospheric pressure, 0.1 MPa. To prepare sandwich panels for flexural testing, pressure and temperature of 0.04 MPa and 135 °C, respectively, were used to bond the face panels to the core. Any greater pressure was seen to cause collapse of the cores at these temperatures. To prevent the application of pressure to the edges of the honeycomb cores, a metallic frame was placed around the sandwich panels. This is necessary since the in-plane compressive strength of the honeycomb cores is negligible, so these structures may readily collapse by applying any compressive loading in the plane of the honeycomb. The samples were placed under vacuum, and placed in the oven. The panels were heated to 135 °C as determined by thermocouples mounted on the surface of the vacuum bag. This temperature was maintained for 15 min to achieve stable temperature, and the system was cooled. To prevent any relaxation of the oriented all-PP structure at elevated temperatures [13], the sandwich panels were cooled to ambient temperature before pressure was removed. Although the GMT and woven glass/PP faced composites did not experience this risk, the same procedure was used for processing these sandwich panels as was used for creating all-PP faced sandwich panels.

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The specimens produced for flexural testing are shown in Table 3. A limited supply of the core materials meant that not all combinations of sandwich faces and cores could be investigated. Fig. 8 shows representative photographs of sections of all-PP and woven/glass faced honeycomb sandwich panels. 5. Experimental method 5.1. Differential scanning calorimetry To determine the melting behaviours of both of the copolymer adhesive films investigated in this paper, differential scanning calorimetry (DSC) was performed on 5 mg samples of polymer taken from the pellet form, using a TA Instruments DSC Q1000 differential scanning calorimeter. To remove the effect of thermal history of the DSC results, samples were heated in the DSC from ambient temperature to 180 °C at 10 °C min1 and then cooled to 20 °C, also at 10 °C min1. Immediately, the samples were reheated to 180 °C at 10 °C min1 and the endothermic data taken from this second heating stage. The data obtained from these experiments show each endothermic event as a vertical peak.

5.2. Peel testing of sandwich face panel A key property of sandwich panels is the strength of the coreto-face bonding. Ideally, the copolymer films used as hot-melt adhesives should be heated above their melting temperature in order to obtain an optimum bond. However, sealing is possible below the melting temperature as these copolymers have a broad melting peak. The strength of core-to-face bonding of sandwich panels depends on many variables such as the nature of the hot-melt adhesive, the core and face materials and the temperature of bonding. Panels using both of the copolymer melt adhesives and the different core materials were manufactured. Sandwich panels containing the honeycomb core were peeled along the L direction of the honeycombs (see Fig. 4). All of the sandwich panels were produced in a hot press as described earlier. In order to create a starting tab to initiate the peel loading, a 70 mm long region of the allPP fabric face was separated from the core by using a PTFE release film (12 lm thick) placed between the all-PP fabric and the copolymer melt adhesive film. This acted as a physical barrier to prevent contact between the melt adhesive and the all-PP fabric and thus gave an area of no adhesion. This unbonded area was used as a

Table 3 Construction of sandwich panels described in this paper Sandwich panel code

Core material 3

Face material

Core thickness (mm)

Face thickness (mm)

Areal density (kg m2)

Woven glass/PP GMT all-PP (9 layers) all-PP (3 layers)

20 20 20 20

0.50 1.12 1.19 0.42

3.15 4.15 3.46 2.26

80H1 80H2 80H3 80H4

Honeycomb (80 kg m

120H1 120H2 120H3

Honeycomb (120 kg m3)

Woven glass/PP GMT all-PP (9 layers)

20 20 20

0.50 1.12 1.19

4.01 5.01 4.32

F1

Foam core

all-PP (3 layers)

9

0.42

1.56

)

Upper all-PP face panel Honeycomb core Lower all-PP face panel 1cm

Upper woven glass/PP face panel Honeycomb core Lower woven glass/PP face panel 1cm Fig. 8. Photograph of a cross-section of an all-PP faced honeycomb sandwich panel (above) and a woven glass/PP faced honeycomb sandwich panel (below). The woven nature of the all-PP laminate is visible on the upper surface of the upper sandwich panel, as is the texture of the woven glass on the upper surface of the lower sandwich panel.

