Investigation of failure modes and mechanical properties of hybrid joints with different interface patterns using digital image correlation

Xichang Wang1,2,3,a*, Joseph Ahn2,b,Junyi Lee2,c, Bamber R K Blackman2,d

1Science and Technology on Power Beam Processes Laboratory,

Beijing Aeronautical Manufacturing Technology Research Institute, Beijing, 100024, China

2Department of Mechanical Engineering, Imperial College London, London,SW7 2AZ, UK

3Huazhong University of Science and Technology, Wuhan, 430074,China

axichangw08@163.com,*corresponding author

bjoseph.ahn08@imperial.ac.uk,cjunyi.lee108@imperial.ac.uk,db.blackman@imperial.ac.uk

Keywords:Hybrid joining; Carbon fibre; Mechanical properties; Digital image correlation

Abstract:An advanced hybrid joining technology for joining metal and composite is introduced. Protrusions formed on the surface of the metal by electron beam are embedded into carbon fibre reinforced polymer layers and thus forms an integrated joints. The performance of a joint produced using such method was found to be better than traditional joints. In this paper, the properties of two different patterns of protrusions, including a linear pattern and a cylindrical pattern were studied by uniaxial tensile testing of double lap composite structures using digital image correlation. The distributions of strains in the composites tested varied and were found to be influenced by the shape of the protrusions which also resulted in different failure modes. The joints with a linear pattern failed between laminate layers, whereas the joints with a cylindrical pattern fractured at the interface of metal and composite. Furthermore, the tensile properties such as the ultimate tensile strength and elongation to failure of the joint with the linear pattern were around twice the value of the joint with the cylindrical pattern.Consequently, the performance of hybrid joints can be improved significantly by optimising the protrusion pattern.

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1. IntroductionComposites, especially resin based composites have many advantages including low density, high strength, and relatively low manufacturing costs. As such, they have been widely used in a variety of fields, including aerospace, automotive, shipping, and sports [1-3]. In order to maximize structural performance, several components are designed and manufactured with different materials, for example carbon fiber reinforced polymer (CFRP) and aluminium, to form an integrated component. In this aspect, the joining of dissimilar materials is essential, so that these components can be produced efficiently with no structural issues.Thus, joining metal and composite materials, like carbon fiber reinforced polymer (CFRP), or glass fiber reinforced polymer (GFRP) have been extensively researched in recent years. Apart from the traditional joining technologies [4-6], such as adhesive joining and mechanical joining (riveting and bolting), several new joining methods have been devised. An example of these new joining methods is Comeld[7] joining developed by the Welding Institute (TWI). This method is a novel hybrid joining technology that is based on electron beam Surfi-Sculpt process that effectively produces protrusions on metal surfaces. Comeld joints have been reported to withstand higher loads before failure compared to conventional joining methods [8, 9].Other research focused on the simulation of protrusion geometry effect on the properties of joints [10] and investigation of protrusions[11].The protrusion used in these joints can also be produced using other manufacturing methods, for example,cold-metal transfer(CMT), which is based on arc welding to weld the pins on the surface of stainless steel. Many experimental studies, including failure mode analysis [12], have been conducted for these types of joints. Additive manufacturing have also been investigated on advanced hybrid joining, based on 3D printing technology, by using selective laser melting or laser metal deposition [13].Although titanium alloys are known for having high strength and stiffness while having relatively low densities, which makes it ideal for aerospace and other lightweight applications, not many studies have been conducted for hybrid joints with titanium alloys instead of aluminium or steel. This may be a result of difficulty in manufacturing titanium alloys due to low thermal coefficient.Severalresearch results [9-10,12-13] showed pins or protrusions made by different methods would have significant effects on the properties and failure mode of joints. Additionally, different parameters of pins or protrusions will lead to different mechanical properties and failure modes even when the constituent materials are joint using the same technique.Therefore, this work aims to investigate effects of different patterns protrusions on Comeld hybrid joints and to show damage development by using digital image correlation (DIC) during tensile testing. The results from this study will serve as aguide for the design or manufacture of hybrid joints in practical engineering application. This study will also contribute to the overall literature on the subject of hybrid joints with the metal part of the joint made from titanium alloy, which is uncommon in the literature.2. Materials and MethodsJoint manufacture

