- Email: [email protected]
Using thin-plies to improve the damage resistance and tolerance of aeronautical CFRP composites
Composites: Part A 86 (2016) 31–38
Contents lists available at ScienceDirect
Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Using thin-plies to improve the damage resistance and tolerance of aeronautical CFRP composites T.A. Sebaey ⇑,1, E. Mahdi Mechanical and Industrial Engineering Department, College of Engineering, Qatar University, 2713 Doha, Qatar
a r t i c l e
i n f o
Article history: Received 17 October 2015 Received in revised form 19 March 2016 Accepted 24 March 2016 Available online 1 April 2016 Keywords: A. Hybrid B. Compression after impact B. Damage tolerance B. Impact behavior Thin-plies
a b s t r a c t Thin-ply composites are currently receiving specific attention from researchers due to their capabilities to delay matrix cracking. In this paper, the aim is to design a hybrid laminate that contains both thin- and normal plies. The objective is to improve the tolerance of normal plies by adding thin-plies to the composite in different configurations. Two alternatives were designed, tested, and compared to the specimens made of traditional plies. Impact and compression after impact tests were conducted on each configuration at different impact energies. After being impacted, the specimens were c-scanned to define the delamination pattern. Results showed that surrounding each normal ply with two thin-plies improved the delamination threshold by 15% as compared to the specimens made all of normal plies. Under compression, 15% improvements in the compression after impact strength were obtained. By using thin-plies, the size of each individual delamination was reduced, resulting in small threads instead of peanut delaminations. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Laminated composites are currently being used in several structural and non-structural applications including aeronautics, automotive, marine, and piping. The main advantages of laminated composite materials are the high specific strength and stiffness as well as the dimensional stability in comparison to the conventional metallic structures. Conventional, well-known composites are made of several layers of fiber reinforced plastics of a nominal ply thickness greater than 0.1 mm [1]. Reducing the ply thickness down to this limit is usually referred to as thin-plies. Few efforts were found in the literature to reduce the ply thickness beyond this limit [2,3]. In 2007, Tsai and Kawabe [4] published a patent that addressed a new tow spreading technology that enables the production of plies of thickness down to 0.02 mm thickness. The report also addressed some of the potential improvements in the structural response by using such laminates. The interest in thinplies can be justified by the possibility of improving the ply in situ strength as the ply thickness is reduced [5]. Detailed potential advantages of using thin-plies in composite design are summarized by Arteiro et al. [6]. Shin et al. [7] presented
⇑ Corresponding author. E-mail addresses: [email protected], [email protected] (T.A. Sebaey). On leave from Mechanical Design and Production Dept., Faculty of Engineering, Zagazig University, 44519 Zagazig, Sharkia, Egypt. 1
http://dx.doi.org/10.1016/j.compositesa.2016.03.027 1359-835X/Ó 2016 Elsevier Ltd. All rights reserved.
a study on using thin-plies (0.04 mm ply thickness) in comparison to traditional plies of 0.2 mm thickness. The comparison was made based on the results of the unnotched tensile, open-hole tensile, and impact as well as compression after impact tests. Tensile tests were conducted both statically and under fatigue. Two different types of specimens were considered. Both laminates composed of the same amount of fiber in each direction. For the first stacking sequence ‘‘THIN” the authors did not allow clustering resulting in a nominal ply thickness of 0.04 mm whereas, for the second one ‘‘THICK”, each 5 layers of the same orientation angle were clustered together resulting in a nominal layer thickness of 0.2 mm. The results of the tensile tests showed an improvement of 10% in both the strength and stiffness. In addition, less sensitivity to the number of cycles was recorded by using thin-plies. In open hole tensile tests, the advantage was given to the thick-plies with 10% higher strength and stiffness in comparison to the thin-plies. The authors justified this response by the stress relaxation near the hole edge after the initial failure. As a general observation in all the tests, the fracture is more brittle in comparison to the traditional ones. These improvements were justified by the stress relaxation and the energy release rate at the crack tip through meso-scale FE modeling by Saito et al. [8]. In addition, the capability of thin-plies to resist the propagation of large cracks does play a role in this improvement [9]. This higher capability helped also to push the damage away of the thin-plies in the experiments designed by Guillamet et al. [10] to monitor the damage at the free edge under
32
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
tension. Under tensile loading, Fuller and Wisnom [11,12] presented an experimental and analytical study to investigate the behavior of ½h5 s ; h = 15, 20, 25, 30, and 45 thin-plies. In addition to the good agreement between the experimental and the analytical predictions, for all angles, delaminations were suppressed, allowing considerable pseudo-ductile strains to develop. Other tensile tests were preformed by Moon et al. [13] under harsh low earth orbit (LEO) conditions. The authors used two different ply thicknesses of 0.125 mm and 0.06 mm. The authors measured the tensile strength and stiffness of both laminates before and after the LEO aging. The authors concluded that using the thin-plies resulted in improvements of 8.9% and 12.1% at ambient and LEO conditions, respectively. On the contrary, reduction of the stiffness of 8.6% and 12.8% was recorded by using the thin-plies at both environmental conditions, respectively. It is worth noting that the fiber volume fraction of the two specimens was different (about 10% difference) which makes it very difficult to draw a clear conclusion from these results. The notch sensitivity of the thinplies were assessed (experimentally and using the finite fracture mechanics approach) by Arteiro et al. [6,14] for two different stacking sequences and compared with results from the literature for normal plies. The results showed that the notch sensitivity of laminates made of thin-plies can be equivalent to that made of thick plies. It is worth remarking that the brittleness is a characteristic of the damage of laminates made of thin plies as it was noted by others. Based on the fracture mechanics approach, Czél and Wisnom [15] suggested a hybridisation system between glass and carbon fibers layers that introduced some ductile characteristics to the damage of the thin-ply composites. Yokozeki et al. [16] compared the compression properties of laminates made of thin-plies of 0.07 mm ply thickness with the ones made of normal ply thickness of 0.14 mm. The authors performed compression, open-hole compression, and compression after impact tests. The results showed improvements of 16%, 9% and 8% for the three mentioned compression tests, respectively. The authors reported a significant reduction in the amount of damage as measured by the X-ray and the acoustic emission