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Influence of pinning on static strength of co-cured metal-GFRP hybrid single lap joints
Composites: Part A 84 (2016) 196–208
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Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Influence of pinning on static strength of co-cured metal-GFRP hybrid single lap joints M.S. Islam a,b, Liyong Tong a,⇑ a b
School of Aerospace Mechanical and Mechatronic Engineering, The University of Sydney, Sydney, NSW 2006, Australia Cooperative Research Centre for Advanced Composite Structures (CRC-ACS), 1/320 Lorimer Street, Port Melbourne, Victoria 3207, Australia
a r t i c l e
i n f o
Article history: Received 2 September 2015 Received in revised form 11 January 2016 Accepted 17 January 2016 Available online 23 January 2016 Keywords: A. Hybrid A. Polymer-matrix composites (PMCs) D. Mechanical testing E. Joints/joining
a b s t r a c t This study presents an experimental investigation into the effects of through-thickness pinning reinforcement on the static strength and damage tolerance of hybrid mild steel–glass fibre prepreg co-cured composite single lap joints (SLJ). Stainless steel pins of 0.3 mm in diameter were inserted as mechanical fastening, in addition to adhesive bonding, to form hybrid joints between metal and glass fibre reinforced polymer substrates. Using the hybrid SLJ tensile testing, the failure modes and static strength were experimentally determined for mild steel–glass fibre prepreg co-cured composites. It is revealed that pinning can improve the static failure load via bridging mechanism by as much as 58% depending on the number and location of pins and the presence of clamping due to bent-ends. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Hybrid joints use both mechanical fastening and adhesive bonding simultaneously to join similar or dissimilar structural components in order to increase joining strength and fatigue performance and to decrease weight and stress concentrations [1,2]. Single lap joints have been used to assess the benefits of combining mechanical fastening and adhesive bonding in terms of static tensile strength. The mechanical fastening entity used in composite–composite or metal–composite joints has different diameter scales ranging from bolts in several millimetres [3–7] to pins or stitches in sub millimetres [8–13] and even in nanometres [14,15]. In the previous studies using bolts of several millimetre diameters, single and multiple bolts were used in bonded-bolted single lap joints subjected to tensile load [3–6]. Single steel bolt of 6.35 mm diameter was used with adhesive bonding in the static and fatigue tests of adhesive/bolted composite joints by Fu and Mallick [3]. Kweon et al. [5] conducted experiments on tensile testing of hybrid bonded-bolted composite-to-aluminium double lap joints with single bolt (of 4.763 mm bolt hole diameter) in the centre of the overlap and reported an improved hybrid joint strength when mechanical fastening was stronger than adhesive bonding. Recently, Matsuzaki et al. [6] reported that hybrid GFRPaluminium single lap joints using bolted/co-cured hybrid method ⇑ Corresponding author. Tel.: +61 2 93516949. E-mail address: [email protected] (L. Tong). http://dx.doi.org/10.1016/j.compositesa.2016.01.011 1359-835X/Ó 2016 Elsevier Ltd. All rights reserved.
with patterned M2 12 mm steel bolts had 1.84 times higher maximum shear strength and a quarter of the standard deviation compared with the adhesive failure strength of the co-cured joints for GFRP/aluminium single lap joints. It is evident that bolt diameter used in the tests in [3–6] has been reduced from 6.35 mm to 2 mm. In addition to bolt diameter, other parameters, such as washer size, clamping force [3] and bolt-to-washer clearance [7], also play a role in affecting the joint strength of hybrid single lap joints. Technologies, such as stitching and pinning, have been an alternative method to introduce mechanical fastening using stitches and pins as bolts with small diameter on the order of millimetre, sub millimetre [2] or even nanometre [14,15]. They have been used in three dimensional reinforced composites and structures and also in structural joints of similar and dissimilar materials. An improvement in static strength of stitched and pinned composite-tocomposite single lap joints has been observed [8–10]. The improvement in joint strength is remarkable given the volume content of stitches or pins on the order of 1–5%, and it could be attributed to many factors including the crack bridging process or the clamping effect due to stitch pretension. Recently, various additive manufacturing techniques have been developed to produce mechanically pinned and adhesively bonded hybrid metal–composite single lap joints [11–13]. Ucsnik et al. [11] employed a cold metal transfer (CMT) process to arc weld small pins of 0.8 mm in diameter with and without ball-heads on metal surface and prepared double lap joints. Their test results show that the static tensile strength can be improved by 11% and 52% for the
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cases of pins without and with ball-heads. Graham et al. [12] used laser metal deposition (LMD) to add pins of 1 mm diameter in the form of an array of macro-scale projections for hybrid joining between metal and CFRP prepreg laminates. They have found an improvement of 60–157% in measured ultimate strength with hybrid joints over plain co-cured joints. Parkes et al. [13] employed the direct metal laser sintering process to build six by six pin (of 1 mm diameter) arrays from a bed of titanium powder and prepared hybrid titanium-CFRP single lap joints. Their results show that the static ultimate tensile strength can be improved by 6.5 times compared to adhesively bonded joints. Additive manufacturing processes often require special equipments and tools, usually involving intellectual property, and thus accessing to such technologies may be limited. To our best knowledge, there is no report available in open literature studying behaviour of hybrid joints with metallic pins with sub millimetre diameter inserted by using conventional mechanical fastening methods, such as bolting and riveting. Hence the aim of this study is to investigate the influence of metallic pins on static strength of hybrid joints by using traditional mechanical fastening methods to insert pins. In the present study, we develop a mechanical fastening and cocuring (bonding) process that allows insertion of metallic pins in the overlap of hybrid mild-steel and glass fibre reinforced polymer (GFRP) single lap joints. Metallic pins of 0.3 mm in diameters were used in selected patterns and with and without clamping to assess their effects on the static failure loads and modes of the mild-steel/ glass-fibre prepreg co-cured interface. Clamping was introduced by using washers and bending both free ends of each pin. Experimental results show that the presence of the metallic pins can increase the static failure load by 58% when compared to that of adhesively bonded joints. The proposed hybrid joint can find its applications in aerospace, defence, automotive and oil and gas infrastructure. 2. Experimental 2.1. Materials The metal substrates used in this study were Bluescope XLERPLATE (AS/NZS 3678-2505) equivalent to A36. The plate was 3 mm thick mild carbon steel with a composition of 0.22% C, 1.7% Mn, 0.55% Si, 0.03% S, 0.1% Al, 0.04% Ti and 0.04% P and yield strength of 280 MPa and elongation at break of 23–45%. Stainless steel pins with 0.3 mm in diameter and 22 mm in length were sourced from Fine Science Tools, Canada. The GFRP prepreg used in this study comes in 300 mm width rolls and comprises a Duomat 750 E-glass fabric. Duomat 750 is made up of a 150 gsm layer of chopped stranded mat (CSM), attached to a woven roving composed of 350 gsm glass in the warp direction and 280 gsm glass in the weft direction, via 5 gsm of polyester stitching yarns. The GFRP fabric was pre-impregnated to a total areal weight of about 1730 gsm to produce a GFRP prepreg with a thickness of approximately 1 mm. An amine cured epoxy resin (out of water usable epoxy primer) was used between the GFRP prepreg and surfaced-prepared steel at a thickness of about 1150 gsm. 2.2. Methods 2.2.1. Surface preparation Based on the results of our previous study on six surface preparation methods, i.e. garnet grit blasting, disk sanding, needle gun, flap wheel, strip wheel and wire brush [18], garnet grit blasting technique was adopted in this study. Firstly, the steel plates were cleaned using acetone and lint free cloth. Then the top faces of the steel plates were prepared using a grit blasting technique with
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a garnet grit of 30–60 grit size (sourced from Burwell Technologies, Australia). The grit blasting was carried out using compressed air regulator set to 0.55 MPa, a nozzle angle of approximately 45°, and a nozzle diameter of 7 mm positioned approximately 100 mm from the surface under preparation using a Hafco Metalmaster SB-420 sand blasting cabinet. 2.2.2. Manufacture of SLJ specimens Two lines 24 mm apart were drawn from the edge along the width of a steel adherend plate (overlapping region of the plate) with 310 mm length, 105 mm width and 3 mm thickness using a scribe as shown in Fig. 1. Twenty five 0.5 mm holes were drilled on each line in such a way that they are located close to the overlap ends. First half of the steel plate had 15 holes (left hand side) on each line to accommodate three holes in each of the five specimens while second half of the plate had 10 holes (right hand side) on each of the lines to accommodate two holes on each of the five specimens. 33.5 mm space on the right hand side of the plate was left free of holes to be able to clamp it during water jet cutting. The choice of placing pins near overlap ends in a SLJ was based on our previous study of the stitched composite SLJ [8,16] and the stitched composite double cantilever beam (DCB) [17], in which stitches were placed near overlap ends. The steel substrate had a thickness of 3 mm, whilst the prepreg substrate thickness was approximately 6 mm. The thicknesses of the substrates differed from those prescribed in ASTM D5868 [16] due to the steel having a 3 mm minimum thickness, whilst the steel substrate thickness prescribed in the standard was 1.5 mm. The required GFRP adherend thickness was determined by bending rigidity of the steel and GFRP adherends in which the second moment of area was varied. The surface of the steel in the overlapping region was prepared in accordance with ISO standard 8501-1:2007 by grit blasting as described above. The overlapping region of the steel adherend plate was wiped down with a lint free cloth and placed on a 20 mm thick cork sheet placed on an aluminium tool plate covered with release film. The base steel spacer plate with 310 mm length, 150 mm width and also 3 mm thickness, was covered with release film, and placed into contact with the adherend plate on the cork sheet, such that the edges oriented along with width were placed into contact with one another. Single lap joints (SLJs) comprised of A36 steel and GFRP substrates, with an out of water usable epoxy resin primer in between, were manufactured for tests per guideline detailed in ASTM D5868 [16]. The upper Aluminium spacer plate with 310 mm length, 77 mm in width and 10 mm thickness, was covered in release film and placed on top of the steel adherend plate, such that the steel adherend plate had a distance of 25 mm in the length direction that was left uncovered. This upper spacer plate acted as a resin dam during consolidation and cure. Weights were placed on top of the aluminium plate during layup to prevent it from moving. The prepreg stack was laid up with laminate configuration of six ply [0°]6 (all in weft direction) and [90°]6 (all in warp direction) and with the CSM layer in contact with the primer. A GFRP adherend with [0°]6 or [90°]6 refers to the case where the weft or warp direction of all plies align with the loading direction of the SLJ. Pressure was applied using an 80 mm 20 mm diameter fin consolidation roller between each ply. A layer of release film was placed over laminate, such that it extended a minimum distance of 10 mm from each corner. 0.3 mm pins were then pushed through the 0.5 mm holes of the adherend plate. Cure temperature was 55 °C for 48 h. A steel plate of 25 mm thickness and of equal in-plane dimensions to the laminate was placed on top of the release film and underlying laminate to be used as weights to apply consolidation pressure of approximately 2 kPa to the surface of the flat laminate parts during the curing process which produced 0.2–0.4 mm thick adhesive
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Fig. 1. Dimensions of the steel plate and pin-hole arrangements. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
layer between the steel and GFRP adherends. The manufacturing set-up of the specimens is shown in Fig. 2. Once the sample curing was completed, protruded part of the half of the pins of the plate was cut-off from the surface of metallic and GFRP adherend ends. The other half of the protruded pins was clamped by putting 5 mm washer and bending the pins as shown in Figs. 3 and 4. For the bent pin specimens with six pins, four of the pins were bent in the loading direction and two of them were bent transverse to the loading direction. For the bent pin specimens with four pins, two of them were bent along the loading direction and two of them were bent transverse to the loading direction. The dimensions of the SLJ specimens obtained using water jet cutting are shown in Fig. 4. Approximately 10 specimens were obtained from each of the manufactured plates. The laminate transverse direction (90°) was aligned with the 185 mm side, hence, the joint loading direction. Approximately 50 mm was allocated for clamping. All the SLJ specimens are labelled using the format given in Table 1 where i represent the sample number varying from 1 to 6. B-4 and P-6 correspond to hybrid specimens with GFRP adherend of [0°]6 (all in weft direction) with 4 bent and 6 unbent pins respectively; B-6 and P-4 refer to specimens with GFRP adherend of [90°]6 (all in warp direction) with 6 bent and 4 unbent pins respectively. For example, C-0-1-weft means controlled specimen number 1 with weft direction aligned in the loading direction. A total of 20 specimens were fabricated for pinned hybrid SLJ testing with the dimensions given in Fig. 4. For controlled (without reinforcing pins) testing another 11 specimens with the same dimensions were manufactured using the same method described above and tested. 4 specimens for each category of pinned hybrid SLJ samples were tested.