The joint will be designed as a single step and double lap structure, as shown in Fig.1. The metal part onthe left side is made from Ti-6Al-4V, the composite part in the right side was made from carbon fibre reinforced polymer (CFRP). The CFRP used in this study was HexPly®8552/33%/UD268/IM7 (12K) unidirectional carbon fibre prepreg plies manufactured by Hexcel. The overlapped area was 33x30 mm. The composite part contains three laminate, upper (3.5mm), middle (4mm), and bottom (3.5mm) sections. The upper and bottom sections consist of 14 layers with the same layup [-45/45/0/45/90/-45/0]sym, while the middle section was designed to have16 layers with the layup [0/-45/45/90/45/0/-45/0]symaccording to the cured ply thickness (0.251 mm) of the raw prepreg.

Fig.1 Dimension of Comeld hybrid joint design

Fig.2 shows the different patterns protrusions on the surface after the electron beam Surfi-Sculpt process. The specimen at the top of the figure has a cylindrical pattern; the processing parameters areas follows: an accelerating voltage of 150 kV, a beam current of 3.5 mA, a scanning frequency of 1500 Hz, a processing time of 11 s, a scanning waveform with a linear array (8×5). The dimensions of the protrusions geometry were as shown in Fig.3a: height, h=1.2mm; diameter, d=0.9 mm; distribution density, σ = 4 pins/cm2; Δx=5.5 mm (the nearest distance between two protrusions along x direction); Δy=3.2 mm (the nearest distance between two protrusions along y direction).The specimen at the bottom of the figure has a linear pattern; the processing parameters areas follows: an accelerating voltage of 150 kV, a beam current of 3.5 mA, a scanning frequency of 1500 Hz, a processing time of 11 s, a scanning waveform with a linear array (23×8). The dimensions of the protrusions geometry were shown in Fig.3b: height, h=1.2mm; angle, θL =90o, θR =45o;top width, w=0.4 mm, thickness, δ=0.4mm;distribution density, σ =18 pins/cm2;Δx=3.6 mm; Δy=1.2 mm.

Fig.2Cylindrical and linear patterns protrusions by electron beam Surfi-Sculpt

Fig.3Profiles of different patterns protrusions(a) Cylindrical protrusion (b) Linear protrusion

The curing of the joint was performed in an autoclave at a pressure of 7 bar and a temperature of 180°C with control vacuum of 0.2 bar.Fig.4 shows the joints cured in the autoclave after laying up and joining.

Fig.4A cross-sectional view of the interface between Ti-6Al-4V and GFRP of the Comeld joint after curing

Experimental Setup and Digital Image Correlation (DIC)Tensile tests were conducted in order to investigate the effects of the protrusion geometry on the failure mechanisms of the joint. The samples were tested on a 250kN tensile testing machine under quasi-static conditions. The cross-head displacement (extension) rate of 1 mm/min were applied to all samples and the tests were conducted at ambient temperature and humidity (23°C,50%RH).The real time full field surface strain distributions of the specimens were obtained using an ARAMISv6.3.0 Digital Image Correlation (DIC) system [14]. Digital Image Correlation(DIC) is a method of measuring the displacement and deformation of the object by comparing recorded images with a stochastic color spray pattern [14].This method offers several advantages, such as non-contact, capable of providing strain information at high temporal and local resolution, and it can also be used in high temperature environments[14]. Therefore, DIC is widely used in the different fields of engineering, such as bio-medical engineering[15,16], structural dynamics[17] and many others, for non-contact strain or displacement measurements. Some studies[18]have also recommended the method to be used in future research of composite joints as it allows the measurement of the strain distribution for detailed study of the failure mechanisms. However, despite these benefits there are only a small amount of studies, like [19], that have been conducted on hybrid joints with the DIC method, especially for Comeld joints. The use of DIC will facilitate the investigation as the failure mechanisms for Comeld joints can be deduced for the surface strain distributions. In this paper, digital images recorded at a set time interval of one image per second during tensile tests were used with the ARAMIS DIC system to calculate the deformation of the specimens.