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available in very thin sheets (a single ply of all-PP fabric has a thickness of approximately 150 lm). The unbonded tab is gripped in the lower jaws of the tensile testing machine, while the sandwich core is freely supported horizontally on two rollers in the test fixture. The action of peeling the face downwards 90° to the sandwich panel plane, causes the sandwich panel to roll horizontally. An Instron tensile testing machine equipped with a 1 kN load cell was used to peel the all-PP away from the core at a cross-head speed of 100 mm min1 while recording the peel force. The peel specimens had dimensions of 250 mm length and 45 mm width. As described earlier, the lay-up was non-symmetrical (Fig. 9) using a 3 mm thick GMT sheet as the upper face panel to achieve the necessary panel stiffness, and the all-PP thin sheet to be peeled off as the lower face. 5.3. Flexural testing of sandwich panels

Fig. 9. Peeling of a single layer all-PP ply from a honeycomb core sandwich panel. Specimen slides to the right as bottom face is peeled off. (Note the 3 mm thick GMT upper face).

tab to load the all-PP fabric in order to start the peeling of the allPP face from the core. Fig. 9 is a picture of the peel jig used to determine the interfacial strength of the all-PP face panels to the sandwich cores, which is also shown schematically in Fig. 10. The standard test method to measure debonding strength of sandwich panel faces, the climbing drum peel test (ASTM D1781), was not necessary for all-PP faces as unlike conventional sandwich panel faces (e.g. GMT) they are

The mechanical properties of the sandwich panels were determined by four-point bending test (ASTM C393) using the same tensile machine as for peel testing, at a cross-head speed of 1 mm min1. The distance between the load points was equalled to half the span length. The specimens were 50 mm wide. The length of the sandwich panels was parallel to the L direction of the honeycombs (see Fig. 4). To prevent local buckling of the inner face during flexure, load pads were used for all specimens to evenly distribute the load. A displacement transducer was used to measure the mid-span deflection of the honeycomb panels while the deflection at the load point was used for the foam cored sandwich panels. The maximum load achieved before failure is identified, defined as the flexural strength, and the failure mode recorded.

Thick GMT Upper Face Sandwich Core all-PP Peel Face Free Rotating Rollers

Sandwich panel translates horizontally

all-PP peel face is pulled downwards Fig. 10. Schematic of test set-up to measure peel strength of all-PP faces to sandwich core. As the thin all-PP face is pulled downwards, the face peels away from the core, and the sandwich panel translates horizontally.

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The deflection ðdÞ of a sandwich panel in a four-point bending test is the sum of a flexural deflection ðdb Þ and a shear deflection ðdc Þ

The value of k is 96 for deflection at the load point and 768/11 for the mid-span deflection [60]. Gc is the core shear modulus, b, the width of the panel and s, the distance between the centres of the faces. Eq. (1) can be rewritten as follows: 2

d=PL ¼ ðL =kEIÞ þ ð1=8SÞ

ð4Þ

Parameters e and d are defined as shown in Fig. 4. A value of 460 MPa was taken as shear modulus, G, of polypropylene. The specific flexural stiffness is evaluated by dividing the flexural stiffness by the areal weight of the sandwich panel. The weight of the adhesive film was neglected, so that the sandwich weight per unit area equals:

W ¼ 2qf tf þ qc t c

ð5Þ

Because the sandwich face is very thin compared to the core ðtc ot f Þ and the core in-plane Young’s moduli are neglected, the flexural stiffness of the panel can be derived by a simplified model [1]:

D¼

Ef btf ðt c þ t f Þ2 2

O

106 C

ð3Þ

If the deflection is measured for different load spans, the shear stiffness, S, is derived by plotting d=PL as a function of L2 . Sandwich beams were tested for three different load spans to determine the shear modulus of each core: 150, 220 and 420 mm for the honeycomb core panels and 80, 140 and 220 mm for the foam core panels. The flexural stiffness (EI) of the panels can then be evaluated using Eq. (1) from the deflection of the largest span length, i.e. 420 mm for the honeycomb core and 220 mm for the foam core panels. Before the flexural stiffness of the sandwich panels can be derived, the core shear modulus is determined as discussed before (see Eq. (3)) for the different core materials. Meraghni et al [61] derived an analytical solution based on the laminate plate theory for the determination of the shear modulus in planes perpendicular to the faces of a tubular honeycomb:

2eG Gxz ¼ Gxy ¼ pffiffiffid  3 2 þe

O

80 C

ð1Þ ð2Þ Heat Flow (Endotherm Up)

d ¼ ðPL3 =kEIÞ þ ðPL=8SÞ S ¼ Gc bs

HS Copolymer LS Copolymer

ð6Þ

where Ef is the Young’s modulus of the sandwich face, and b, is the width of the panel. The theoretical specific flexural stiffness is obtained by dividing Eq. (6) by Eq. (5). It should be noted that this model is simplified, and provides a first estimation of the flexural stiffness of the panel. The model does not incorporate effects due to shear deflection, which may account for the discrepancy between the predicted values and the values derived experimentally. Also, the model does not account for the fact that the all-PP sandwich faces do not possess equal moduli in tension and compression. The modulus of all-PP composites used in these models is the tensile modulus, which is due to molecular orientation of the precursor tapes, and it is significantly greater than the compressive modulus, which does not increase due to molecular orientation. 6. Results and discussion 6.1. Differential scanning calorimetry Fig. 11 shows the DSC data obtained for copolymer LS and HS melt adhesives. Also indicated on Fig. 11 are the manufacturers recommended sealing temperatures. It is clear that copolymer LS

40

60

80

100

120

140

160

O

Temperature ( C) Fig. 11. DSC curves of PP copolymer melt adhesive films. The minimum sealing temperatures suggested by the manufacturer are shown by the arrows.

melt adhesive has a lower temperature endothermic event associated with the manufacturers melt sealing temperature, which is absent from copolymer HS melt adhesive. The lower melting temperature of copolymer LS melt adhesive allows adhesion of the sandwich faces to the sandwich cores at lower temperatures, aiding processing but also limiting the upper operating temperature of the resulting sandwich panel. This lower temperature endotherm seems to be due to an increased fraction of ethylene monomer units. 6.2. Peel testing of sandwich face panel The peel forces which occurred due to the peeling a single ply of all-PP fabric away from the core material were measured. The peel forces measured were normalised per unit width of specimen. The recorded peel forces were seen to oscillate with a regular pattern. Two representative peel force graphs are shown in Fig. 12, as a function of peel displacement. Fig. 12a shows the peel force per unit width of an all-PP fabric ply bonded with a film of copolymer LS melt adhesive to a 80 kg.m3 honeycomb. The distance between two successive peaks is on average 7.1 ± 0.7 mm. Observation revealed that these peaks correspond to the sections A–B with highest density of the honeycombs (Fig. 4), i.e. where the greatest area of bonded surface was present, as the peel propagates. Fig. 12b shows the peel force profile for the foam core, which is also periodic but the distance between the peaks here is considerably smaller, with an average distance between successive peaks of 2.2 ± 0.4 mm. This distance equals the width of the tape present in the all-PP fabric, so here the structure of the all-PP fabric is likely to be the cause of the profile. A similar effect has been reported when two layers of bonded all-PP fabric are pulled apart in a peel test; the plot of peel force as a function of displacement also has a period equal to the width of the tape [14]. The large unbonded areas seen in the peeling of all-PP fabrics from the honeycomb cores (Fig. 12a), corresponding to the centre of the honeycomb cells, present a hazard for the sandwich structures as they can be regarded as unbonded areas. In this respect, the more homogeneous foam core appears attractive, providing a more uniformly distributed bonding. For the remainder of this paper, only the average value of the peel force profile will be presented. The effect of different melt adhesive films and different sandwich production temperatures on the average peel force per unit width is presented in Fig. 13, and will now be discussed. Each data