using an Xradia MicroXCT-400 Micro-Computed Tomography machine. The pinned SLJ specimens were placed onto a platform vertically and placed into a chamber. X-ray energy source of 50 keV was used with an average scan time of 10 min for each specimen.
2.2.3. Micro-Computed Tomography (Micro-CT) analysis To determine the location and arrangements of the pins, 3D Micro-Computed Tomography (Micro-CT) analysis was performed
The load–displacement curves of all the controlled SLJ and hybrid SLJ specimens are shown in Fig. 7(a)–(f). From the curves, it can be seen that for the controlled SLJ specimens, C-1-i-warp
2.2.4. Single Lap Shear Joint (SLJ) testing The SLJ testing for the controlled and pinned specimens was performed in accordance with ASTM D5868-01. A 30 kN load cell with a crosshead speed of 0.5 mm/min was used. The SLJ testing for controlled and pinned reinforced samples were carried out at room temperature. Fig. 5 shows the set-up for SLJ testing. A gauge length of 135 mm was maintained leaving a total of 50 mm for gripping (25 mm on each end) for each of the specimens. 3. Results and discussion 3.1. Micro-CT scans for Pin location and alignment Due to the use of a hand pinning process in the current study, the difficulty in aligning the pins in the vertical direction through the joint was encountered. Fig. 6 presents the 3D view of the location and arrangement of the pins of the overlap joints of the pinned hybrid SLJ specimens that shows misalignment of the pins from the vertical direction but by a very low angle. The average misalignment from vertical direction was approximately 1° for most of the specimens although misalignment as high as 3° was also observed. The misalignment was found to be random relative to the vertical direction. The cause of the misalignment could be attributed to the pressure exerted on the pins during curing as well as cutting and bending of the pins after the curing process. 3.2. Load–displacement curves
310 mm x 105 mm x 25 mm steel mass
Glass fibre prepreg Epoxy resin primer
310 mm x 105 mm x 3 mm steel Aluminium tool plate
Fig. 2. Manufacture of SLJ test specimen.
Release Film
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4 mm trim width 3.5 mm
1 mm water jet cut thickness 10 mm
2 mm
Specimen width 25 mm
9 mm
12.5 mm
Sample length 310 mm
Pin diameter 25 mm ∅ = 0.3 mm overlap width
Composite width 105 mm
Steel width 105 mm
(a)
Specimen width 25 mm
30 mm clamp width for water jet cut
(b)
Fig. 3. (a) Dimensions of the manufactured plates and the dimensions of the specimens for water jet cutting and (b) Manufactured plates ready for water jet cutting. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
Fig. 4. Test configurations for SLJ specimen. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
SLJ specimens had higher load carrying capacity than the C-0-i-weft SLJ specimens. From the comparison of the controlled and hybrid SLJ specimens, it can be seen that the hybrid SLJ specimens had a higher load carrying capability than the controlled SLJ specimens, indicating that the inserted pin has a reinforcing effect
in the overlap region of the pinned SLJ specimens. It can also be seen that SLJ specimens with six bent pins in the warp direction (B-6-i-warp) had the greatest load carrying capability, followed by SLJ specimens with four unbent pins in the warp direction (P-4-i-warp), then SLJ specimens with six unbent pins in the weft
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Table 1 Labelling of the SLJ specimens. Specimen type
GFRP adherend [0°]6 (all in weft direction)
GFRP adherend [90°]6 (all in warp direction)
Controlled specimens without pin reinforcement Hybrid Specimens
C-0-i-weft
C-1-i-warp
B-4-i-weft P-6-i-weft
B-6-i-warp P-4-i-warp
direction (P-6-i-weft) and finally SLJ specimens with four bent pins in the weft direction (B-4-i-weft) which had the lowest load carrying capability. Table 1 can be seen for the description of the specimens. From this figure, it can also be seen that for all the specimens the failure mode was brittle in nature as can be identified by the sudden drop of the failure load of the curves. For all the controlled SLJ and pinned SLJ specimens, the curves were found to increase at a linear rate up to failure load except for the P-4-iwarp SLJ specimens. For the P-4-i-warp SLJ specimens, transient drop in the load-capacity of the joint was recorded for specimens 1 and 2, although the specimens quickly recovered and was able to withstand further loading. For specimens 3 and 4, gradual drop in the load-capacity of the joint was seen and these specimens were found to gradually recover and withstand further loading. These unbent four-pin specimens were found to eventually fail at a higher load and displacement limit. 3.3. Maximum loads and displacements Table 2 tabulates the average maximum failure loads, displacements and corresponding failure modes as well as one standard deviation for the controlled SLJ and hybrid SLJ specimens. For specimens C-0-i-weft, B-4-i-weft and P-6-i-weft with GFRP all in weft direction, the increase of the average maximum load is 51% for the P-6-i-weft specimens and 32% for the B-4-i-weft specimens, respectively, compared to that of the corresponding controlled specimen C-0-i-weft. For specimens C-1-i-warp, B-6-i-warp and P-4-i-warp with GFRP all in warp direction, the increase in the
average maximum load is 58% for the B-6-i-warp specimens and 57% for the P-4-i-warp specimens, respectively, compared to that of the corresponding controlled specimen C-1-i-warp. It is worth noting that the average load at first load-drop for P-4-i-warp is 8.30, which is 23% higher than that of C-1-i-warp. From Table 2, it can be seen that the greatest average displacement at break was found to be 3.64 mm for P-4-i-warp specimens, followed by 2.67 mm for B-6-i-warp specimens, then 2.08 mm B-4i-weft specimens, 1.82 mm P-6-i-weft specimens and finally 1.66 mm and 1.63 mm for the controlled C-1-i-warp and C-0-iweft specimens, which had the lowest maximum displacement at failure. The corresponding increase of the average maximum displacement at failure is 61% and 119% for B-6-i-warp and P-4-iwarp specimens comparing to that of C-1-i-warp specimens, and is 28% and 12% for B-4-i-weft and P-6-i-weft specimens when compared to that of C-0-i-weft specimens. As noted in Table 1, the average displacement at which the first load drop occurs for P-4i-warp specimens is 2.41 mm, which is only 45% larger than that of C-1-i-warp specimens. It is evident that the average maximum failure loads and displacements increase with the presence of pin reinforcement as shown by the results in Table 2 and Fig. 7. Specimen group B-6-iwarp exhibits the largest failure loads and displacements. The results in Table 2 also reveal that the average maximum failure loads for specimens with GFRP in weft or warp direction increase with the number of pins used for the two considered pin arrangement cases. This means that for specimens with GFRP in weft or warp direction the associated average maximum failure loads with 6 pins are larger than those with 4 pins, which are subsequently larger than the corresponding unpinned.
3.4. Failure modes and mechanisms The failure modes and strengthening mechanism of the pinned joints were found to vary along with the number of pins and whether the pins were bent or unbent i.e. whether the clamping force of the pins was present or not. Three types of failure modes for the controlled SLJ and pinned SLJ specimens were seen as shown in Fig. 8.