The required stochastic speckles on the specimen surfaces for the DIC strain measurement [20] were produced by using black and white spray-paints as seen in Fig.5.The average speckle diameter (aerosol spray) was approximately 0.1 mm. Additionally, the facet size and facet step was set up according to the configuration with the average speckle diameter and the resolution of the image in order to ensure that the results are accurate.

Fig. 5The surface of the specimen viewed from the side after the application of a speckle pattern

3. Results and discussion

Uniaxial tensile testing of hybrid joints

Two sets of Comeld joint samples, each set with 5 samples, were tested in this study, a set with the linear pattern Comeld joints and another set with the cylindrical pattern Comeld joints. The results from the tensile tests are summarized in Tables 1 and 2 for the linear and cylindrical pattern of Comeld joints respectively. The values of tensile stresses,, and shear stresses,, in the tables were calculated using Eq. 1 and Eq. 2respectively.Eq 1.

Eq 2.

Where is the force, is the cross sectional area ( mm 11 mm), and is the overlapped area

Table 1Results of tensile testing of linear pattern Comeld joints (Ti6Al4V/CFRP)

No.

F (kN)

Displacement (mm)

σ (MPa)

τ (MPa)

1

65.97

4.73

200.0

33.3

2

73.10

5.60

221.5

36.9

3

75.12

5.87

227.6

37.9

4

72.91

5.49

220.9

36.8

5

76.85

5.10

232.9

38.8

Av

72.79

5.36

220.6

36.8

STD

4.14

0.45

12.54

2.09

Table 2 Results of tensile testing of cylindrical pattern Comeld joints (Ti6Al4V/CFRP)

No.

F (kN)

Displacement (mm)

σ (MPa)

τ (MPa)

1

25.86

2.14

78.36

13.06

2

33.09

2.44

100.27

16.71

3

36.22

3.36

109.76

18.29

4

28.77

2.31

87.18

14.53

5

23.80

2.03

72.12

12.02

Av

29.55

2.46

89.54

14.92

STD

5.11

0.53

15.47

2.58

The loading histories of the Comeld joints up to failure for the two patterns are shown in Fig.6. For both patterns, the force increases linearly with increasing displacements until failure, as seen in Fig.6.Additionally, the behaviour of the specimens was similar to the behaviour of the bulk CRFP composite. However, as seen in Fig.6, the overall strength and stiffness of the specimen with the cylindrical pattern is lower than those for the specimen with the linear pattern. For example, the specimen with the cylindrical pattern failed at 35 kN, which is approximately half of that for the specimen with the linear patter which failed at 75 kN. By comparing the specimens with the linear and cylindrical pattern after failure, shown in Fig.7 (a) and (b) respectively, the large differences in failure strengths can be attributed to the difference in failure mechanisms. In Fig.7 (a), the failure of the specimens with the linear pattern occurred within the CFRP laminates as some plies remained on the overlapped region, in which the protrusions are connected to the composite. Conversely, in Fig.7 (b), the specimens with the cylindrical protrusions appears to have failed due to the pull-out of the protrusions from the specimen, as the protrusions are still intact and no plies are connected at the overlapped region. This result suggests that the stresses required to cause the pull-out of the protrusions in the specimens with the cylindrical protrusions in this study is significantly lower than that to cause the failure of the CFRP laminates, which is likely to be a result that the geometries of the cylindrical protrusions were not being optimised to provide maximum strength. Further discussions of the failure mechanism will be presented in the following section.