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Average Peel Force per Width (N/cm)

12

10

8

No Melt Adhesive Film 3 Honeycomb 120kg/m Copolymer Melt Adhesive Film HS 3 Honeycomb 80kg/m 3 Honeycomb 120kg/m

6

4

2

0 125

130

135

140

145 O

Sandwich Production Temperature ( C)

Average Peel Force per Width (N/cm)

12

10

Copolymer Melt Adhesive Film LS 3 Honeycomb 80kg/m 3 Honeycomb 120kg/m 3 Foam Core 100kg/m

8

6

4

2

0 Fig. 12. Peel force as a function of displacement. Peel force is normalised per unit width of specimen to accurately compare specimens of slightly varying width. (a) all-PP face bonded at 135 °C using copolymer LS to the 80 kg m3 honeycomb. Note the peaks correspond to the dense A–B sections of the honeycomb (see Fig. 4). (b) all-PP face bonded at 130 °C to the foam core.

point in this graph represents the average of five test specimens, while the error bars represent the standard deviation. Fig. 13a shows the average peel force per unit width of sandwich panels bonded either with copolymer HS or no melt adhesive film at a range of processing temperatures. Sandwich panels produced at temperatures below 125 °C showed negligible adhesion, despite this temperature being above the ‘‘sealing temperature” recommended by the melt adhesive manufacturer. Generally, all of the sandwich panels produced showed that peel force increased with increasing production temperature. The higher density honeycomb yielded a greater peel force, due to the greater wall thickness of the tubular sections, and hence larger bonding area. Also noticeable is that the sandwich panel created with the high density honeycomb but no melt adhesive film, showed relatively low peel forces (1.96 N cm1), emphasising the need for an adhesive melt film. Fig. 13b shows the average peel force per unit width of sandwich panels bonded with copolymer LS, with honeycomb or foam cores at a range of processing temperatures. As seen before, increasing production temperatures resulted in greater peel forces [10]. The sandwich panels with the foam core gave higher peel forces, most likely due to a greater contact area and more homogenous bonding. These results show that the use of a melt adhesive with a lower sealing temperature allows greater bonding at lower temperatures. For this reason, all sandwich panels created for flex-

125

130

135

140

145 O

Sandwich Production Temperature ( C) Fig. 13. Average peel force as a function of temperature. Peel force is normalised per unit width of specimen to accurately compare specimens of slightly varying width. (a) Core-to-face adhesion either using films of copolymer HS or direct bonding of the all-PP face to the honeycomb. No adhesion could be measured at 125 °C for the 80 kg m3 honeycomb and below 145 °C when no copolymer film was used. (b) Core-to-face bonding using films of copolymer LS. Error bars represent standard deviation of five specimens.

ural testing use the copolymer LS as a melt adhesive film between face panels and the core, and were processed at 135 °C. 6.3. Core shear modulus From Eq. (1), the foam core has a shear modulus of 7.1 MPa, the low density tubular honeycomb has a shear modulus of 14 Mpa, while the high density tubular honeycomb has a shear modulus of 24 MPa. It should be noted that only the shear modulus in the x–z plane of the honeycombs (Fig. 4) is determined, as the length of the flexural specimens was parallel to the ribbon direction. In spite of a higher density, the shear modulus of the foam core is only half that of the 80 kg m3 honeycomb. It is an expected result as two-dimensional cellular solids generally have better out-ofplane properties than homogeneous cores [62]. This is identified as the main weakness of foam cores. The values derived from Eq. (4) are 25 MPa for the 80 kg m3 honeycomb and 38 MPa for 120 kg m3 honeycomb, which are higher than the experimental values, respectively 14 MPa and 24 MPa. A similar discrepancy between Eq. (4) and experimental