Fig. 5. Testing arrangements for the SLJ specimens. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. Pictures taken using Micro-CT scan showing the location and arrangements of the pins (a) B-6-3-warp (b) P-6-1-weft (c) B-4-1-weft and (d) P-4-1-warp.
Failure mode type 1 (Fig. 9(a)) or failure by bond line debonding by rapid delamination cracking that caused the separation of steel and prepreg adherents was observed for the controlled SLJ specimens. Approximately 3% and 5% cohesive failure was also observed for C-0-i-weft and C-1-i-warp specimens respectively.
Failure mode type 2 or failure by debonding of bond line and plastic shear deformation of the pins were observed for P-4-iwarp and P-6-i-weft SLJ specimens. P-4-i-warp SLJ specimens were found to have a much larger pull-out and shear deformation of the pins (Fig. 9(c)) than the P-6-i-weft SLJ specimens (Fig. 9(e)). For
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8
4 2 0
0
0.5
12
1.5
0
0.5
(b) 12
0.5 1 1.5 2 Displacement (mm)
8 6 4 2 0
2.5
0
0.5
1 1.5 2 Displacement (mm)
2.5
3
(d)
P-4-1-warp P-4-2-warp P-4-3-warp P-4-4-warp
12
P-6-1-weft P-6-2-weft P-6-3-weft P-6-4-weft
10
8
8 Load (kN)
Load (kN)
2
B-6-1-warp B-6-2-warp B-6-3-warp B-6-4-warp
10
(c)
10
1.5
(a)
2
12
1
Displacement (mm)
4
0
2
Displacement (mm)
6
0
4
0
2
Load (kN)
8 Load (kN)
1
B-4-1-weft B-4-2-weft B-4-3-weft B-4-4-weft
10
C-1-1-warp C-1-2-warp C-1-3-warp C-1-4-warp C-1-6-warp
6 Load (kN)
6 Load (kN)
8
C-0-1-weft C-0-2-weft C-0-3-weft C-0-4-weft C-0-5-weft
6 4 2
6 4 2
0
0 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 Displacement (mm)
0
0.5
1 1.5 2 Displacement (mm)
(e)
(f)
2.5
Fig. 7. Load–displacement curves of SLJ specimens: (a) C-0-i-weft, (b) C-1-i-warp, (c) B-4-i-weft, (d) B-6-i-warp, (e) P-4-i-warp and (f) P-6-i-weft. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
P-4-i-warp pins, all pins were pulled out with plastic shear deformation except specimen P-4-3-warp where one pin at the edge of resin end failed. Both interfacial and cohesive failure with matrix cracking and fibre breakage around the location of the pins was observed. For P-6-i-weft pins, small pull-outs and mostly interfacial debonding with a very little cohesive failure was observed. Fractured surface of these specimens were found to have debonding and matrix cracking around the location of the pins. Failure mode type 3 or failure by bond line debonding, fracture and pull-out of the pins were observed for B-4-i-weft and B-6-iwarp SLJ specimens with relatively more fracture and pull-out of the pins (Fig. 9(b)) were observed for B-6-i-warp specimens than B-4-i-weft specimens (Fig. 9(d)). In addition to that, fractured surface of the B-6-i-warp specimens was found to appear with larger
amount of matrix cracking and fibre breakage around the location of the pins than that of B-4-i-weft specimens. For B-4-i-weft specimens, all pins were intact with a bit pull out except specimen B-41-weft where both of the pins at the resin end failed with the pins at the steel end pulled out with plastic shear deformation. For B-6i-warp specimens, all three pins at the resin end failed for B-6-1warp and B-6-4-warp, one pin at the edge of resin end failed for B-6-2-warp and all six pins failed with pins at the steel end strained/pulled out for B-6-3-warp specimen. All of the other pins pulled out with plastic shear deformation. The failure mode type of each of the controlled SLJ and pinned SLJ specimens are summarised in Table 2. The transfer of the applied load between the steel and prepreg adherents by the pins is believed to be responsible for the
M.S. Islam, L. Tong / Composites: Part A 84 (2016) 196–208 Table 2 Average maximum failure loads and displacements and summary of failure modes for the controlled and pinned SLJ (values in the parentheses correspond to one standard deviation (STDV)) specimens. Specimen types
Mean of max. failure load in kN (STDV)
Mean of max displacement at failure in mm (STDV)
Failure mode types
C-0-i-weft C-1-i-warp Pa-4-i-warp P-6-i-weft B-4-i-weft B-6-i-warp
5.65 (0.19) 6.72 (0.56) 10.58 (0.05) 8.51 (0.15) 7.46 (0.70) 10.61 (0.42)
1.63 1.66 3.64 1.82 2.08 2.67
Type Type Type Type Type Type
(0.15) (0.05) (0.70) (0.13) (0.19) (0.12)
1 1 2 2 3 3
a Mean load and displacement of the first kink was 8.30(1.36) kN and 2.41(0.36) mm. The values in the parentheses are one standard deviation.
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carefully. From the figure it can be seen that for the P-4-i-warp specimens, the area around the pins has matrix cracking and fibre breakage along with debonding of the pins. For all other pinned SLJ specimens, no fibre breakage and matrix cracking was observed. The P-4-i-warp specimens had a large amount of shear deformation and pull-out of the pins which may be responsible for fibre breakage and matrix cracking around the location of the pins. For the P-6-i-weft specimens, small pin pull-out and delamination was observed indicating a smaller amount of pin pull-out and shear deformation for these specimens. On the other hand, for the bent specimens, only delamination around the pins was observed indicating larger tensile plastic deformation of the pins and a very small amount of pull-out of the pins. 3.5. Influence of pin arrangement
improvement in the shear strength of the pinned SLJ specimens when compared to the controlled SLJ specimens. The pinned SLJ specimens were failed by delamination cracking of the prepreg and adhesive and shear-induced rupture of the pins at the bond-line. Shear deformation was found to be very significant in the P-4-i-warp specimens. Failure started with the initiation of delamination cracks at both ends of the bonded region. The delamination grew a short distance into the bonded region, before being arrested by bridging pins. The initial growth and arrest of these cracks might be responsible for the initial transient and gradual load drop in the load–displacement curves. Growth of these delaminations reinstated when the applied load was increased above the load drop and subsequent crack growth along the bond-line was stabilised by the bridging action of the pins. When the applied load reached the maximum failure load, the delamination propagated rapidly along the remaining region of the bond-line, causing complete failure of the joint. At the maximum failure load, the pins were pulled out with shear deformation along the bond-line. SLJ specimens reinforced with six pins displayed similar strengthening mechanisms and failure mode as the four pin specimens. However, in contrast to low pin content, the higher density of pins permanently arrested delamination cracking of the bond-line and ultimate failure was achieved by adherent failure inside the pinned region. Fig. 10 shows the optical micrographs of the typical top surfaces of the prepreg of the pinned SLJ specimens. To look at the surface around the pins, the washer of the bent specimens were removed
Fig. 11 shows that majority of all the pins of the hybrid specimens are located within 3 mm from the edges of the overlap along the loading direction although the exact location of each pin varies. For example, as can be seen for P-6-i-weft specimens (Fig. 11(b), from the edge the average location of the four cornered pins (pin number 1, 3, 4 and 6) towards loading direction was found to be in 0.9, 21, 1.3 and 21.6 mm of the overlap region. Similar to the experimental results for stitched single lap joint [2,8] and double cantilever beam [17] tests, this distribution pattern of placing pins near overlap ends is considered to be the major determinant factor for the increase in the measured failure loads and displacements of the hybrid specimens. As well-known [2], the high peel stresses near overlap ends are believed to be a key factor affecting the static strength of an adhesively bonded single lap joint in tension, placing pins near overlap ends is thus more effective in reducing the peel stress and increasing the join strength. This is because the reinforcing effect of the pins and adhesive bonding might act simultaneously to raise the failure load above the controlled specimens. For B-4-i-weft, B-6i-warp and P-6-i-weft specimens, it is believed that, once the adhesive debonding takes place, the pins were unable to carry out any further load and hence the interfacial failure was the main failure mode of the specimens. In this case the balanced position of the presence of pins might have suppressed bond-line failure of joints for a while and helped increase the failure load. However, as shown in Fig. 7, specimens P-4-i-warp exhibit a nonlinear behaviour with a load-drop (kink) prior to reaching ultimate failure load. This is
Fig. 8. Type of failures observed for the pinned SLJ specimens.