techniques. Under out-of-plane loading, Yokozeki et al. [17] compared the results of the two test specimens using indentation tests. The results showed that in thin-plies damage initiated by accumulation of matrix cracking and localized delaminations and then sudden fiber breakage. For standard plies (0.14 mm ply thickness), more visible delaminations appeared after matrix cracking and then fiber breakage. The authors concluded that thin-ply laminates have high damage resistance against matrix cracking and delaminations near the back surface due to the absent of back face delaminations. A limited number of studies addressed the impact of thin-ply laminates, and if found, it is a part of the study with few impact energies and limited discussions on the results. Yokozeki et al. [16] compared the impact properties of laminates made of thinplies of 0.07 mm ply thickness with the ones made of normal ply thickness of 0.14 mm at 6.7 J impact energy. The results did not show any significant effect on the damage area. The reason could be the lower value of the impact energy which, most probably, is lower than the delamination threshold [18,19]. For three different ply thicknesses, Amacher et al. [20] compared the impact response. The results showed that reducing the ply thickness down to 30 lm resulted in a brittle failure with extensive translaminar cracks. Moreover, the optimum ply thickness was found to be 100 lm which is almost the thickness of the traditional plies. Based on the results and the conclusions summarized, it can be inferred that using thin-plies results in brittle damage which usually leads to catastrophic undesirable failure with some improved strength. However, for low velocity impacts, there were no
recorded improvements in any of the parameters used to assess the damage tolerance in the literature. In the current paper, the woven thin-plies were used inside the laminate with the traditional plies in two different ways. The idea is to improve the damage tolerance and the impact resistance as well as avoid the brittle failure mode obtained by Amacher et al. [20]. Low velocity impacts were conducted at different energy levels and impact velocities. 2. Materials and manufacturing Spread tow fabrics/epoxy composites were supplied by Oxeon. Two grades of woven carbon fibers fabrics were used: TeXtreme 80 PW UTS50S of 80 gsm and TeXtreme 320 PW UTS50S of 320 gsm. One TeXtreme +45/45 fabric layer is equivalent to four layers of unidirectional. The nominal ply thickness of the traditional (320 gsm) and the thin-plies (80 gsm) were 85 and 330 lm, respectively. The MTFA500 (DF044) epoxy resin (delivered by SHD composites) was used as a matrix to manufacture the specimen in an autoclave based process. Considering the previous results on low velocity impact of thinplies and to avoid brittle failure, hybrid specimens were manufactured from both traditional and thin-plies and the damage resistance was compared with specimens made of all traditional plies. Three stacking sequences were considered. The first group of specimens was the baseline or the reference laminate. The reference laminate consisted of 12 layers of TeXtreme 320 PW. The fiber orientations of the different layers are shown in Fig. 1. It was considered, during the design stage to have the same amount of fiber in 0°, 90°, and 45°. The total laminate thickness was 3:980:038 mm. In the first alternative (A1), the laminate was designed to have traditional plies at the specimen surfaces and a block of thin-plies at the mid-plane. The idea of this design was initiated by Sebaey et al. [21]. In that paper, the authors supposed that consuming the impact energy, by different damage mechanisms, close to outer surfaces can leave a cluster of the undamage layers at the middle. This cluster could withstand more compression loading after impacts, and consequently, the damage tolerance is improved. To push the damage outside, the core of the specimen was designed by using the thin-plies technology. To have a fair comparison, the amount of fiber in each direction is designed to be equal to that of the laminate BL. The total thickness of the A1 laminate is 4:100:027 mm. For the second alternative (A2 laminate), the effect of the neighboring plies on the strength of each layer was taken into account. It is well-known that the crack initiation/propagation is highly affected by the neighboring plies (i.e., the higher the strength and stiffness of the neighboring plies, the higher the damage resistance of the ply under consideration) [8]. For this reason, in our design (A2), each traditional ply is surrounded by two thin-plies to improve its damage resistance. The total thickness of the A2 laminate is 4:110:033 mm. With those two designs, the user can benefit from the higher strength of the thin-plies and avoid the brittle nature associated with the laminates made all of thin-plies under impact loading. The three plates were manufactured with 700 700 mm. Then, the margins (40 mm all around) were cut. The fiber volume fraction was measured by the ignition test standard, ASTM D258411. The values of the fiber volume fraction were 50:410:52 %, 49:091:03 % and 50:010:28 % for BL, A1 and A2, respectively. 3. Impact and CAI tests After being cured, the plates were scanned using c-scan, and the positions with defects were eliminated. The plates were then cut using a computer controlled diamond wheel into low velocity
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 1. Schematic of the cross-section properties of the three stacking sequences. BL refers to the reference laminate whereas, A1 and A2 represent the alternatives 1 and 2, respectively.
impact specimens. The nominal in-plane dimensions of the impact specimen are 100 150 mm as per the test standards [22–24]. The test setup of the impact events is shown in Fig. 2. The two halves of the clamping system, after being collected together with the pneumatic clamp, provided an unclamped area of 75 mm diameter. An impactor of 16 mm diameter was impacting the specimens at different impact velocities [22]. The impactor mass ðM i Þ was chosen to be constant at 8.15 kg. The two side location pins were added to the standard IMATEK test jig provided with the machine in order to ensure central impacts. The impact energies ðEi Þ and the correpffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sponding velocity ðV 0 ¼ 2Ei =M i Þ and height ðHi ¼ Ei =ðM i gÞÞ values are shown in Table 1. After being impacted, the specimens were taken to the c-scan to assess the projected delamination area associated with each impact energy level. The Compression After Impact (CAI) tests were performed using a Hydraulic test machine and a 250 kN load-cell according to the guidelines introduced in the AITM1-0010 [22] and the ASTM D7137-07 [24]. Test setup is introduced in Fig. 3 test speed was set as 1 mm/min. The compression after impact strength ðrCAI Þ was calculated based on the ultimate load ðP max Þ and the specimen cross-section dimensions (rCAI ¼ Pmax =ðb hÞ, where b and h were the specimen width and thickness, respectively). In addition to the specimens summarized in Table 1, two non-impacted specimens of each stacking sequence were compressed up to failure to determine the non-impacted strength ðr0 Þ. The compression test setup of the non-impacted specimen is exactly as the compression after impact test. The residual strength can be calculated as the ratio between the CAI strength and the non-impacted specimen strength.