Fig.6Results of tensile testing of Comeld joints (Ti6Al4V/CFRP) with two difference patterns

Fig.7FracturedComeld specimens after tensile testing (a) Linear protrusions (b) Cylindrical protrusions

Analysis of DIC results

Strain distribution on the front surface the jointsFig.8 (a) and (b) shows the strain fields at the joints on the front plane of the specimens with the linear protrusions and cylindrical protrusions respectively prior to failure while Fig.9 (a) and (b) shows the strain values at selected locations in Fig.8 for both specimens. As seen in Fig.8 and Fig.9, at the point of failure, the strains at the interface between the metal and CFRP are the highest while the other regions experienced negligible amounts of strains. This result suggests that the failure initiates at the interface between the metal and CFRP component as a result of the failure of the adhesive bonds. Additionally, for both types of specimens, the strains along the interface are almost constant. The strains for the specimens with the cylindrical protrusions are also lower at approximately 5 % compared to that for the specimens with the linear protrusions which are at approximately 11 %. This result is expected as the specimens with the cylindrical protrusions fails at lower levels of forces and displacements comparedto the specimens with the linear protrusions.

Fig.8 Comparison of the strain distribution at 99% strain to failure (a) specimen with linear protrusions (b) specimen with cylindrical protrusions.

Fig.9 Axial strains at lines shown in Fig 8 for (a) specimen with linear protrusions (b) specimen with cylindrical protrusions.

Distribution of transverse strain on the side surface of the joints

In addition to the axial strains at the front face, the transverse strains at the side face of the specimens for both the specimens with linear and cylindrical protrusions are shown in Fig.10 and Fig.11 respectively. By observing a section across the side face (line 4 and 4') for both types of specimens shown in Fig.10 (b) and Fig.11 (b), the deformations appears to be large at the faces between the two faces of the metal with the protrusions and the CFRP as denoted by the two peaks. Since the material is symmetrical, the peaks in Fig.10 (b) is also symmetrical as expected while the in Fig.11 (b) the peak on the left appear to be higher. This is likely to be a manufacturing defect that causes some stress concentration on the left face.

By comparing the strain field in Fig.10 (a) and Fig.11 (a), the strain distributions for both types of specimens appear to be quite different. In Fig.10 (a), the largest strains at the onset of failure were found to be not at the interface of the metal and CFRP but within the CFRP laminates. This result suggests that for the specimens with the linear protrusions, the failure of the CFRP laminates also contribute to the overall failure of the joint. This result is consistent with the observations made in Fig.7 (a) and the deduction that the failure of the specimens with linear protrusions occurs within the CFRP laminates. On the other hand, for the specimens with the linear protrusions, it can be seen that the strains are concentrated along the interface between the metal and the CFRP as seen in Fig.11 (a), this result suggests that the mechanism that causes the failure in the specimens with the cylindrical protrusions occurs mainly at the interface between the metal and CFRP interface. Again, this is consistent with the observations made in Fig.7 (b) in which the joints appear to be caused by the pull-out of the protrusions and little damage was found on the CFRP laminates.

Fig.10 (a) Computed transverse strain full field distribution at 99% of the failure time on side surface for the specimen with the linear protrusions. (b) transverse strain distribution on the section line 4 at different strains to failure.

Fig.11 (a) Computed transverse strain full field distribution at 99% of the failure time on side surface for the specimen with the cylindrical protrusions. (b) transverse strain distribution on the section line 4' at different strains to failure.

Side surfaceresults of the joints (Axial strain)

Lastly, the axial strains at the side face have also been calculated for both sets of specimens with linear and cylindrical protrusions and these are shown in Fig.12 and Fig.13 respectively. As seen in the strain fields and the axial strains of selected sections in the figures, both the specimens with the linear and cylindrical protrusions showed similar trends, in which the maximum strains are largest at the metal and CFRP interface. As seen in the figures, Section 5 (5') and Section 7 (7') in Fig.12 and Fig.13 (b) and (d) for both types of specimens are almost symmetrical with the largest strains observed at the top section (Section 6 (6')) in Fig.12 and Fig.13 (c). This result suggests that the first failure occurred at the region in line 6 (6') for both types of specimens.

Similar to the observations made for the strains at the front face, the axial strains at the sides for the specimens with the cylindrical protrusions are approximately half of that for the strains for the specimens with the linear protrusions. These results are consistent with earlier observations and that the specimens with cylindrical protrusions failed at lower stress levels.

Fig.12 (a)Computed axial strain (ε22) full field distribution at 99% of the failure time on side surface for the specimen with the linear protrusions. (b) axial strain (ε22) distribution on the section line 5. (c) axial strain (ε22) distribution on the section line 6. (d) axial strain (ε22) distribution on the section line 7.