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Fig. 9. Optical micrographs of the typical fractured surfaces of controlled (a) C-0-1-weft and pinned (b) B-6-3-warp (c) P-6-4-weft (d) B-4-1-weft and (e) P-4-1-warp SLJ specimens. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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205
Fig. 9 (continued)
quite different from B-4-i-weft, B-6-i-warp and P-6-i-weft specimens. As shown in Fig. 11(d) for P-4-i-warp, specimens 1 and 2 were found to have average pin locations (2.3, 14.3, 2.8 and 14.8) and specimens 3 and 4 were found to have average pin locations (8.5, 20.3, 4.3, 20.8) along the horizontal direction perpendicular to the load direction in the overlap region. Specimens 1 and 2 were found to follow a sharp initial kink and specimens 3 and 4 were found to follow a gradual kink in the load displacement curves. However, the load was found to pick up from the kink and the specimens were found to fail close to the average failure load of B-6-i-warp specimens (Fig. 11(a)). The similar load–displacement curves for specimens 1 and 2 and different load–displacement curve for specimens 3 and 4 support the dependency of the load on the location of the pins along the horizontal direction as well as the distance between the two pins at each overlap end. As shown in Fig. 11(c) and (d), the distance in horizontal direction between the two pins at each overlap end for P-4-i-warp is larger than that for B-4-i-weft, in addition to the difference in the hori-
zontal distance between a pin and the nearest free overlap edge. These are also regarded as a major contributing factor to the discrepancy between the load–displacement curves between specimens B-4-i-weft and P-4-i-warp. Due to strain hardening of steel during drilling holes in the steel plate, it was difficult to control the pin symmetry about the specimen centreline. However, as discussed above, in experimental testing minor errors in load and displacement was observed for centreline deviation of the location of the pins. 3.6. Relative efficiency of mechanical fastening in hybrid joints Since the material and geometrical parameters and loading conditions used in the literatures [5,6,11–13] for hybrid joints are different, the measured failure loads and load variation due to insertion of fastening vary significantly. Thus it is difficult to make a direct comparison amongst the reported data. Therefore, to measure the efficiency of mechanical fastening in a hybrid joint and to make a comparison amongst various hybrid joints, we introduce
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Fig. 10. Optical micrographs of the typical top surfaces of the prepreg of the pinned SLJ specimens (a) B-6-3-warp (b) P-6-1-weft (c) B-4-1-weft and P-4-1-warp. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
the following mechanical fastening efficiency factor based on relative change in measured apparent shear strength as used in ASTM standards such as D5868:
g¼
ðPh Pb ÞðAf þ Ab Þ ðrh rb ÞðAf þ Ab Þ ¼ rb Af P b Af
where g is a factor for relative efficiency of mechanical fastening in a hybrid joint, Ph and rh denotes the failure load or stress of the hybrid joint, P b and rb represents the failure load and stress of adhesively bonded joint (with no mechanical fastening), Af denotes the total cross sectional area of all the mechanical fasteners, and Ab represents the overall bonding area in a hybrid joint. It is evident that the term ðPh Pb Þ=Pb or ðrh rb Þ=rb represents the relative failure load or stress variation due to insertion of mechanical fastening in adhesively bonded joints, and the term Af =ðAf þ Ab Þ denotes the ratio of the mechanical fastening area to the overall overlap area. To certain extent this efficiency factor reflects the efficiency of mechanical fastening insertion in affecting the failure strength because of the inclusion of percentage changes
in failure loads/strengths in the definition of the factor. It is well known that both peel and shear stresses are present at the ends of an overlap of a balanced or unbalanced bonded joint, such as a single lap joint under tensile loading even aligned along the mid-plane of an adhesive layer or a double lap joint due to localized bending effect; and the pins placed at the end of an overlap play a role in reducing peak peel and shear stresses [16]. It is also worth noting that the measured apparent shear strength for single-lap and double-lap joints contains effects of various geometrical and material arrangements, such as adhesive thickness, mixed peel and shear effects. Table 3 lists the values of these four parameters and the calculated efficiency factors for the present hybrid joints and others reported in the literature. The comparison reveals that the pin distribution used in this study is highly efficient in terms of the static strength of hybrid joints. This is because for single-lap and doublelap joints in tension, maximum peel stress occurs near overlap ends. Thus it is naturally more effective to place pins or fasteners near overlap ends to counteract the peel stress as observed in previous studies [8,16,17].
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M.S. Islam, L. Tong / Composites: Part A 84 (2016) 196–208 0
0
Pin 2
Pin 1
Pin 3
Pin 3
B-6-1-warp B-6-2-warp B-6-3-warp B-6-4-warp
10
15
20
Pin 4
Pin 5
5
10
15
20
P-6-1-weft P-6-2-weft P-6-3-weft P-6-4-weft
Pin 4
Pin 6
25 0
Loading direction
15
Location of Pins (mm)
10
20
Pin 5
Pin 6
25
25
0
5
10
15
20
Location of Pins (mm)
Location of Pins (mm)
(a)
(b)
0
25
0
Pin 2
Pin 1 5
B-4-1-weft B-4-2-weft B-4-3-weft B-4-4-weft
Loading direction
15
Location of Pins (mm)
10
10
15
P-4-1-warp P-4-2-warp P-4-3-warp P-4-4-warp
Pin 4
Pin 3
20
20
Pin 2
Pin 1
5
Loading direction
Location of Pins (mm)
Pin 2
5
Loading direction
Location of Pins (mm)
Pin 1 5
Pin 4
Pin 3
25
25 0
5
10
15
20
25
0
5
10
15
Location of Pins (mm)
Location of Pins (mm)
(c)
(d)
20
25
Fig. 11. Diagram for the experimental location of the pins in the overlap region of the hybrid SLJ specimens.
Table 3 A list of the values of four parameters (Ph =rh ; P b =rb ; Af ; Ab ) and the calculated efficiency factor (g) for the present hybrid joints and others reported in the literature. References
P h (kN) /rh (MPa)
P b (kN) /rb (MPa)
Af (mm2)
Af þ Ab (mm2)
g
Present Parkes et al. [13] (Baseline geometry) Parkes et al. [13] (Alternate geome try) Matsuzaki et al. [6] Ucsnik et al. [11] Kweon et al. [5] (Film Type Adhesive) Kweon et al. [5] (Paste Type Adhesive) Graham et al. [12]
10.61 kN 25 kN
6.72 kN 5 kN
0.4 28.3
625.4 673.4
854 95
32.5 kN
5 kN
28.3
673.4
131
6.09 MPa 32.85 kN 440 MPa
3.37 MPa 21.5 kN 453 MPa
12.6 35.2 35.6
512. 6 1535.2 1326.0
33 23 1
192 MPa
67.1 MPa
35.6
1326.0
69
17.43 MPa
8.76 MPa
24.4
649.4
26
4. Conclusions In this study, the failure modes and shear strength were experimentally determined for mild steel–glass epoxy co-bonded composite hybrid single-lap joints. It has been demonstrated that pin reinforcement improves the static strength of hybrid single lap joints for steel glass fibre prepreg co-bonded composites. The major factor which influenced the static strength of the pinned SLJ specimens the most was whether the GFRP prepreg was in warp or weft direction. The dominant strengthening mechanism was crack bridging by the specimens with six pins, which increased the static strength by as much as 58%. The calculated
mechanical fastening efficiency factor revealed significantly high distribution efficiency of the selected pin arrangement in the current study. Further studies can be performed to assess the fatigue behaviours of the pinned SLJ composites, to investigate the effects of various geometrical and material parameters, such as adhesive thickness, pin arrangements and peak peel-shear stress mixity, as well as surface preparation methods on failure loads and modes. Acknowledgements This work was undertaken as part of a CRC-ACS research program, established and supported under the Australian Government’s Cooperative Research Centres Program. The authors would like to thank Dr Paul Falzon, General Manager of ACS, Australia Pty Ltd, Port Melbourne, Australia for his suggestions as Project Leader. LT would like to acknowledge the support of Australian Research Council (DP130103958). The authors also thankful to Mr Michael Marelli, Senior Technical Officer of CRC-ACS, Prestons, NSW for technical assistance and Mr Trevor Shearing, Senior Technical Officer of AMME, The University of Sydney for his assistance with testing. References [1] Hart-Smith LJ. Bonded-bolted composite joints. J Aircraft 1985;22 (11):993–1000. [2] Tong L, Mouritz AP, Bannister MK. 3D fiber reinforced polymer composites. Oxford, UK: Elsevier; 2002. [3] Fu M, Mallick PK. Fatigue of hybrid (adhesive/bolted) joints in SRIM composites. Int J Adhes Adhes 2001;21:145–59. [4] Kelly G. Load transfer in hybrid (bonded/bolted) composite single-lap joints. Compos Struct 2005;69:35–43.
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[5] Kweon J-H, Jung J-W, Kim T-H, Choi J-H, Kim D-H. Failure of carbon compositeto-aluminum joints with combined mechanical fastening and adhesive bonding. Compos Struct 2006;75(1–4):192–8. [6] Matsuzaki R, Shibata M, Todoroki A. Improving performance of GFRP/ Aluminum single lap joints using bolted/co-cured hybrid method. Compos A Appl Sci Manuf 2008;39:154–63. [7] Tong L. Bearing failure of composite bolted joints with non-uniform bolt-towasher clearance. Compos A Appl Sci Manuf 2000;31(6):609–15. [8] Tong L, Jain LK, Leong KH, Kelly D, Hertzberg I. Failure of transversely stitched RTM lap joints. Compos Sci Technol 1998;58(2):221–7. [9] Chang P, Mouritz AP, Cox BN. Tensile properties and failure mechanisms of zpinned composite lap joints. Compos Sci Technol 2006;66:2163–76. [10] Cartie DDR, Dell Anno G, Poulin E, Partridge IK. 3D reinforcement of stiffenerto-skin t-joints by z-pinning and tufting. Eng Fract Mech 2006;73:2532–40. [11] Ucsnik S, Scheerer M, Zaremba S, Pahr DH. Experimental investigation of a novel hybrid metal-composite joining technology. Compos A Appl Sci Manuf 2010;41:369–74. [12] Graham DP, Rezai A, Baker D, Smith PA, Watts JF. The development and scalability of a high strength, damage tolerant, hybrid joining scheme for composite–metal structures. Compos A Appl Sci Manuf 2014;64:11–24.
[13] Parkes PN, Butler R, Meyer J, de Oliveira A. Static strength of metal-composite joints with penetrative reinforcement. Compos Struct 2014;118:250–6. [14] Yang L, He XD, Mei L, Tong L, Wang RG, Li YB. Interfacial shear behavior of 3D composites reinforced with chemically connected CNTs. Compos A Appl Sci Manuf 2012;43(8):1410–8. [15] Tong L, Sun X, Tan P. Effect of long multi-walled carbon nanotubes on enhancing delamination toughness of laminated composites. J Compos Mater 2008;42(1):5–23. [16] Tong L, Steven GP. Analysis and design of structural bonded joints. Boston, USA: Kluwer Academic Publishers; 1999. [17] Wood MD, Sun X, Tong L, Kaltzos A, Rispler A, Mai Y-W. Effect of stitch distribution on mode I delamination toughness of stitched laminated composites – experimental results and simulation. Compos Sci Technol 2007;67(6):1058–72. [18] Islam MS, Tong L, Falzon PJ. Influence of metal surface preparation on its surface profile, contact angle, surface energy and adhesion with glass fibre prepreg. Int J Adhes Adhes 2014;51:32–41.