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Table 4 Flexural stiffness of sandwich panels described in this paper Sandwich panel code

Core material 3

Face material

Areal density (kg m2)

Flexural stiffnessa [N m2 cm1]

Specific flexural stiffness (N m3 kg1)

Woven glass/PP GMT all-PP (9 layers) all-PP (3 layers)

3.15 4.15 3.46 2.26

12.70 ± 0.39 8.95 ± 0.20 12.65 ± 0.53 3.87 ± 0.09

403 ± 12 215 ± 5 366 ± 15 172 ± 4 282 ± 2 179 ± 4 282 ± 11

80H1 80H2 80H3 80H4

Honeycomb (80 kg m

120H1 120H2 120H3

Honeycomb (120 kg m3)

Woven glass/PP GMT all-PP (9 layers)

4.01 5.01 4.32

11.31 ± 0.07 8.98 ± 0.20 12.18 ± 0.49

F1

Foam core

all-PP (3 layers)

1.56

0.79 ± 0.01

a

)

51 ± 1

Per unit width of specimen.

values of shear moduli was reported by Meraghni et al [61]. It is these experimentally obtained values which are used for further calculation in Eq. (2). 6.4. Flexural stiffness of sandwich panels The flexural stiffness of the panels is presented in Table 4. The foam panel specific rigidity is very low because the core was thickness supplied was low. The flexural stiffnesses of the panels 80H1, 80H2 and 80H3 can be considered identical to the value of the panels 120H1, 120H2 and 120H3. The two sets of panels are only differentiated by the density of the honeycomb core used (see Table 3). This result means that it is correct to consider the tubular honeycomb cores as anti-plane cores, i.e. their in-plane moduli of elasticity are negligible. The evaluation of the different face materials is not as straight forward because the geometrical parameters, such as sandwich core and face thickness vary as well as the sandwich face material. However, the panels H1, containing woven glass/PP faces, have the same specific rigidity as the panels H3 which contain considerably thicker all-PP faces. The lower density of all-PP compensates for the higher modulus of the woven glass/PP material, if the specific stiffnesses are considered. In Fig. 14, the experimental specific flexural stiffness values are plotted against the theoretical values. It can be observed that the theoretical data are approximately 20% lower than the experimental data, independent of the panel composition. 6.5. Flexural strength of sandwich panels Three modes of failure were observed: shear failure which appeared to initiate at the interface, outward buckling or compressive

Fig. 14. Comparison of the experimental flexural stiffness with the theoretical values. The values are obtained from Eq. (6), and divided by the areal weight of the panels. Parameters of the different panels can be found in Table 3. Note that experimental values are 20% lower than calculated values.

Fig. 15. The different failure modes observed during flexure of all-PP sandwich structures.

failure of the face (Fig. 15). Both the core shear stress and the face stress at failure were calculated according to ASTM C393. The flexural strength of the panels based on 80 or 120 kg m3 honeycomb core indicate that the heavier honeycomb does not improve the flexural strength of the panel. This is to be expected since none of the panels failed in the core. The H1 panels (woven glass/PP faces) are the weakest, while the H2 panels (GMT faces) have the highest flexural strength. However, the strength of the H3 panels (all-PP faces) is similar to the strength of the H1 and H2 panels, which means that if the areal weight of the panels is considered, the all-PP faced sandwich panels will have the highest specific

Fig. 16. Four-point flexure of an all-PP faced honeycomb sandwich panel. Note the outward buckling failure of the upper face.