Fig.13 (a) Computed axial strain (ε22) full field distribution at 99% of the failure time on side surface for the specimen with the cylindrical protrusions. (b) axial strain (ε22) distribution on the section line 5'. (c) transverse strain (ε22) distribution on the section line 6'. (d) axial strain (ε22) distribution on the section line 7'

Three dimensional results from specimensFig.14 and Fig.15 show the computed major strains fields at 99 % failure strain for the specimens with the linear protrusions and cylindrical protrusions respectively. As seen in Fig.14 (b), there are peaks of strains at regular intervals suggesting that stress concentrations occur near the protrusions, which was about 1.5 mm far from the edge of the side surface. Although the specimens with cylindrical protrusions in Fig.15 (b) show similar behaviour, the peaks are less significant for farther distance (roughly 2.5 mm) from the edge. On the other hand, the dimension of cylindrical protrusion was bigger than that of linear protrusion, as shown in Fig.3.Furthermore, the overall strains for the specimens with the linear protrusions in Fig.14 (b) are higher compared to the overall strains for the specimens with cylindrical protrusions in Fig.15 (b). This observation may be attributed to the fact that the specimens with the cylindrical protrusions failed at a lower force and displacement levels. However, at onset of failure, the difference between the strains at the interfaces with the protrusions and the interfaces without the protrusions for the linear specimens are higher at about 20 % compared to 5 % of the cylindrical specimens. The result suggests that the strains for the cylindrical protrusions are distributed in a more homogenous manner.

Fig.14 (a) Computed major strain full field distribution at 99% of the failure time on side surface of the specimen with the linear protrusions. (b) 3D plot of local major strain field distribution at 99% of the failure time on the side surface of the specimen.

Fig.15 (a) Computed major strain full field distribution at 99% of the failure time on side surface of the specimen with the cylindrical protrusions. (b) 3D plot of local major strain field distribution at 99% of the failure time on the side surface of the specimen.Failure mode analysisThe strains fields presented in the previous sections and the observations from the specimens after failure in Fig.16 to Fig.18 can be used to deduce the failure mechanisms. Firstly, it is obvious that the failure mechanisms for the specimens with the linear protrusions and the cylindrical protrusions experiences different failure mechanisms. For the specimens with the linear protrusions, the failure occurs in a two stage process. In the first stage, the adhesives bonding the metal to the CFRP at the top interfaces (Regions C in Fig.16) will fail first followed by a similar debonding process at the bottom interfaces (Regions A and B). The weaker bond strength of the adhesives at these regions is the key contributors to the initiation of failure at these zones and this conclusion is supported by the high strains at these regions shown in Fig.12. The second stage of the failure starts when all three interfaces have failed. Unlike the cases described in [13, 15], in which the specimens failed via the fracture of the protrusions, the protrusions in the specimen for these study has been optimized to prevent fracturing of the pins. Therefore, in this study the failure occurs between the piles of the composite materials as seen in Fig.16 and Fig.17. The damage to the composite consists of both shear and tensile modes. Here, the plies remain bonded to the protrusions and the amount of layers were found to be 6 plies, where each layer is about 0.251 mm thick based on the observation of the surface layer direction (-45 ° on both sides). Hence, the maximum load is limited to the damage that can be taken by the remaining layers and shear between the laminates.

Fig.16Failure of the Comeld specimen with linear patterned protrusions

Fig.17 Delamination failure on the interface between composite layers: (a) failure behaviour on the metal part. (b)failure behavior on the CFRP part.The failure mechanism for the specimens with the cylindrical protrusions is also a two stage process. The first stage of the failure mechanism of the specimens with the cylindrical protrusions is similar to that of the specimens with the linear protrusions, where the bonds at the three interfaces failed. The second stage of the failure of the specimens with the cylindrical protrusions starts when all three interfaces failed. However, unlike the specimens with the linear protrusions, these specimens fail by the pulling out of the pins from the CFRP component, as seen in Fig.18, since the pins are sufficiently strong that they do not fracture, instead of the failure of the CFRP plies. Additionally, during the pull-out process a layer of the CFRP laminates will also experience de-bonding in which a crack is observed in Fig.18 (a) in the CFRP laminate. Since the strength of the bonds to hold the pins are lower than the strength of the CRFP laminates, the overall strength and hence the force at failure for the specimen with the cylindrical protrusions are lower than that for the specimens with the linear protrusions. This result highlights the importance of optimising the pins for the cylindrical protrusions to maximise its failure stresses.