Contents lists available at ScienceDirect
Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Influence of pinning on static strength of co-cured metal-GFRP hybrid single lap joints M.S. Islam a,b, Liyong Tong a,⇑ a b
School of Aerospace Mechanical and Mechatronic Engineering, The University of Sydney, Sydney, NSW 2006, Australia Cooperative Research Centre for Advanced Composite Structures (CRC-ACS), 1/320 Lorimer Street, Port Melbourne, Victoria 3207, Australia
a r t i c l e
i n f o
Article history: Received 2 September 2015 Received in revised form 11 January 2016 Accepted 17 January 2016 Available online 23 January 2016 Keywords: A. Hybrid A. Polymer-matrix composites (PMCs) D. Mechanical testing E. Joints/joining
a b s t r a c t This study presents an experimental investigation into the effects of through-thickness pinning reinforcement on the static strength and damage tolerance of hybrid mild steel–glass fibre prepreg co-cured composite single lap joints (SLJ). Stainless steel pins of 0.3 mm in diameter were inserted as mechanical fastening, in addition to adhesive bonding, to form hybrid joints between metal and glass fibre reinforced polymer substrates. Using the hybrid SLJ tensile testing, the failure modes and static strength were experimentally determined for mild steel–glass fibre prepreg co-cured composites. It is revealed that pinning can improve the static failure load via bridging mechanism by as much as 58% depending on the number and location of pins and the presence of clamping due to bent-ends. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Hybrid joints use both mechanical fastening and adhesive bonding simultaneously to join similar or dissimilar structural components in order to increase joining strength and fatigue performance and to decrease weight and stress concentrations [1,2]. Single lap joints have been used to assess the benefits of combining mechanical fastening and adhesive bonding in terms of static tensile strength. The mechanical fastening entity used in composite–composite or metal–composite joints has different diameter scales ranging from bolts in several millimetres [3–7] to pins or stitches in sub millimetres [8–13] and even in nanometres [14,15]. In the previous studies using bolts of several millimetre diameters, single and multiple bolts were used in bonded-bolted single lap joints subjected to tensile load [3–6]. Single steel bolt of 6.35 mm diameter was used with adhesive bonding in the static and fatigue tests of adhesive/bolted composite joints by Fu and Mallick [3]. Kweon et al. [5] conducted experiments on tensile testing of hybrid bonded-bolted composite-to-aluminium double lap joints with single bolt (of 4.763 mm bolt hole diameter) in the centre of the overlap and reported an improved hybrid joint strength when mechanical fastening was stronger than adhesive bonding. Recently, Matsuzaki et al. [6] reported that hybrid GFRPaluminium single lap joints using bolted/co-cured hybrid method ⇑ Corresponding author. Tel.: +61 2 93516949. E-mail address: [email protected] (L. Tong). http://dx.doi.org/10.1016/j.compositesa.2016.01.011 1359-835X/Ó 2016 Elsevier Ltd. All rights reserved.
with patterned M2 12 mm steel bolts had 1.84 times higher maximum shear strength and a quarter of the standard deviation compared with the adhesive failure strength of the co-cured joints for GFRP/aluminium single lap joints. It is evident that bolt diameter used in the tests in [3–6] has been reduced from 6.35 mm to 2 mm. In addition to bolt diameter, other parameters, such as washer size, clamping force [3] and bolt-to-washer clearance [7], also play a role in affecting the joint strength of hybrid single lap joints. Technologies, such as stitching and pinning, have been an alternative method to introduce mechanical fastening using stitches and pins as bolts with small diameter on the order of millimetre, sub millimetre [2] or even nanometre [14,15]. They have been used in three dimensional reinforced composites and structures and also in structural joints of similar and dissimilar materials. An improvement in static strength of stitched and pinned composite-tocomposite single lap joints has been observed [8–10]. The improvement in joint strength is remarkable given the volume content of stitches or pins on the order of 1–5%, and it could be attributed to many factors including the crack bridging process or the clamping effect due to stitch pretension. Recently, various additive manufacturing techniques have been developed to produce mechanically pinned and adhesively bonded hybrid metal–composite single lap joints [11–13]. Ucsnik et al. [11] employed a cold metal transfer (CMT) process to arc weld small pins of 0.8 mm in diameter with and without ball-heads on metal surface and prepared double lap joints. Their test results show that the static tensile strength can be improved by 11% and 52% for the
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cases of pins without and with ball-heads. Graham et al. [12] used laser metal deposition (LMD) to add pins of 1 mm diameter in the form of an array of macro-scale projections for hybrid joining between metal and CFRP prepreg laminates. They have found an improvement of 60–157% in measured ultimate strength with hybrid joints over plain co-cured joints. Parkes et al. [13] employed the direct metal laser sintering process to build six by six pin (of 1 mm diameter) arrays from a bed of titanium powder and prepared hybrid titanium-CFRP single lap joints. Their results show that the static ultimate tensile strength can be improved by 6.5 times compared to adhesively bonded joints. Additive manufacturing processes often require special equipments and tools, usually involving intellectual property, and thus accessing to such technologies may be limited. To our best knowledge, there is no report available in open literature studying behaviour of hybrid joints with metallic pins with sub millimetre diameter inserted by using conventional mechanical fastening methods, such as bolting and riveting. Hence the aim of this study is to investigate the influence of metallic pins on static strength of hybrid joints by using traditional mechanical fastening methods to insert pins. In the present study, we develop a mechanical fastening and cocuring (bonding) process that allows insertion of metallic pins in the overlap of hybrid mild-steel and glass fibre reinforced polymer (GFRP) single lap joints. Metallic pins of 0.3 mm in diameters were used in selected patterns and with and without clamping to assess their effects on the static failure loads and modes of the mild-steel/ glass-fibre prepreg co-cured interface. Clamping was introduced by using washers and bending both free ends of each pin. Experimental results show that the presence of the metallic pins can increase the static failure load by 58% when compared to that of adhesively bonded joints. The proposed hybrid joint can find its applications in aerospace, defence, automotive and oil and gas infrastructure. 2. Experimental 2.1. Materials The metal substrates used in this study were Bluescope XLERPLATE (AS/NZS 3678-2505) equivalent to A36. The plate was 3 mm thick mild carbon steel with a composition of 0.22% C, 1.7% Mn, 0.55% Si, 0.03% S, 0.1% Al, 0.04% Ti and 0.04% P and yield strength of 280 MPa and elongation at break of 23–45%. Stainless steel pins with 0.3 mm in diameter and 22 mm in length were sourced from Fine Science Tools, Canada. The GFRP prepreg used in this study comes in 300 mm width rolls and comprises a Duomat 750 E-glass fabric. Duomat 750 is made up of a 150 gsm layer of chopped stranded mat (CSM), attached to a woven roving composed of 350 gsm glass in the warp direction and 280 gsm glass in the weft direction, via 5 gsm of polyester stitching yarns. The GFRP fabric was pre-impregnated to a total areal weight of about 1730 gsm to produce a GFRP prepreg with a thickness of approximately 1 mm. An amine cured epoxy resin (out of water usable epoxy primer) was used between the GFRP prepreg and surfaced-prepared steel at a thickness of about 1150 gsm. 2.2. Methods 2.2.1. Surface preparation Based on the results of our previous study on six surface preparation methods, i.e. garnet grit blasting, disk sanding, needle gun, flap wheel, strip wheel and wire brush [18], garnet grit blasting technique was adopted in this study. Firstly, the steel plates were cleaned using acetone and lint free cloth. Then the top faces of the steel plates were prepared using a grit blasting technique with
197
a garnet grit of 30–60 grit size (sourced from Burwell Technologies, Australia). The grit blasting was carried out using compressed air regulator set to 0.55 MPa, a nozzle angle of approximately 45°, and a nozzle diameter of 7 mm positioned approximately 100 mm from the surface under preparation using a Hafco Metalmaster SB-420 sand blasting cabinet. 2.2.2. Manufacture of SLJ specimens Two lines 24 mm apart were drawn from the edge along the width of a steel adherend plate (overlapping region of the plate) with 310 mm length, 105 mm width and 3 mm thickness using a scribe as shown in Fig. 1. Twenty five 0.5 mm holes were drilled on each line in such a way that they are located close to the overlap ends. First half of the steel plate had 15 holes (left hand side) on each line to accommodate three holes in each of the five specimens while second half of the plate had 10 holes (right hand side) on each of the lines to accommodate two holes on each of the five specimens. 33.5 mm space on the right hand side of the plate was left free of holes to be able to clamp it during water jet cutting. The choice of placing pins near overlap ends in a SLJ was based on our previous study of the stitched composite SLJ [8,16] and the stitched composite double cantilever beam (DCB) [17], in which stitches were placed near overlap ends. The steel substrate had a thickness of 3 mm, whilst the prepreg substrate thickness was approximately 6 mm. The thicknesses of the substrates differed from those prescribed in ASTM D5868 [16] due to the steel having a 3 mm minimum thickness, whilst the steel substrate thickness prescribed in the standard was 1.5 mm. The required GFRP adherend thickness was determined by bending rigidity of the steel and GFRP adherends in which the second moment of area was varied. The surface of the steel in the overlapping region was prepared in accordance with ISO standard 8501-1:2007 by grit blasting as described above. The overlapping region of the steel adherend plate was wiped down with a lint free cloth and placed on a 20 mm thick cork sheet placed on an aluminium tool plate covered with release film. The base steel spacer plate with 310 mm length, 150 mm width and also 3 mm thickness, was covered with release film, and placed into contact with the adherend plate on the cork sheet, such that the edges oriented along with width were placed into contact with one another. Single lap joints (SLJs) comprised of A36 steel and GFRP substrates, with an out of water usable epoxy resin primer in between, were manufactured for tests per guideline detailed in ASTM D5868 [16]. The upper Aluminium spacer plate with 310 mm length, 77 mm in width and 10 mm thickness, was covered in release film and placed on top of the steel adherend plate, such that the steel adherend plate had a distance of 25 mm in the length direction that was left uncovered. This upper spacer plate acted as a resin dam during consolidation and cure. Weights were placed on top of the aluminium plate during layup to prevent it from moving. The prepreg stack was laid up with laminate configuration of six ply [0°]6 (all in weft direction) and [90°]6 (all in warp direction) and with the CSM layer in contact with the primer. A GFRP adherend with [0°]6 or [90°]6 refers to the case where the weft or warp direction of all plies align with the loading direction of the SLJ. Pressure was applied using an 80 mm 20 mm diameter fin consolidation roller between each ply. A layer of release film was placed over laminate, such that it extended a minimum distance of 10 mm from each corner. 0.3 mm pins were then pushed through the 0.5 mm holes of the adherend plate. Cure temperature was 55 °C for 48 h. A steel plate of 25 mm thickness and of equal in-plane dimensions to the laminate was placed on top of the release film and underlying laminate to be used as weights to apply consolidation pressure of approximately 2 kPa to the surface of the flat laminate parts during the curing process which produced 0.2–0.4 mm thick adhesive
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Fig. 1. Dimensions of the steel plate and pin-hole arrangements. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
layer between the steel and GFRP adherends. The manufacturing set-up of the specimens is shown in Fig. 2. Once the sample curing was completed, protruded part of the half of the pins of the plate was cut-off from the surface of metallic and GFRP adherend ends. The other half of the protruded pins was clamped by putting 5 mm washer and bending the pins as shown in Figs. 3 and 4. For the bent pin specimens with six pins, four of the pins were bent in the loading direction and two of them were bent transverse to the loading direction. For the bent pin specimens with four pins, two of them were bent along the loading direction and two of them were bent transverse to the loading direction. The dimensions of the SLJ specimens obtained using water jet cutting are shown in Fig. 4. Approximately 10 specimens were obtained from each of the manufactured plates. The laminate transverse direction (90°) was aligned with the 185 mm side, hence, the joint loading direction. Approximately 50 mm was allocated for clamping. All the SLJ specimens are labelled using the format given in Table 1 where i represent the sample number varying from 1 to 6. B-4 and P-6 correspond to hybrid specimens with GFRP adherend of [0°]6 (all in weft direction) with 4 bent and 6 unbent pins respectively; B-6 and P-4 refer to specimens with GFRP adherend of [90°]6 (all in warp direction) with 6 bent and 4 unbent pins respectively. For example, C-0-1-weft means controlled specimen number 1 with weft direction aligned in the loading direction. A total of 20 specimens were fabricated for pinned hybrid SLJ testing with the dimensions given in Fig. 4. For controlled (without reinforcing pins) testing another 11 specimens with the same dimensions were manufactured using the same method described above and tested. 4 specimens for each category of pinned hybrid SLJ samples were tested.