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N.O. Cabrera et al. / Composites: Part B 39 (2008) 1183–1195 Table 5 Stresses experienced by sandwich panels described in this paper Sandwich panel code

Core material 3

Face material

Strengtha (N mm1)

Failure mode

Core shear stress (MPa)

Face stress at failure (MPa)

Woven glass/PP GMT all-PP (9 layers) all-PP (3 layers)

584 ± 33 1140 ± 51 930 ± 10 268 ± 6

Buckling Shear Buckling Buckling

0.14 0.27 0.22 0.07

30.0 25.5 19.4 17.2

Buckling Shear Buckling

0.13 0.27 0.27

28.1 25.1 24.0

Compression

0.16

21.6

80H1 80H2 80H3 80H4

Honeycomb (80 kg m

120H1 120H2 120H3

Honeycomb (120 kg m3)

Woven glass/PP GMT all-PP (9 layers)

551 ± 17 1150 ± 40 1150 ± 74

F1

Foam core

all-PP (3 layers)

288 ± 8

a

)

Sandwich flexural strength is defined as the failure load normalised by panel width.

strength (flexural strength/areal weight). The H1 and H3 panels both failed by buckling of the face panels, while H2 panels failed in shear mode at the adhesive bond. A typical specimen which has failed by buckling is shown in Fig. 16. The face stress at failure is similar independently of the sandwich composition or failure mode (see Table 5). The reason for this is not clear but the low tensile yield stress of the copolymer film, LS (Table 2), may account for the outward buckling failure of the compressive face. Representative load vs. deflection curves are shown in Fig. 17 for a selection of panels. Both panels based on all-PP faces have ductile behaviour due to yielding of the face in compression. So even if the core-to-face bond of the H3 panels was improved, it is believed that they would fail at the same load by compressive failure of the face in a similar way to the foam core sandwich panels (F1). Provided the core-to-face bond does not fail, the strength of these panels would be limited by the compressive strength of the faces, so that panels based on glass reinforced PP faces would be stronger. Because the sandwich panels were assembled in an isothermal process, the temperature was limited to 135 °C due to the core and the all-PP temperature dependent behaviour. In a non-isothermal process, the adhesive could be heated to its optimum temperature (145 °C) while the all-PP face and the core would not reach the critical temperature (135 °C). This was beyond the scope of this research although such methods are well documented (e.g. the two step heating technique suggested by Rozant et al [63]). Another method to improve the honeycomb-to-face bond may be to use foamed hot melt adhesives [64–66] in order to increase the surface contact between the core and the face. However, the blowing temperature of the PP foam (215 °C) may prove to be too high. Generally, the low yield stress of the PP melt adhesive film appears to be the limiting factor of the bond strength of these sandwich panels, and the core properties and hence the strength of the sandwich structure will not always depend on the compressive strength of the face panels. It should be considered that if the design parameters of a sandwich panel are driven by the performance of the core, a different mono-component material solution could also be envisaged. Jullien et al [67] reported that for a door panel of a high-speed train, the PP tubular honeycomb was not selected for a sandwich core. Instead, an aluminium honeycomb was preferred because of its higher shear properties and low density (83 kg m3). Aluminium (see Table 1) is a serious alternative to 100% PP structures, since it is also possible to have high-performance structures which are 100% aluminium and also fully recyclable.