Fig.18Failure of the Comeld specimen with cylindrical patterned protrusions (a) side surface. (b) front surface.

It was determined that the mode of failure in these hybrid joints was influenced by the pin geometry and pattern. The use of cylindrical pins resulted in mainly interfacial failure at the adhesive and adherend interface, and shear failure of the composite matrix on the surface of the adherend. On the other hand, the use of linear pins resulted in mainly inter-laminar failure in the adherend laminate and failure of the surface lamina due to intra ply matrix cracks. The wider pin section at the root of linear pins improved the joint performance by bridging the incipient crack more effectively, whereas, the small pin radius of cylindrical pins decreased the bridging performance of the pins and as a result, debonding of metal-composite interfaces became the critical damage mode. The larger pin density used for linear pins enhanced the global pin bridging traction and increased the joint load carrying capability so the adhesive-adherend interface was able to remain intact. The mechanical performance of the hybrid joints can therefore be optimized by a careful balance between protrusion shape, height, diameter angle and density. It is essential to keep the pin radius large enough to provide a sufficient strength and stiffness so that shear failure of the pins does not occur and also to prevent it from deforming when the uncured composite laminate gets inserted. The effect of electron beam Surfi-Sculpt processing parameters on the height of Ti-6Al-4V protrusions was investigated by Xichang et al. and the optimum height was found to be around 1.6 mm[21]. Increasing the pin density will reduce the debonding behavior of the hybrid joint by transferring the bridging load of the pin more effectively and the shear strain at the tip of the protrusions.

4. Conclusions

Double lap joint specimens were manufactured by joining aerospace titanium alloy and CFRP using an advanced Comeld hybrid joining technique for two different types of pin configurations, linear pattern and cylindrical pattern. Uniaxial tensile tests were perform on the specimens until failure and the deformations of the specimens were measured using full field DIC to determine the failure mechanisms for the specimens.

In this study, the failure strength of the specimens with the linear protrusions was found to be approximately twice the value of that for the cylindrical pattern. In order to determine the cause of this observation and the failure mechanism, the DIC strain distributions were studied. The DIC strain distributions for both types of specimens were similar. Significant stress concentrations at the metal-CFRP interface were observed for both types of specimens. Additionally, the DIC measurements of the strain distribution at the side of the specimens revealed that the pins will result in stress concentrations around the regions with the pins as distinct peaks can be found at these regions. However, the specimens with the cylindrical protrusions appear to have a more homogenous distribution of the stresses compared to that for the linear protrusions for bigger size, farther distance from the edge of surface and lower tensile force.

The measured strain distributions and observations made on the final state of the failed specimens were used to deduce the failure mechanism for the specimens. Both types of specimens experienced two stage failure mechanisms with the failure mechanism for the first stage being identical. In the first stage, the failure occurs at the tensile metal-CFRP interfaces as the adhesives bonding the metal and CFRP portions without the pins will be the first to fail as the strength of the adhesive bonds are the weakest. The second stage of the damage will initiate when all three interfaces failed. The second stage of the damage mechanism differs for both the specimens with the linear and cylindrical protrusions. Since the pins are sufficiently strong for both cases, no damage in the pins was observed in this study. For the specimens with linear protrusions, the failure occurs within the CFRP laminates in which the debonding occurs at the plies right above the pins. On the other hand, for the specimens with the cylindrical protrusions, the failure was caused by the pins being pulled out from the CFRP laminate. Since the strength of the adhesive bonds holding the pins to the CFRP laminates are weaker than the CFRP laminates, the overall strength of the specimen, which are limited by the strength that causes failure, for the specimens with the cylindrical protrusions are lower than that for the specimens with the linear protrusions. Therefore, the pin designs for the cylindrical protrusions should be optimised so that the failure occurs when the CFRP laminates fail in order to maximise the strength of the specimens.