using an Xradia MicroXCT-400 Micro-Computed Tomography machine. The pinned SLJ specimens were placed onto a platform vertically and placed into a chamber. X-ray energy source of 50 keV was used with an average scan time of 10 min for each specimen.
2.2.3. Micro-Computed Tomography (Micro-CT) analysis To determine the location and arrangements of the pins, 3D Micro-Computed Tomography (Micro-CT) analysis was performed
The load–displacement curves of all the controlled SLJ and hybrid SLJ specimens are shown in Fig. 7(a)–(f). From the curves, it can be seen that for the controlled SLJ specimens, C-1-i-warp
2.2.4. Single Lap Shear Joint (SLJ) testing The SLJ testing for the controlled and pinned specimens was performed in accordance with ASTM D5868-01. A 30 kN load cell with a crosshead speed of 0.5 mm/min was used. The SLJ testing for controlled and pinned reinforced samples were carried out at room temperature. Fig. 5 shows the set-up for SLJ testing. A gauge length of 135 mm was maintained leaving a total of 50 mm for gripping (25 mm on each end) for each of the specimens. 3. Results and discussion 3.1. Micro-CT scans for Pin location and alignment Due to the use of a hand pinning process in the current study, the difficulty in aligning the pins in the vertical direction through the joint was encountered. Fig. 6 presents the 3D view of the location and arrangement of the pins of the overlap joints of the pinned hybrid SLJ specimens that shows misalignment of the pins from the vertical direction but by a very low angle. The average misalignment from vertical direction was approximately 1° for most of the specimens although misalignment as high as 3° was also observed. The misalignment was found to be random relative to the vertical direction. The cause of the misalignment could be attributed to the pressure exerted on the pins during curing as well as cutting and bending of the pins after the curing process. 3.2. Load–displacement curves
310 mm x 105 mm x 25 mm steel mass
Glass fibre prepreg Epoxy resin primer
310 mm x 105 mm x 3 mm steel Aluminium tool plate
Fig. 2. Manufacture of SLJ test specimen.
Release Film
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4 mm trim width 3.5 mm
1 mm water jet cut thickness 10 mm
2 mm
Specimen width 25 mm
9 mm
12.5 mm
Sample length 310 mm
Pin diameter 25 mm ∅ = 0.3 mm overlap width
Composite width 105 mm
Steel width 105 mm
(a)
Specimen width 25 mm
30 mm clamp width for water jet cut
(b)
Fig. 3. (a) Dimensions of the manufactured plates and the dimensions of the specimens for water jet cutting and (b) Manufactured plates ready for water jet cutting. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
Fig. 4. Test configurations for SLJ specimen. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
SLJ specimens had higher load carrying capacity than the C-0-i-weft SLJ specimens. From the comparison of the controlled and hybrid SLJ specimens, it can be seen that the hybrid SLJ specimens had a higher load carrying capability than the controlled SLJ specimens, indicating that the inserted pin has a reinforcing effect
in the overlap region of the pinned SLJ specimens. It can also be seen that SLJ specimens with six bent pins in the warp direction (B-6-i-warp) had the greatest load carrying capability, followed by SLJ specimens with four unbent pins in the warp direction (P-4-i-warp), then SLJ specimens with six unbent pins in the weft
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Table 1 Labelling of the SLJ specimens. Specimen type
GFRP adherend [0°]6 (all in weft direction)
GFRP adherend [90°]6 (all in warp direction)
Controlled specimens without pin reinforcement Hybrid Specimens
C-0-i-weft
C-1-i-warp
B-4-i-weft P-6-i-weft
B-6-i-warp P-4-i-warp
direction (P-6-i-weft) and finally SLJ specimens with four bent pins in the weft direction (B-4-i-weft) which had the lowest load carrying capability. Table 1 can be seen for the description of the specimens. From this figure, it can also be seen that for all the specimens the failure mode was brittle in nature as can be identified by the sudden drop of the failure load of the curves. For all the controlled SLJ and pinned SLJ specimens, the curves were found to increase at a linear rate up to failure load except for the P-4-iwarp SLJ specimens. For the P-4-i-warp SLJ specimens, transient drop in the load-capacity of the joint was recorded for specimens 1 and 2, although the specimens quickly recovered and was able to withstand further loading. For specimens 3 and 4, gradual drop in the load-capacity of the joint was seen and these specimens were found to gradually recover and withstand further loading. These unbent four-pin specimens were found to eventually fail at a higher load and displacement limit. 3.3. Maximum loads and displacements Table 2 tabulates the average maximum failure loads, displacements and corresponding failure modes as well as one standard deviation for the controlled SLJ and hybrid SLJ specimens. For specimens C-0-i-weft, B-4-i-weft and P-6-i-weft with GFRP all in weft direction, the increase of the average maximum load is 51% for the P-6-i-weft specimens and 32% for the B-4-i-weft specimens, respectively, compared to that of the corresponding controlled specimen C-0-i-weft. For specimens C-1-i-warp, B-6-i-warp and P-4-i-warp with GFRP all in warp direction, the increase in the
average maximum load is 58% for the B-6-i-warp specimens and 57% for the P-4-i-warp specimens, respectively, compared to that of the corresponding controlled specimen C-1-i-warp. It is worth noting that the average load at first load-drop for P-4-i-warp is 8.30, which is 23% higher than that of C-1-i-warp. From Table 2, it can be seen that the greatest average displacement at break was found to be 3.64 mm for P-4-i-warp specimens, followed by 2.67 mm for B-6-i-warp specimens, then 2.08 mm B-4i-weft specimens, 1.82 mm P-6-i-weft specimens and finally 1.66 mm and 1.63 mm for the controlled C-1-i-warp and C-0-iweft specimens, which had the lowest maximum displacement at failure. The corresponding increase of the average maximum displacement at failure is 61% and 119% for B-6-i-warp and P-4-iwarp specimens comparing to that of C-1-i-warp specimens, and is 28% and 12% for B-4-i-weft and P-6-i-weft specimens when compared to that of C-0-i-weft specimens. As noted in Table 1, the average displacement at which the first load drop occurs for P-4i-warp specimens is 2.41 mm, which is only 45% larger than that of C-1-i-warp specimens. It is evident that the average maximum failure loads and displacements increase with the presence of pin reinforcement as shown by the results in Table 2 and Fig. 7. Specimen group B-6-iwarp exhibits the largest failure loads and displacements. The results in Table 2 also reveal that the average maximum failure loads for specimens with GFRP in weft or warp direction increase with the number of pins used for the two considered pin arrangement cases. This means that for specimens with GFRP in weft or warp direction the associated average maximum failure loads with 6 pins are larger than those with 4 pins, which are subsequently larger than the corresponding unpinned.
3.4. Failure modes and mechanisms The failure modes and strengthening mechanism of the pinned joints were found to vary along with the number of pins and whether the pins were bent or unbent i.e. whether the clamping force of the pins was present or not. Three types of failure modes for the controlled SLJ and pinned SLJ specimens were seen as shown in Fig. 8.