Wc ¼ 2  Wf

ð7Þ

This relationship is derived from Eqs. (5) and (6). Theulen and Peijs verified this criterion experimentally for sandwich constructions consisting of a foamed PVC core and glass fibre reinforced epoxy faces [68]. A similar criterion for the strength performance of a sandwich panel is based on the assumption that the sandwich strength is limited by only one mode of failure. It has been shown earlier in this paper that shear failure of the core and compressive or buckling failure of the face may limit the strength of PP based sandwich panels. However, the minimum weight criterion with respect to stiffness is the ultimate tool to compare the performance of the different polypropylene composites faces used in this study. If a panel with a given core material satisfies Eq. (7), both the thickness of the face and the weight of the panel are linear functions of the core thickness. In addition, given the core thickness and density, the flexural stiffness of the panel, as derived from Eq. (6), varies only due to the face properties (i.e. density and stiffness). Considering a core of 80 kg m3 and a sandwich panel weight of up to 5.5 kg m2, the maximum flexural stiffness of 100% PP sandwich structures is plotted as a function of the panel weight in Fig. 18. In agreement with the results from Section 5.3, the Young’s moduli values from Table 1 were reduced by 20%. As expected, the performance of all-PP as a face material lies between that of GMT and woven glass/PP but closer to the latter. The all-PP face can be twice as thick because its density is only half that of the woven glass/PP material. However the resulting sandwich panel is only 2.6% thicker because the total thickness of the panel depends mainly on the core thickness. It is important to note that the three curves of Fig. 18 are unrealistic, as these thermoplastic composite

6.6. Critical weight criterion Kuenzi [15] described how the weight of a sandwich panel can be minimised for a given flexural stiffness or strength independently of the material solution. Provided certain conditions are satisfied, the core weight should be twice that of the faces for a minimum weight sandwich of specified stiffness:

Fig. 17. Typical load-deflection curves for different sandwich panels. Panels with all-PP faces (solid lines) show ductile failure.

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Fig. 18. Flexural stiffness of polypropylene sandwich panels with a 80 kg m3 core according to the minimum weight criterion ðW c ¼ 2  W f Þ. Flexural stiffness was calculated using Eq. (6). The symbols represent discrete values of the faces as these materials are only available as laminated sheets of discrete thickness.

laminates are only available in discrete thicknesses. This is especially true for glass fibre reinforced PP, such as TwintexÒ. The areal weight of the woven materials is a multiple of the nominal weight of the fabric plies. Because all-PP fabric (105 g m2) is seven times lighter than a light weight woven glass/PP laminate (745 g m2), sandwich panels can be designed with all-PP faces closer to their minimum weight criterion. Each data point for all-PP faces in Fig. 18 represents the optimum flexural stiffness for a given allPP face thickness. Although this paper concerns the production and comparison of planar sandwich panels, no reference is made to the manufacture of more complex geometries of sandwich panels. Although the forming of all-PP composite laminates into various three-dimensional geometries has been investigated [54], the processing of three-dimensional sandwich structures is outside the scope of this paper. The ability to design optimum sandwich panels may reduce the cost and mass of the sandwich panels, although an accurate assessment of the total cost of a sandwich panel product would also need to incorporate costs for sandwich face production and subsequent forming steps, which are not considered here. The question of recyclability of these materials needs full assessment. A study simulating the effect of recycling all-PP composites into new all-PP tapes has been performed and is reported elsewhere [14]. In summary, an extruder feedstock was prepared which was composed of the proportional content (approximately 11%) of matrix copolymer phase blended with PP homopolymer, to represent feedstock which would result from granulating allPP composite. This material was used as the core (reinforcement) phase, and co-extruded with a virgin copolymer skin (matrix) layer and ‘‘second generation” tapes were drawn as in the normal tape production route described earlier in this paper. These drawn second generation tapes showed a small decrease in mechanical performance compared to the first generation tapes (17% decrease in tensile modulus), but this decrease may be addressed by combining recyclate material with virgin material during recycling, to reduce this effect. This indicates the potential recyclability of all-PP composites. 7. Conclusions The use of material selection tools to assess the performance of all-PP composites in structural sandwich applications reveals that