33
A summary of the peak loads and the delamination threshold loads are presented in Fig. 5. The delamination threshold can be defined as the load level at which the first unstable delamination appears. On the load–displacement curve, this load can be considered as the first sharp decrease in the stiffness [28]. The delamination threshold is shown on the load-time plot as a horizontal line, Fig. 4. For the impact event in Fig. 4(a), the maximum load is lower than the delamination threshold and, for that reason, the horizontal line does not appear. For the maximum load, the load is linearly proportional to the impact energy up to a certain limit (between 30 J and 45 J impact energies). After that limit, the maximum load seems to be constant. This response was also experienced elsewhere [29] for multidirectional laminates. The reason of this behavior could be justified by the initiations of fiber breakage at this energy limit. It is worth remarking that the current TeXtreme fabrics (either with or without the thin-plies) provide higher fiber breakage limit in comparison to the normal laminates made of UD fibers [21,25,29–31] with similar laminate thickness. The comparison between the laminates BL, A1, and A2 does not show any advantage to any of the laminates for up to 15 J impact energy. For impact energies higher than 15 J, the maximum load is usually higher in case of BL, as compared to A1 and A2. The difference between the three laminates can be seen when comparing the delamination threshold load. As it is mainly used to delay the matrix cracking and consequently the other damage mechanisms, the laminates that contain thin-plies have a higher delamination threshold than the ones that do not have thin-plies. The delamination threshold of each configuration shows constant value with respect to the impact energy values. This phenomenon was noted for other tested in the literature [18,21,25]. The values of the delamination threshold load are 4:790:12 , 5:560:17 , and 5:480:28 for the BL, A1, and A2 laminates, respectively. The comfor A1 and A2 does not show parison between the values of F Thr d any significant difference. This result could be justified by the higher number of ply interfaces resulted from using thin plies (to obtain the same thickness, more plies were used and hence, the number of interfaces increased). With higher number of interfaces, the energy is expected to be absorbed in higher number of narrow delaminations, as compared to the traditional plies. This delayed the appearance of the first unstable delamination and, consequently delayed the delamination threshold. The value of the impact energy Ei can be divided into two main components. The first component is energy required to initiate and propagate the different damage mechanisms, ED . The second component is the one consumed into elastic deformation, EE . Both values can be extracted from the energy-time plot. The difference between the maximum energy on the plot and the energy at the end of the impact event is EE , whereas the energy at the end of the impact event is ED . There is another component that is being
4. Results and discussion 4.1. Impact results In order to examine the repeatability of the impact tests, sample load-time plots are shown in Fig. 4. Only impact energies with more than one specimen are considered in this comparison. For the plots shown in Fig. 4 and other similar plots, the impact tests show excellent repeatability which is usually experienced elsewhere [21]. Typical half-sinusoidal load-time diagrams are obtained for the specimens impacted with smaller impact energies as a result of the very localized impact damage [25]. As the impact energy increases, large drops in the load-time plots are usually obtained due to the uncontrolled propagation of delaminations [26–28].
lost into friction of the machine component, etc., EL . The impact energy can be defined as Ei ¼ ED þ EE þ EL . In our case, the value of EL is measured for all the impact events, and its value is 0:210:05 J. In comparison to the desired impact energies, the value of the energy losses is less than 5% of the minimum energy value and can be neglected for the rest of the analysis. The energy values consumed in damage ðED Þ as a function of the impact energy can be shown in Fig. 6(a). A summary of the maximum displacement and the displacement at 0 kN load is also shown in Fig. 6(b). It is clear on the plots in Fig. 6 that all the plates (BL, A1, and A2) behaved in the same way under low velocity impact up to 30 J impact energy. After 30 J, the laminate A2 showed an improvement of 15% in the energy consumed in damage in comparison to BL and A1. This means that the laminate A2 consumed more energy in elastic deformation than the other laminates. This is clear on
34
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Upper clamp with 75 mm diameter hole
V0 Pneumac clamping pressure
Composite Speciemen
Lower clamp with 75 mm diameter hole
Locaon system for the specimen
Fig. 2. Schematic sketch of the components of the location and clamping systems used for the impact tests.
Table 1 Impact energies, mass, velocities and drop weight heights. Impact energy (J)
Impact velocity (m/s) Impact height (mm) Number of specimens
5
7.5
10
15
20
30
45
1.11 62.5 2
1.36 93.8 1
1.57 125.1 2
1.92 187.6 1
2.22 250.1 1
2.71 375.2 2
3.32 562.8 1
Fig. 3. Compression after impact test setup.