Acknowledgement

The strong support from Aviation Industry Corporation of China (AVIC) for this funded research is much appreciated. The research was performed at the AVIC Centre for Structural Design and Manufacture at Imperial College London. The authors acknowledge the financial support from National Natural Science Foundation of China (No.50975268), and thank Prof John P Dear of Imperial College London for his guidance on DIC.

References

[1] http://www.airbus.com.

[2] Hart-Smith LJ. Adhesive Bonded Double-lap Joints. NASA report CR112235;1973.

[3] Moutrille M P, Derrien K, Baptiste D, Balandraud X, Grédiac M. Through-thickness strain field measurement in a composite/aluminum adhesive joint. Composites Part A 2009;40:985-996.

[4] Camanho P P, Matthews F L. Stress analysis and strength prediction of mechanically fastened joints in FRP: a review. Composites Part A 1997;28A:529-547.

[5] Kweon J, Jung J, Kim T, Choi J, Kim D. Failure of carbon composite-to-aluminum joints with combined mechanical fastening and adhesive bonding. Compos Struct 2006;75:192–198.[6] Dano M, Gendron G, Picard A. Stress and failure analysis of mechanically fastened joints in composite laminates. Compos Struct 2000;50:287-296.

[7] Dance B G I, Kellar E J C. Workpiece structure modification. International patent publication number WO2004028731 A1.

[8] Buxton A L, Dance B G I. Surfi-Sculpt - Revolutionary surface processing with an electron beam. Proc. ISEC.Minnesota,August,2005.

[9] Wang X C, Guo E M, Gong S L, Dance B G I. Study of electron beam Surfi-sculpt during composite materials joining. Rare Metal Mat Eng 2011;40(S4):292-296.

[10] Tu W, Wen P H, Hogg P J, Guild F J. Optimisation of the protrusion geometry in Comeld™ joints. Compos Sci Technol 2011;71:868-876.

[11] Wang X C, Guo E M, Gong S L, Li B. Realization and experimental analysis of electron beam Surfi-Sculpt on Ti-6Al-4V alloy. Rare Metal Mat Eng 2014;43(4):0819-0822.

[12] Ucsnik S, Scheerer M, Zaremba S, Pahr D H. Experimental investigation of a novel hybrid metal-composite joining technology. Composites Part A 2010;41:369-374.

[13] Parkes P N, Butler R, Meyer J, Oliveira A de. Static strength of metal-composite joints with penetrative reinforcement. Compos Struct 2014;118:250-256.

[14]GOM. ARAMIS, User Manual - Software 2009;6.1:129.

[15]Zhang D, Arola DD. Applications of digital image correlation to biological tissues. J Biomed Opt 2004;9:691–9.

[16]Sutradhar A, Park J, Carrau D, Miller MJ. Experimental validation of 3D printed patient-specific implants using digital image correlation and finite element analysis. Comput Biol Med 2014;52:8–17.

[17]Helfrick M, Niezrecki C. 3D digital image correlation methods for full-field vibration measurement. Mech Syst. 2011;25:917–27.

[18]Löbel T, Kolesnikov B, Scheffler S, Stahl a., Hühne C. Enhanced tensile strength of composite joints by using staple-like pins: Working principles and experimental validation. Compos Struct 2013;106:453–60.

[19]Hashim S, Berggreen C, Tsouvalis N, McGeorge D, Chirica I, Moore P, et al. Fabrication, testing and analysis of steel/composite DLS adhesive joints 2011:37–41.

[20]Cheng P, Sutton M a., Schreier H, McNeill SR. Full-field Speckle Pattern Image Correlation with B-Spline Deformation Function. Exp Mech 2002;42:344–52.

[21] Wang, X., Ahn, J., Bai, Q., Lu, W., & Lin, J. (2015). Effect of forming parameters on electron beam Surfi-Sculpt protrusion for Ti–6Al–4V. Materials & Design, 76, 202-206.