Fig. 5. Testing arrangements for the SLJ specimens. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. Pictures taken using Micro-CT scan showing the location and arrangements of the pins (a) B-6-3-warp (b) P-6-1-weft (c) B-4-1-weft and (d) P-4-1-warp.
Failure mode type 1 (Fig. 9(a)) or failure by bond line debonding by rapid delamination cracking that caused the separation of steel and prepreg adherents was observed for the controlled SLJ specimens. Approximately 3% and 5% cohesive failure was also observed for C-0-i-weft and C-1-i-warp specimens respectively.
Failure mode type 2 or failure by debonding of bond line and plastic shear deformation of the pins were observed for P-4-iwarp and P-6-i-weft SLJ specimens. P-4-i-warp SLJ specimens were found to have a much larger pull-out and shear deformation of the pins (Fig. 9(c)) than the P-6-i-weft SLJ specimens (Fig. 9(e)). For
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8
4 2 0
0
0.5
12
1.5
0
0.5
(b) 12
0.5 1 1.5 2 Displacement (mm)
8 6 4 2 0
2.5
0
0.5
1 1.5 2 Displacement (mm)
2.5
3
(d)
P-4-1-warp P-4-2-warp P-4-3-warp P-4-4-warp
12
P-6-1-weft P-6-2-weft P-6-3-weft P-6-4-weft
10
8
8 Load (kN)
Load (kN)
2
B-6-1-warp B-6-2-warp B-6-3-warp B-6-4-warp
10
(c)
10
1.5
(a)
2
12
1
Displacement (mm)
4
0
2
Displacement (mm)
6
0
4
0
2
Load (kN)
8 Load (kN)
1
B-4-1-weft B-4-2-weft B-4-3-weft B-4-4-weft
10
C-1-1-warp C-1-2-warp C-1-3-warp C-1-4-warp C-1-6-warp
6 Load (kN)
6 Load (kN)
8
C-0-1-weft C-0-2-weft C-0-3-weft C-0-4-weft C-0-5-weft
6 4 2
6 4 2
0
0 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 Displacement (mm)
0
0.5
1 1.5 2 Displacement (mm)
(e)
(f)
2.5
Fig. 7. Load–displacement curves of SLJ specimens: (a) C-0-i-weft, (b) C-1-i-warp, (c) B-4-i-weft, (d) B-6-i-warp, (e) P-4-i-warp and (f) P-6-i-weft. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
P-4-i-warp pins, all pins were pulled out with plastic shear deformation except specimen P-4-3-warp where one pin at the edge of resin end failed. Both interfacial and cohesive failure with matrix cracking and fibre breakage around the location of the pins was observed. For P-6-i-weft pins, small pull-outs and mostly interfacial debonding with a very little cohesive failure was observed. Fractured surface of these specimens were found to have debonding and matrix cracking around the location of the pins. Failure mode type 3 or failure by bond line debonding, fracture and pull-out of the pins were observed for B-4-i-weft and B-6-iwarp SLJ specimens with relatively more fracture and pull-out of the pins (Fig. 9(b)) were observed for B-6-i-warp specimens than B-4-i-weft specimens (Fig. 9(d)). In addition to that, fractured surface of the B-6-i-warp specimens was found to appear with larger
amount of matrix cracking and fibre breakage around the location of the pins than that of B-4-i-weft specimens. For B-4-i-weft specimens, all pins were intact with a bit pull out except specimen B-41-weft where both of the pins at the resin end failed with the pins at the steel end pulled out with plastic shear deformation. For B-6i-warp specimens, all three pins at the resin end failed for B-6-1warp and B-6-4-warp, one pin at the edge of resin end failed for B-6-2-warp and all six pins failed with pins at the steel end strained/pulled out for B-6-3-warp specimen. All of the other pins pulled out with plastic shear deformation. The failure mode type of each of the controlled SLJ and pinned SLJ specimens are summarised in Table 2. The transfer of the applied load between the steel and prepreg adherents by the pins is believed to be responsible for the
M.S. Islam, L. Tong / Composites: Part A 84 (2016) 196–208 Table 2 Average maximum failure loads and displacements and summary of failure modes for the controlled and pinned SLJ (values in the parentheses correspond to one standard deviation (STDV)) specimens. Specimen types
Mean of max. failure load in kN (STDV)
Mean of max displacement at failure in mm (STDV)
Failure mode types
C-0-i-weft C-1-i-warp Pa-4-i-warp P-6-i-weft B-4-i-weft B-6-i-warp
5.65 (0.19) 6.72 (0.56) 10.58 (0.05) 8.51 (0.15) 7.46 (0.70) 10.61 (0.42)
1.63 1.66 3.64 1.82 2.08 2.67
Type Type Type Type Type Type
(0.15) (0.05) (0.70) (0.13) (0.19) (0.12)
1 1 2 2 3 3
a Mean load and displacement of the first kink was 8.30(1.36) kN and 2.41(0.36) mm. The values in the parentheses are one standard deviation.
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carefully. From the figure it can be seen that for the P-4-i-warp specimens, the area around the pins has matrix cracking and fibre breakage along with debonding of the pins. For all other pinned SLJ specimens, no fibre breakage and matrix cracking was observed. The P-4-i-warp specimens had a large amount of shear deformation and pull-out of the pins which may be responsible for fibre breakage and matrix cracking around the location of the pins. For the P-6-i-weft specimens, small pin pull-out and delamination was observed indicating a smaller amount of pin pull-out and shear deformation for these specimens. On the other hand, for the bent specimens, only delamination around the pins was observed indicating larger tensile plastic deformation of the pins and a very small amount of pull-out of the pins. 3.5. Influence of pin arrangement
improvement in the shear strength of the pinned SLJ specimens when compared to the controlled SLJ specimens. The pinned SLJ specimens were failed by delamination cracking of the prepreg and adhesive and shear-induced rupture of the pins at the bond-line. Shear deformation was found to be very significant in the P-4-i-warp specimens. Failure started with the initiation of delamination cracks at both ends of the bonded region. The delamination grew a short distance into the bonded region, before being arrested by bridging pins. The initial growth and arrest of these cracks might be responsible for the initial transient and gradual load drop in the load–displacement curves. Growth of these delaminations reinstated when the applied load was increased above the load drop and subsequent crack growth along the bond-line was stabilised by the bridging action of the pins. When the applied load reached the maximum failure load, the delamination propagated rapidly along the remaining region of the bond-line, causing complete failure of the joint. At the maximum failure load, the pins were pulled out with shear deformation along the bond-line. SLJ specimens reinforced with six pins displayed similar strengthening mechanisms and failure mode as the four pin specimens. However, in contrast to low pin content, the higher density of pins permanently arrested delamination cracking of the bond-line and ultimate failure was achieved by adherent failure inside the pinned region. Fig. 10 shows the optical micrographs of the typical top surfaces of the prepreg of the pinned SLJ specimens. To look at the surface around the pins, the washer of the bent specimens were removed
Fig. 11 shows that majority of all the pins of the hybrid specimens are located within 3 mm from the edges of the overlap along the loading direction although the exact location of each pin varies. For example, as can be seen for P-6-i-weft specimens (Fig. 11(b), from the edge the average location of the four cornered pins (pin number 1, 3, 4 and 6) towards loading direction was found to be in 0.9, 21, 1.3 and 21.6 mm of the overlap region. Similar to the experimental results for stitched single lap joint [2,8] and double cantilever beam [17] tests, this distribution pattern of placing pins near overlap ends is considered to be the major determinant factor for the increase in the measured failure loads and displacements of the hybrid specimens. As well-known [2], the high peel stresses near overlap ends are believed to be a key factor affecting the static strength of an adhesively bonded single lap joint in tension, placing pins near overlap ends is thus more effective in reducing the peel stress and increasing the join strength. This is because the reinforcing effect of the pins and adhesive bonding might act simultaneously to raise the failure load above the controlled specimens. For B-4-i-weft, B-6i-warp and P-6-i-weft specimens, it is believed that, once the adhesive debonding takes place, the pins were unable to carry out any further load and hence the interfacial failure was the main failure mode of the specimens. In this case the balanced position of the presence of pins might have suppressed bond-line failure of joints for a while and helped increase the failure load. However, as shown in Fig. 7, specimens P-4-i-warp exhibit a nonlinear behaviour with a load-drop (kink) prior to reaching ultimate failure load. This is
Fig. 8. Type of failures observed for the pinned SLJ specimens.