Fig. 6(b), i.e., laminate A2 bent with the smallest value of the maximum displacement ðdmax Þ and while rebound or unloading, it remained in contact with the impactor for longer distances. This resulted in the minimum dF¼0N . The load–displacement and energy-time plots at 45 J as shown in Fig. 7 can be easily used to justify this point. As they all behave in the same way at lower energy values, the difference after 30 J impact energy is due to the appearance of fiber breakage in BL and A1 before A2. This result shows that introducing thin-plies as a cover to normal plies highly delayed fiber breakage. However, higher energy values should be checked to generalize this phenomenon. 4.2. Damage assessment A summary of the value of the damage area measured by the cscan for all the specimens and impact energies presented in Table 1
is shown in Fig. 8. The shapes of the projected delamination areas at various impact energies are shown in Fig. 9. The projected delamination area, Fig. 8, showed similar values as the impact energy increased from 5 J to 10 J. For impact energies higher than 10 J, the projected delamination area of the laminate A1 was very high in comparison to the other laminates. The laminate A2 started to have the same trend at 45 J impact energy as compared to BL. The reason for this response could be justified by the shape of the damage area, Fig. 9. For all the laminates at 5, 7.5, and 10 J impact energies, the projected delamination area was very concentrated at the impact location. For higher energy values, the projected delamination area of the BL laminate extended in the traditional manner to form the well-known circular profile which is a projection of individual peanut shaped delaminations [32]. For the A1 laminate, the damage started in the traditional plies by normal delaminations forming the central continuous damage area. As the damage increased in the traditional plies, the stiffness decreased which led to higher deformation by which the thin plies started to crack forming these discontinuations in the projected damage area. The discontinuity in the damage area implies that these delaminations were formed at different interfaces, and they are, most probably, not connected. Although they resulted in larger area, these unconnected threads might lead to improvement in the damage tolerance due to the discontinuity of the damage. For the A2 laminate, the same phenomenon existed. However, due to the lower deformation (dmax is smaller which led to smaller bending strain) at higher impact energies, which resulted from the distribution of the thin-plies surrounding the normal ones, Fig. 6(b), the discontinuity decreased, and the formation of the unconnected thread was delayed up to 45 J impact energy. The pattern obtained for delamination in the laminate A1 can be justified by the potential of the thin-plies to decrease the energy release rates, delaying the propagation of both intra- and interlaminar cracks [15]. Jalalvand et al. [33,34] and Czél and Wisnom [15] presented analytical models for different damage mechanisms that can be predicted in laminates made of hybrid thin and traditional plies. For such configurations, the authors of these papers suggested that the low strain to failure plies are expected to fail first with extra matrix cracking. This cracking is followed by small delaminations, and instead of delamination propagation, the analytical models showed the possibility to have multiple small delaminations at the same interface up to the final saturation. Although it was developed for UD hybrid laminates, the phenomena can be used to justify the unconnected threads observed for the laminate A1 and A2, as shown in Fig. 9.
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
35
6 5 4 3 2 1 0
0
1
2
3
0
1
2
3
4
5
6
8 7 6 5 4 3 2 1 0
4
5
6 Fig. 5. Maximum load and the delamination threshold load as functions of the impact energy.
16 14 12 10 8 6 4 2 0 0
Fig. 4. Sample thresholds).
1
load-time
2
plots
3
(centerlines
4
represents
5
6
the
delamination
4.3. CAI results The failure load and stress of the specimens under compression after impact (CAI) are shown in Fig. 10. As noticed on the plots there is a higher compressive strength of the non-impacted specimens of the second alternative (A2) as compared to the other specimens. Although they have the same in-plane stiffness, an improvement of 10% is recorded for the laminate A2 over the other two laminates. This improvement can be justified from the in situ strength of each individual ply. Taking into account that the in situ phenomenon is a fracture based property rather than a strength related one [35,36], the current results reveal that using the mix of thin and normal plies in the same panel highly changes the fracture toughness properties. It can also be noted that all the configurations start the CAI strength degradation at the same impact
Fig. 6. Energy and displacement history as functions of the impact energy for the three laminates.
36
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 9. Projected delamination area at some impact energies as measured by the cscan.
Fig. 7. Load, displacement and energy history for the three laminates at 45 J impact energy.
Fig. 8. Summary of the projected delamination area for the three laminates at different impact energy.
energy (around 10 J). This result shows that the effect of adding thin-plies in the two proposed schemes is insignificant in terms of the impact energy required to initiate the strength degradation. The laminate A2 kept its improved strength for the whole range of the examined impact energy. The compression after impact of the two alternatives A1 and A2 as normalized to the strength of the BL laminate at the same impact energy is shown in Fig. 11. The result showed that the CAI strength of the A1 laminate is either lower than or equivalent to the baseline laminate for most of the examined impact energies and even for the non-impacted specimens. The maximum reduction of 10% in the CAI strength was measured at 7.5 J impact energy. This result revealed that clustering of thin plies at the middle of the specimen has insignificant effect on the compression
Fig. 10. Compression after impact load and strength.
after impact strength. On the other hand, for the A2 laminate, the compression after impact strength is higher than that for the BL for all the examined impact energies. An improvement of up to 15% is recorded in the compression after impact strength in comparison to the traditional normal ply laminates.
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 11. Compression after impact strength normalized to its value of BL.
5. Conclusions The current study aimed to propose a hybrid system of the traditional plies and the thin-plies in order to indicate the advantages of both of them (especially the thin-plies) in damage resistance and impact tolerance of laminated composites. Two laminates with thin-plies of TeXtreme fabrics were designed and tested in addition to the baseline which is made of all normal ply thickness. The specimens were then impacted using the drop-weight tower with impact energies ranging from 5 J to 45 J. After impact, the specimens were c-scanned and then compressed in order to define the residual strength. The result of the impact test showed that up to 30 J impact energy there is no difference between the three laminates in the contact time, displacement, maximum load, absorbed energy, etc. Only at 45 J, the laminate A2 experienced a lower value of the absorbed energy and the displacement which implies that, for this laminate, the fiber breakage was delayed in comparison to the other laminates. However, this idea still needs several investigations for impact energies higher than 45 J. The most significant difference in the impact result is the delamination threshold. As expected, the delamination threshold of the specimens containing thin-plies is higher than that of the normal plies due to the effect of the ply thickness on the in situ strength. The projected delamination area for the laminates containing thin-plies is the same as the normal plies up to a certain limit (depends on the laminate), and then, it showed higher values, compared to the BL. The larger damage area results from the formation of unconnected, narrow, and dispersed delaminations. The results of the compression after impact showed the importance of this study. The use of both the thin-plies and the normal ones in the same laminate guaranteed the benefits of the higher strength from the thin-plies and avoid the brittle nature of thin-ply composites. Improvement of up to 15% in the compression after impact was recorded to the laminate A2 which was designed to have each traditional ply surrounded by two thin-plies. Acknowledgement This paper was made possible by NPRP grant # NPRP05-12982-560 from the Qatar National Research Fund (a member of Qatar Foundation). The statements made herein are solely the responsibility of the authors. References [1] Mazumdar SK. Composites manufacturing, materials, product and process engineering. Florida, USA: CRC Press LLC; 2002. [2] Kawabe K, Tomoda S, Matsuo T. A pneumatic process for spreading reinforcing fiber tow. In: The 42nd international SAMPE symposium and exhibition. Anaheim, CA, May; 1997.