there is a potential to reduce the weight of applications where GFRPP is currently used. This is particularly true in the case of stiffness limited design. It was, however, stressed that the low compressive strength of all-PP can be a limiting factor for the use of all-PP structures loaded in flexure or compression. It has been shown that 100% PP sandwich panels can be manufactured without losing the mechanical properties of all-PP sheets. PP tubular honeycombs and rigid PP foam were used to manufacture sandwich structures. The bonding strength between the face and the core was optimised through the evaluation of a peel test where parameters such as the temperature, the nature of the core material and the copolymer film were varied. It was shown that the foam-to-face bond was more homogeneous due to a higher surface contact area. Sandwich panels with GMT, woven glass/PP or all-PP faces were tested under four-point bending. Unlike panels with GFRPP faces, all-PP faced panels exhibit a ductile behaviour due to yielding of the compressive facing. The stiffness of a sandwich panel depends on the material choice as well as geometrical parameters such as the core and faces thickness. In order to compare the performance of all-PP with GFRPP faces, the minimum weight criterion for a given flexural stiffness was used. It was shown that the low density of all-PP allows the creation of 100% PP sandwich structures with a similar specific flexural stiffness to panels with commingled woven glass/PP facings. The possibility to create all-PP face panels in incremental layers of thin plies gives the designer a greater freedom to create optimum low weight sandwich panels. Acknowledgement This work was sponsored by the Dutch Government’s Economy, Ecology and Technology (EET) programme for sustainable development, under Grant No. EETK97104. References [1] Allen HG. Analysis and design of structural sandwich panels. Pergamon Press; 1969. [2] Tiganis BE, Shanks RA, Long Y. Effects of processing on the microstructure melting behaviour and melting temperature of polypropylene. J Appl Polym Sci 1996;59:663–71. [3] Gonzalez-Gonzalez VA, Neira-Velazquez G, Angulo-Sanchez JL. Polypropylene chain scission and molecular weight changes in multiple extrusion. Polym Degrad Stabil 1998;60:33–42. [4] Brydson J. Plastics materials. Butterworth; 1999. [5] Bledzki AK, Gassan J. Natural fibre reinforced plastics. In: Cheremisinof NP, editor. Handbook of engineering polymeric materials. New York: Marcel Dekker; 1997. [6] Garkhail SK, Heijinrath RWH, Peijs T. Mechanical properties of natural-fibremat-reinforced thermoplastics based on flax fibres and polypropylene. Appl Compos Mater 2000;7:351–72. [7] Heijenrath R, Peijs T. Natural-fibre-mat-reinforced thermoplastic composite based on flax fibres and polypropylene. Adv Compos Lett 1996;5(3):81–5. [8] Leterrier Y. Life cycle engineering of composites in comprehensive composite materials. Polymer matrix composites, vol. 2. Amsterdam: Elsevier; 2000. [9] Alcock B, Cabrera NO, Barkoula NM, Loos J, Peijs T. The mechanical properties of unidirectional all-polypropylene composites. Compos Part A: Appl Sci Manuf 2006;37(5):716–26. [10] Alcock B, Cabrera NO, Barkoula NM, Loos J, Peijs T. Interfacial properties of highly oriented coextruded polypropylene tapes for the creation of recyclable all-polypropylene composites. J Appl Polym Sci 2007;104(1):118–29. [11] Alcock B, Cabrera NO, Barkoula NM, Peijs T. Low velocity impact performance of recyclable all-polypropylene composites. Compos Sci Technol 2006;66(11– 12):1724–37. [12] Alcock B, Cabrera NO, Barkoula NM, Reynolds CT, Govaert LE, Peijs T. The effect of temperature and strain rate on the mechanical properties of highly oriented polypropylene tapes and all-polypropylene composites. Compos Sci Technol 2007;67(10):2061–70. [13] Alcock B, Cabrera NO, Barkoula NM, Spoelstra AB, Loos J, Peijs T. The mechanical properties of woven tape all-polypropylene composites. Compos Part A: Appl Sci Manuf 2007;38(1):147–61. [14] Alcock B. Single polymer composites based on polypropylene: processing and properties. UK: Queen Mary, University of London. PhD Thesis; 2004. [15] Kuenzi EW. Minimum weight structural sandwich. US Forest Service Research Note, FPL-086; 1965. p. 1–20.

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