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Fig. 9. Optical micrographs of the typical fractured surfaces of controlled (a) C-0-1-weft and pinned (b) B-6-3-warp (c) P-6-4-weft (d) B-4-1-weft and (e) P-4-1-warp SLJ specimens. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 9 (continued)
quite different from B-4-i-weft, B-6-i-warp and P-6-i-weft specimens. As shown in Fig. 11(d) for P-4-i-warp, specimens 1 and 2 were found to have average pin locations (2.3, 14.3, 2.8 and 14.8) and specimens 3 and 4 were found to have average pin locations (8.5, 20.3, 4.3, 20.8) along the horizontal direction perpendicular to the load direction in the overlap region. Specimens 1 and 2 were found to follow a sharp initial kink and specimens 3 and 4 were found to follow a gradual kink in the load displacement curves. However, the load was found to pick up from the kink and the specimens were found to fail close to the average failure load of B-6-i-warp specimens (Fig. 11(a)). The similar load–displacement curves for specimens 1 and 2 and different load–displacement curve for specimens 3 and 4 support the dependency of the load on the location of the pins along the horizontal direction as well as the distance between the two pins at each overlap end. As shown in Fig. 11(c) and (d), the distance in horizontal direction between the two pins at each overlap end for P-4-i-warp is larger than that for B-4-i-weft, in addition to the difference in the hori-
zontal distance between a pin and the nearest free overlap edge. These are also regarded as a major contributing factor to the discrepancy between the load–displacement curves between specimens B-4-i-weft and P-4-i-warp. Due to strain hardening of steel during drilling holes in the steel plate, it was difficult to control the pin symmetry about the specimen centreline. However, as discussed above, in experimental testing minor errors in load and displacement was observed for centreline deviation of the location of the pins. 3.6. Relative efficiency of mechanical fastening in hybrid joints Since the material and geometrical parameters and loading conditions used in the literatures [5,6,11–13] for hybrid joints are different, the measured failure loads and load variation due to insertion of fastening vary significantly. Thus it is difficult to make a direct comparison amongst the reported data. Therefore, to measure the efficiency of mechanical fastening in a hybrid joint and to make a comparison amongst various hybrid joints, we introduce
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Fig. 10. Optical micrographs of the typical top surfaces of the prepreg of the pinned SLJ specimens (a) B-6-3-warp (b) P-6-1-weft (c) B-4-1-weft and P-4-1-warp. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
the following mechanical fastening efficiency factor based on relative change in measured apparent shear strength as used in ASTM standards such as D5868:
g¼
ðPh Pb ÞðAf þ Ab Þ ðrh rb ÞðAf þ Ab Þ ¼ rb Af P b Af
where g is a factor for relative efficiency of mechanical fastening in a hybrid joint, Ph and rh denotes the failure load or stress of the hybrid joint, P b and rb represents the failure load and stress of adhesively bonded joint (with no mechanical fastening), Af denotes the total cross sectional area of all the mechanical fasteners, and Ab represents the overall bonding area in a hybrid joint. It is evident that the term ðPh Pb Þ=Pb or ðrh rb Þ=rb represents the relative failure load or stress variation due to insertion of mechanical fastening in adhesively bonded joints, and the term Af =ðAf þ Ab Þ denotes the ratio of the mechanical fastening area to the overall overlap area. To certain extent this efficiency factor reflects the efficiency of mechanical fastening insertion in affecting the failure strength because of the inclusion of percentage changes
in failure loads/strengths in the definition of the factor. It is well known that both peel and shear stresses are present at the ends of an overlap of a balanced or unbalanced bonded joint, such as a single lap joint under tensile loading even aligned along the mid-plane of an adhesive layer or a double lap joint due to localized bending effect; and the pins placed at the end of an overlap play a role in reducing peak peel and shear stresses [16]. It is also worth noting that the measured apparent shear strength for single-lap and double-lap joints contains effects of various geometrical and material arrangements, such as adhesive thickness, mixed peel and shear effects. Table 3 lists the values of these four parameters and the calculated efficiency factors for the present hybrid joints and others reported in the literature. The comparison reveals that the pin distribution used in this study is highly efficient in terms of the static strength of hybrid joints. This is because for single-lap and doublelap joints in tension, maximum peel stress occurs near overlap ends. Thus it is naturally more effective to place pins or fasteners near overlap ends to counteract the peel stress as observed in previous studies [8,16,17].
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0
Pin 2
Pin 1
Pin 3
Pin 3
B-6-1-warp B-6-2-warp B-6-3-warp B-6-4-warp
10
15
20
Pin 4
Pin 5
5
10
15
20
P-6-1-weft P-6-2-weft P-6-3-weft P-6-4-weft
Pin 4
Pin 6
25 0
Loading direction
15
Location of Pins (mm)
10
20
Pin 5
Pin 6
25
25
0
5
10
15
20
Location of Pins (mm)
Location of Pins (mm)
(a)
(b)
0
25
0
Pin 2
Pin 1 5
B-4-1-weft B-4-2-weft B-4-3-weft B-4-4-weft
Loading direction
15
Location of Pins (mm)
10
10
15
P-4-1-warp P-4-2-warp P-4-3-warp P-4-4-warp
Pin 4
Pin 3
20
20
Pin 2
Pin 1
5
Loading direction
Location of Pins (mm)
Pin 2
5
Loading direction
Location of Pins (mm)
Pin 1 5
Pin 4
Pin 3
25
25 0
5
10
15
20
25
0
5
10
15
Location of Pins (mm)
Location of Pins (mm)
(c)
(d)
20
25
Fig. 11. Diagram for the experimental location of the pins in the overlap region of the hybrid SLJ specimens.
Table 3 A list of the values of four parameters (Ph =rh ; P b =rb ; Af ; Ab ) and the calculated efficiency factor (g) for the present hybrid joints and others reported in the literature. References
P h (kN) /rh (MPa)
P b (kN) /rb (MPa)
Af (mm2)
Af þ Ab (mm2)
g
Present Parkes et al. [13] (Baseline geometry) Parkes et al. [13] (Alternate geome try) Matsuzaki et al. [6] Ucsnik et al. [11] Kweon et al. [5] (Film Type Adhesive) Kweon et al. [5] (Paste Type Adhesive) Graham et al. [12]
10.61 kN 25 kN
6.72 kN 5 kN
0.4 28.3
625.4 673.4
854 95
32.5 kN
5 kN
28.3
673.4
131
6.09 MPa 32.85 kN 440 MPa
3.37 MPa 21.5 kN 453 MPa
12.6 35.2 35.6
512. 6 1535.2 1326.0
33 23 1
192 MPa
67.1 MPa
35.6
1326.0
69
17.43 MPa
8.76 MPa
24.4
649.4
26
4. Conclusions In this study, the failure modes and shear strength were experimentally determined for mild steel–glass epoxy co-bonded composite hybrid single-lap joints. It has been demonstrated that pin reinforcement improves the static strength of hybrid single lap joints for steel glass fibre prepreg co-bonded composites. The major factor which influenced the static strength of the pinned SLJ specimens the most was whether the GFRP prepreg was in warp or weft direction. The dominant strengthening mechanism was crack bridging by the specimens with six pins, which increased the static strength by as much as 58%. The calculated
mechanical fastening efficiency factor revealed significantly high distribution efficiency of the selected pin arrangement in the current study. Further studies can be performed to assess the fatigue behaviours of the pinned SLJ composites, to investigate the effects of various geometrical and material parameters, such as adhesive thickness, pin arrangements and peak peel-shear stress mixity, as well as surface preparation methods on failure loads and modes. Acknowledgements This work was undertaken as part of a CRC-ACS research program, established and supported under the Australian Government’s Cooperative Research Centres Program. The authors would like to thank Dr Paul Falzon, General Manager of ACS, Australia Pty Ltd, Port Melbourne, Australia for his suggestions as Project Leader. LT would like to acknowledge the support of Australian Research Council (DP130103958). The authors also thankful to Mr Michael Marelli, Senior Technical Officer of CRC-ACS, Prestons, NSW for technical assistance and Mr Trevor Shearing, Senior Technical Officer of AMME, The University of Sydney for his assistance with testing. References [1] Hart-Smith LJ. Bonded-bolted composite joints. J Aircraft 1985;22 (11):993–1000. [2] Tong L, Mouritz AP, Bannister MK. 3D fiber reinforced polymer composites. Oxford, UK: Elsevier; 2002. [3] Fu M, Mallick PK. Fatigue of hybrid (adhesive/bolted) joints in SRIM composites. Int J Adhes Adhes 2001;21:145–59. [4] Kelly G. Load transfer in hybrid (bonded/bolted) composite single-lap joints. Compos Struct 2005;69:35–43.
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[5] Kweon J-H, Jung J-W, Kim T-H, Choi J-H, Kim D-H. Failure of carbon compositeto-aluminum joints with combined mechanical fastening and adhesive bonding. Compos Struct 2006;75(1–4):192–8. [6] Matsuzaki R, Shibata M, Todoroki A. Improving performance of GFRP/ Aluminum single lap joints using bolted/co-cured hybrid method. Compos A Appl Sci Manuf 2008;39:154–63. [7] Tong L. Bearing failure of composite bolted joints with non-uniform bolt-towasher clearance. Compos A Appl Sci Manuf 2000;31(6):609–15. [8] Tong L, Jain LK, Leong KH, Kelly D, Hertzberg I. Failure of transversely stitched RTM lap joints. Compos Sci Technol 1998;58(2):221–7. [9] Chang P, Mouritz AP, Cox BN. Tensile properties and failure mechanisms of zpinned composite lap joints. Compos Sci Technol 2006;66:2163–76. [10] Cartie DDR, Dell Anno G, Poulin E, Partridge IK. 3D reinforcement of stiffenerto-skin t-joints by z-pinning and tufting. Eng Fract Mech 2006;73:2532–40. [11] Ucsnik S, Scheerer M, Zaremba S, Pahr DH. Experimental investigation of a novel hybrid metal-composite joining technology. Compos A Appl Sci Manuf 2010;41:369–74. [12] Graham DP, Rezai A, Baker D, Smith PA, Watts JF. The development and scalability of a high strength, damage tolerant, hybrid joining scheme for composite–metal structures. Compos A Appl Sci Manuf 2014;64:11–24.
[13] Parkes PN, Butler R, Meyer J, de Oliveira A. Static strength of metal-composite joints with penetrative reinforcement. Compos Struct 2014;118:250–6. [14] Yang L, He XD, Mei L, Tong L, Wang RG, Li YB. Interfacial shear behavior of 3D composites reinforced with chemically connected CNTs. Compos A Appl Sci Manuf 2012;43(8):1410–8. [15] Tong L, Sun X, Tan P. Effect of long multi-walled carbon nanotubes on enhancing delamination toughness of laminated composites. J Compos Mater 2008;42(1):5–23. [16] Tong L, Steven GP. Analysis and design of structural bonded joints. Boston, USA: Kluwer Academic Publishers; 1999. [17] Wood MD, Sun X, Tong L, Kaltzos A, Rispler A, Mai Y-W. Effect of stitch distribution on mode I delamination toughness of stitched laminated composites – experimental results and simulation. Compos Sci Technol 2007;67(6):1058–72. [18] Islam MS, Tong L, Falzon PJ. Influence of metal surface preparation on its surface profile, contact angle, surface energy and adhesion with glass fibre prepreg. Int J Adhes Adhes 2014;51:32–41.