37
[3] Sasayama H, Kawabe K, Tomoda S, Kageyama K, Ogata N. Effect of lamina thickness on first ply failure in multidirectionally laminated composites. In: The 8th Japan international SAMPE symposium and exhibition (JISSE-8). Tokyo, Japan, November; 2003. [4] Tsai ST, Kawabe K. Thin ply laminates. Tech rep EP 1 793 989 B1. European patent; 2007. [5] Saito H, Morita M, Kawaba K, Kanesaki M, Takeuchi H, Tanaka M, et al. Effect of ply-thickness on impact damage morphology in CFRP laminates. J Compos Mater 2011;30:1097–106. [6] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Notched response of non-crimp fabric thin-ply laminates. Compos Sci Technol 2013;79:97–114. [7] Shin S, Kim RY, Kawabe K, Tsai SW. Experimental studies of thin-ply laminated composites. Compos Sci Technol 2007;67:699–1008. [8] Saito H, Takeuchi H, Kimpara I. A study of crack suppression mechanism of thin-ply carbon-fiber-reinforced polymer laminate with mesoscopic numerical simulation. J Compos Mater 2014;48:2085–96. [9] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Large damage capability of noncrimp fabric thin-ply laminates. Composites: Part A 2014;63:110–22. [10] Guillamet J, Turon A, Costa J, Renart J, Linde P, Mayugo JA. Damage occurrence at edges of non-crimp-fabric thin-ply laminates under off-axis uniaxial loading. Compos Sci Technol 2014;98:44–50. [11] Fuller JD, Wisnom MR. Pseudo-ductility and damage suppression in thin ply CFRP angle-ply laminates. Composites: Part A 2015;69:64–71. [12] Fuller JD, Wisnom MR. Exploration of the potential for pseudo-ductility in thin ply CFRP angle-ply laminates via an analytical method. Compos Sci Technol 2015;112:8–15. [13] Moon JB, Kim MG, Kim CG, Bhowmik S. Improvement of tensile properties of CFRP composites under LEO space environment by applying MWNTs and thinply. Composites: Part A 2011;42:694–701. [14] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Notched response of non-crimp fabric thin-ply laminates: analysis methods. Compos Sci Technol 2013;88:165–71. [15] Czél G, Wisnom MR. Demonstration of pseudo-ductility in high performance glass/epoxy composites by hybridisation with thin-ply carbon prepreg. Composites: Part A 2013;52:23–30. [16] Yokozeki T, Aoki Y, Ogasawara T. Experimental characterization of strength and damage resistance properties of thin-ply carbon fiber/toughened epoxy laminates. Compos Struct 2008;82:382–9. [17] Yokozeki T, Kuroda A, Yoshimura A, Ogasawara T, Aoki T. Damage characterization in thin-ply composite laminates under out-of-plane transverse loadings. Compos Struct 2010;93:49–57. [18] González EV, Maimí P, Camanho PP, Lopes CS, Blanco N. Effect of ply clustering in laminated composite plates under low-velocity impact loading. Compos Sci Technol 2011;71:805–17. [19] González EV, Maimí P, Camanho PP, Turon A, Mayugo JA. Simulation of dropweight impact and compression after impact tests on composite laminates. Compos Struct 2012;94:3364–78. [20] Amacher R, Cugnoni J, Botsis J, Sorensen L, Smith W, Dransfeld C. Thin ply composites: experimental characterization and modeling of size-effects. Compos Sci Technol 2014;101:121–32. [21] Sebaey TA, González EV, Lopes CS, Blanco N, Maimi P, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: effect of the mismatch angle between plies. Compos Struct 2013;101:255–64. [22] AITM1-0010. Fiber reinforced plastics: determination of compression strength after impact. Airbus test method. Blagnac Cedex, France; 2005. [23] ASTM-D7136/D7136M. Standard test method for measuring the damage resistance of a fiber-reinforced polymer matrix composite to a drop-weight impact event. West Conshohocken, PA, USA: ASTM International, American Society for Testing and Materials; 2007. [24] ASTM-D7137/D7137M. Standard test method for compressive residual strength properties of damaged polymer matrix composite plates. West Conshohocken, PA, USA: ASTM International, American Society for Testing and Materials; 2007. [25] Lopes CS, Seresta O, Coquet Y, Gürdal Z, Camanho PP, Thuis B. Low-velocity impact damage on dispersed stacking sequence laminates. Part I: experiments. Compos Sci Technol 2009;69:926–36. [26] Olsson R, Donadon MV, Falzon BG. Delamination threshold load for dynamic impact on plates. Int J Solids Struct 2006;43:3124–41. [27] Davies GAO, Zhang X, Zhou G, Watson S. Numerical modeling of impact damage. Composites 1994;25:342–50. [28] Sebaey TA, González EV, Lopes CS, Blanco N, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: design and optimization. Compos Struct 2013;95:569–76. [29] Sebaey TA, González EV, Lopes CS, Blanco N, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: the bending stiffness effect. Compos Struct 2013;106:30–2. [30] Lopes CS, Camanho PP, Gürdal Z, Maimí P, González EV. Low-velocity impact damage on dispersed stacking sequence laminates. Part II: numerical simulations. Compos Sci Technol 2009;69:937–47. [31] Sebaey TA, González EV, Lopes CS, Blanco P, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: effect of ply clustering. Compos Struct 2013;106:96–103. [32] Abrate S. Impact on composite structures. second ed. Cambridge, UK: Cambridge University Press; 1998.
38
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
[33] Jalalvand M, Czél G, Wisnom MR. Damage analysis of pseudo-ductile thin-ply UD hybrid composites – a new analytical method. Composites: Part A 2015;69:83–93. [34] Jalalvand M, Czél G, Wisnom MR. Parametric study of failure mechanisms and optimal configurations of pseudo-ductile thin-ply UD hybrid composites. Composites: Part A 2015;74:123–31.
[35] Nairn JA, Hu S, Bark S. A critical evaluation of theories for predicting microcracking in composite laminates. J Mater Sci 1993;28:5099–111. [36] Nairn JA. Matrix microcracking in composites. In: Talreja R, Månson JA, editors. Polymer matrix composites. Elsevier Science; 2000. p. 403–32.
Contents lists available at ScienceDirect
Composites: Part A journal homepage: www.elsevier.com/locate/compositesa
Using thin-plies to improve the damage resistance and tolerance of aeronautical CFRP composites T.A. Sebaey ⇑,1, E. Mahdi Mechanical and Industrial Engineering Department, College of Engineering, Qatar University, 2713 Doha, Qatar
a r t i c l e
i n f o
Article history: Received 17 October 2015 Received in revised form 19 March 2016 Accepted 24 March 2016 Available online 1 April 2016 Keywords: A. Hybrid B. Compression after impact B. Damage tolerance B. Impact behavior Thin-plies
a b s t r a c t Thin-ply composites are currently receiving specific attention from researchers due to their capabilities to delay matrix cracking. In this paper, the aim is to design a hybrid laminate that contains both thin- and normal plies. The objective is to improve the tolerance of normal plies by adding thin-plies to the composite in different configurations. Two alternatives were designed, tested, and compared to the specimens made of traditional plies. Impact and compression after impact tests were conducted on each configuration at different impact energies. After being impacted, the specimens were c-scanned to define the delamination pattern. Results showed that surrounding each normal ply with two thin-plies improved the delamination threshold by 15% as compared to the specimens made all of normal plies. Under compression, 15% improvements in the compression after impact strength were obtained. By using thin-plies, the size of each individual delamination was reduced, resulting in small threads instead of peanut delaminations. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Laminated composites are currently being used in several structural and non-structural applications including aeronautics, automotive, marine, and piping. The main advantages of laminated composite materials are the high specific strength and stiffness as well as the dimensional stability in comparison to the conventional metallic structures. Conventional, well-known composites are made of several layers of fiber reinforced plastics of a nominal ply thickness greater than 0.1 mm [1]. Reducing the ply thickness down to this limit is usually referred to as thin-plies. Few efforts were found in the literature to reduce the ply thickness beyond this limit [2,3]. In 2007, Tsai and Kawabe [4] published a patent that addressed a new tow spreading technology that enables the production of plies of thickness down to 0.02 mm thickness. The report also addressed some of the potential improvements in the structural response by using such laminates. The interest in thinplies can be justified by the possibility of improving the ply in situ strength as the ply thickness is reduced [5]. Detailed potential advantages of using thin-plies in composite design are summarized by Arteiro et al. [6]. Shin et al. [7] presented
⇑ Corresponding author. E-mail addresses: [email protected], [email protected] (T.A. Sebaey). On leave from Mechanical Design and Production Dept., Faculty of Engineering, Zagazig University, 44519 Zagazig, Sharkia, Egypt. 1
http://dx.doi.org/10.1016/j.compositesa.2016.03.027 1359-835X/Ó 2016 Elsevier Ltd. All rights reserved.
a study on using thin-plies (0.04 mm ply thickness) in comparison to traditional plies of 0.2 mm thickness. The comparison was made based on the results of the unnotched tensile, open-hole tensile, and impact as well as compression after impact tests. Tensile tests were conducted both statically and under fatigue. Two different types of specimens were considered. Both laminates composed of the same amount of fiber in each direction. For the first stacking sequence ‘‘THIN” the authors did not allow clustering resulting in a nominal ply thickness of 0.04 mm whereas, for the second one ‘‘THICK”, each 5 layers of the same orientation angle were clustered together resulting in a nominal layer thickness of 0.2 mm. The results of the tensile tests showed an improvement of 10% in both the strength and stiffness. In addition, less sensitivity to the number of cycles was recorded by using thin-plies. In open hole tensile tests, the advantage was given to the thick-plies with 10% higher strength and stiffness in comparison to the thin-plies. The authors justified this response by the stress relaxation near the hole edge after the initial failure. As a general observation in all the tests, the fracture is more brittle in comparison to the traditional ones. These improvements were justified by the stress relaxation and the energy release rate at the crack tip through meso-scale FE modeling by Saito et al. [8]. In addition, the capability of thin-plies to resist the propagation of large cracks does play a role in this improvement [9]. This higher capability helped also to push the damage away of the thin-plies in the experiments designed by Guillamet et al. [10] to monitor the damage at the free edge under
32
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
tension. Under tensile loading, Fuller and Wisnom [11,12] presented an experimental and analytical study to investigate the behavior of ½h5 s ; h = 15, 20, 25, 30, and 45 thin-plies. In addition to the good agreement between the experimental and the analytical predictions, for all angles, delaminations were suppressed, allowing considerable pseudo-ductile strains to develop. Other tensile tests were preformed by Moon et al. [13] under harsh low earth orbit (LEO) conditions. The authors used two different ply thicknesses of 0.125 mm and 0.06 mm. The authors measured the tensile strength and stiffness of both laminates before and after the LEO aging. The authors concluded that using the thin-plies resulted in improvements of 8.9% and 12.1% at ambient and LEO conditions, respectively. On the contrary, reduction of the stiffness of 8.6% and 12.8% was recorded by using the thin-plies at both environmental conditions, respectively. It is worth noting that the fiber volume fraction of the two specimens was different (about 10% difference) which makes it very difficult to draw a clear conclusion from these results. The notch sensitivity of the thinplies were assessed (experimentally and using the finite fracture mechanics approach) by Arteiro et al. [6,14] for two different stacking sequences and compared with results from the literature for normal plies. The results showed that the notch sensitivity of laminates made of thin-plies can be equivalent to that made of thick plies. It is worth remarking that the brittleness is a characteristic of the damage of laminates made of thin plies as it was noted by others. Based on the fracture mechanics approach, Czél and Wisnom [15] suggested a hybridisation system between glass and carbon fibers layers that introduced some ductile characteristics to the damage of the thin-ply composites. Yokozeki et al. [16] compared the compression properties of laminates made of thin-plies of 0.07 mm ply thickness with the ones made of normal ply thickness of 0.14 mm. The authors performed compression, open-hole compression, and compression after impact tests. The results showed improvements of 16%, 9% and 8% for the three mentioned compression tests, respectively. The authors reported a significant reduction in the amount of damage as measured by the X-ray and the acoustic emission techniques. Under out-of-plane loading, Yokozeki et al. [17] compared the results of the two test specimens using indentation tests. The results showed that in thin-plies damage initiated by accumulation of matrix cracking and localized delaminations and then sudden fiber breakage. For standard plies (0.14 mm ply thickness), more visible delaminations appeared after matrix cracking and then fiber breakage. The authors concluded that thin-ply laminates have high damage resistance against matrix cracking and delaminations near the back surface due to the absent of back face delaminations. A limited number of studies addressed the impact of thin-ply laminates, and if found, it is a part of the study with few impact energies and limited discussions on the results. Yokozeki et al. [16] compared the impact properties of laminates made of thinplies of 0.07 mm ply thickness with the ones made of normal ply thickness of 0.14 mm at 6.7 J impact energy. The results did not show any significant effect on the damage area. The reason could be the lower value of the impact energy which, most probably, is lower than the delamination threshold [18,19]. For three different ply thicknesses, Amacher et al. [20] compared the impact response. The results showed that reducing the ply thickness down to 30 lm resulted in a brittle failure with extensive translaminar cracks. Moreover, the optimum ply thickness was found to be 100 lm which is almost the thickness of the traditional plies. Based on the results and the conclusions summarized, it can be inferred that using thin-plies results in brittle damage which usually leads to catastrophic undesirable failure with some improved strength. However, for low velocity impacts, there were no
recorded improvements in any of the parameters used to assess the damage tolerance in the literature. In the current paper, the woven thin-plies were used inside the laminate with the traditional plies in two different ways. The idea is to improve the damage tolerance and the impact resistance as well as avoid the brittle failure mode obtained by Amacher et al. [20]. Low velocity impacts were conducted at different energy levels and impact velocities. 2. Materials and manufacturing Spread tow fabrics/epoxy composites were supplied by Oxeon. Two grades of woven carbon fibers fabrics were used: TeXtreme 80 PW UTS50S of 80 gsm and TeXtreme 320 PW UTS50S of 320 gsm. One TeXtreme +45/45 fabric layer is equivalent to four layers of unidirectional. The nominal ply thickness of the traditional (320 gsm) and the thin-plies (80 gsm) were 85 and 330 lm, respectively. The MTFA500 (DF044) epoxy resin (delivered by SHD composites) was used as a matrix to manufacture the specimen in an autoclave based process. Considering the previous results on low velocity impact of thinplies and to avoid brittle failure, hybrid specimens were manufactured from both traditional and thin-plies and the damage resistance was compared with specimens made of all traditional plies. Three stacking sequences were considered. The first group of specimens was the baseline or the reference laminate. The reference laminate consisted of 12 layers of TeXtreme 320 PW. The fiber orientations of the different layers are shown in Fig. 1. It was considered, during the design stage to have the same amount of fiber in 0°, 90°, and 45°. The total laminate thickness was 3:980:038 mm. In the first alternative (A1), the laminate was designed to have traditional plies at the specimen surfaces and a block of thin-plies at the mid-plane. The idea of this design was initiated by Sebaey et al. [21]. In that paper, the authors supposed that consuming the impact energy, by different damage mechanisms, close to outer surfaces can leave a cluster of the undamage layers at the middle. This cluster could withstand more compression loading after impacts, and consequently, the damage tolerance is improved. To push the damage outside, the core of the specimen was designed by using the thin-plies technology. To have a fair comparison, the amount of fiber in each direction is designed to be equal to that of the laminate BL. The total thickness of the A1 laminate is 4:100:027 mm. For the second alternative (A2 laminate), the effect of the neighboring plies on the strength of each layer was taken into account. It is well-known that the crack initiation/propagation is highly affected by the neighboring plies (i.e., the higher the strength and stiffness of the neighboring plies, the higher the damage resistance of the ply under consideration) [8]. For this reason, in our design (A2), each traditional ply is surrounded by two thin-plies to improve its damage resistance. The total thickness of the A2 laminate is 4:110:033 mm. With those two designs, the user can benefit from the higher strength of the thin-plies and avoid the brittle nature associated with the laminates made all of thin-plies under impact loading. The three plates were manufactured with 700 700 mm. Then, the margins (40 mm all around) were cut. The fiber volume fraction was measured by the ignition test standard, ASTM D258411. The values of the fiber volume fraction were 50:410:52 %, 49:091:03 % and 50:010:28 % for BL, A1 and A2, respectively. 3. Impact and CAI tests After being cured, the plates were scanned using c-scan, and the positions with defects were eliminated. The plates were then cut using a computer controlled diamond wheel into low velocity
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 1. Schematic of the cross-section properties of the three stacking sequences. BL refers to the reference laminate whereas, A1 and A2 represent the alternatives 1 and 2, respectively.
impact specimens. The nominal in-plane dimensions of the impact specimen are 100 150 mm as per the test standards [22–24]. The test setup of the impact events is shown in Fig. 2. The two halves of the clamping system, after being collected together with the pneumatic clamp, provided an unclamped area of 75 mm diameter. An impactor of 16 mm diameter was impacting the specimens at different impact velocities [22]. The impactor mass ðM i Þ was chosen to be constant at 8.15 kg. The two side location pins were added to the standard IMATEK test jig provided with the machine in order to ensure central impacts. The impact energies ðEi Þ and the correpffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sponding velocity ðV 0 ¼ 2Ei =M i Þ and height ðHi ¼ Ei =ðM i gÞÞ values are shown in Table 1. After being impacted, the specimens were taken to the c-scan to assess the projected delamination area associated with each impact energy level. The Compression After Impact (CAI) tests were performed using a Hydraulic test machine and a 250 kN load-cell according to the guidelines introduced in the AITM1-0010 [22] and the ASTM D7137-07 [24]. Test setup is introduced in Fig. 3 test speed was set as 1 mm/min. The compression after impact strength ðrCAI Þ was calculated based on the ultimate load ðP max Þ and the specimen cross-section dimensions (rCAI ¼ Pmax =ðb hÞ, where b and h were the specimen width and thickness, respectively). In addition to the specimens summarized in Table 1, two non-impacted specimens of each stacking sequence were compressed up to failure to determine the non-impacted strength ðr0 Þ. The compression test setup of the non-impacted specimen is exactly as the compression after impact test. The residual strength can be calculated as the ratio between the CAI strength and the non-impacted specimen strength.
33
A summary of the peak loads and the delamination threshold loads are presented in Fig. 5. The delamination threshold can be defined as the load level at which the first unstable delamination appears. On the load–displacement curve, this load can be considered as the first sharp decrease in the stiffness [28]. The delamination threshold is shown on the load-time plot as a horizontal line, Fig. 4. For the impact event in Fig. 4(a), the maximum load is lower than the delamination threshold and, for that reason, the horizontal line does not appear. For the maximum load, the load is linearly proportional to the impact energy up to a certain limit (between 30 J and 45 J impact energies). After that limit, the maximum load seems to be constant. This response was also experienced elsewhere [29] for multidirectional laminates. The reason of this behavior could be justified by the initiations of fiber breakage at this energy limit. It is worth remarking that the current TeXtreme fabrics (either with or without the thin-plies) provide higher fiber breakage limit in comparison to the normal laminates made of UD fibers [21,25,29–31] with similar laminate thickness. The comparison between the laminates BL, A1, and A2 does not show any advantage to any of the laminates for up to 15 J impact energy. For impact energies higher than 15 J, the maximum load is usually higher in case of BL, as compared to A1 and A2. The difference between the three laminates can be seen when comparing the delamination threshold load. As it is mainly used to delay the matrix cracking and consequently the other damage mechanisms, the laminates that contain thin-plies have a higher delamination threshold than the ones that do not have thin-plies. The delamination threshold of each configuration shows constant value with respect to the impact energy values. This phenomenon was noted for other tested in the literature [18,21,25]. The values of the delamination threshold load are 4:790:12 , 5:560:17 , and 5:480:28 for the BL, A1, and A2 laminates, respectively. The comfor A1 and A2 does not show parison between the values of F Thr d any significant difference. This result could be justified by the higher number of ply interfaces resulted from using thin plies (to obtain the same thickness, more plies were used and hence, the number of interfaces increased). With higher number of interfaces, the energy is expected to be absorbed in higher number of narrow delaminations, as compared to the traditional plies. This delayed the appearance of the first unstable delamination and, consequently delayed the delamination threshold. The value of the impact energy Ei can be divided into two main components. The first component is energy required to initiate and propagate the different damage mechanisms, ED . The second component is the one consumed into elastic deformation, EE . Both values can be extracted from the energy-time plot. The difference between the maximum energy on the plot and the energy at the end of the impact event is EE , whereas the energy at the end of the impact event is ED . There is another component that is being
4. Results and discussion 4.1. Impact results In order to examine the repeatability of the impact tests, sample load-time plots are shown in Fig. 4. Only impact energies with more than one specimen are considered in this comparison. For the plots shown in Fig. 4 and other similar plots, the impact tests show excellent repeatability which is usually experienced elsewhere [21]. Typical half-sinusoidal load-time diagrams are obtained for the specimens impacted with smaller impact energies as a result of the very localized impact damage [25]. As the impact energy increases, large drops in the load-time plots are usually obtained due to the uncontrolled propagation of delaminations [26–28].
lost into friction of the machine component, etc., EL . The impact energy can be defined as Ei ¼ ED þ EE þ EL . In our case, the value of EL is measured for all the impact events, and its value is 0:210:05 J. In comparison to the desired impact energies, the value of the energy losses is less than 5% of the minimum energy value and can be neglected for the rest of the analysis. The energy values consumed in damage ðED Þ as a function of the impact energy can be shown in Fig. 6(a). A summary of the maximum displacement and the displacement at 0 kN load is also shown in Fig. 6(b). It is clear on the plots in Fig. 6 that all the plates (BL, A1, and A2) behaved in the same way under low velocity impact up to 30 J impact energy. After 30 J, the laminate A2 showed an improvement of 15% in the energy consumed in damage in comparison to BL and A1. This means that the laminate A2 consumed more energy in elastic deformation than the other laminates. This is clear on
34
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Upper clamp with 75 mm diameter hole
V0 Pneumac clamping pressure
Composite Speciemen
Lower clamp with 75 mm diameter hole
Locaon system for the specimen
Fig. 2. Schematic sketch of the components of the location and clamping systems used for the impact tests.
Table 1 Impact energies, mass, velocities and drop weight heights. Impact energy (J)
Impact velocity (m/s) Impact height (mm) Number of specimens
5
7.5
10
15
20
30
45
1.11 62.5 2
1.36 93.8 1
1.57 125.1 2
1.92 187.6 1
2.22 250.1 1
2.71 375.2 2
3.32 562.8 1
Fig. 3. Compression after impact test setup.
Fig. 6(b), i.e., laminate A2 bent with the smallest value of the maximum displacement ðdmax Þ and while rebound or unloading, it remained in contact with the impactor for longer distances. This resulted in the minimum dF¼0N . The load–displacement and energy-time plots at 45 J as shown in Fig. 7 can be easily used to justify this point. As they all behave in the same way at lower energy values, the difference after 30 J impact energy is due to the appearance of fiber breakage in BL and A1 before A2. This result shows that introducing thin-plies as a cover to normal plies highly delayed fiber breakage. However, higher energy values should be checked to generalize this phenomenon. 4.2. Damage assessment A summary of the value of the damage area measured by the cscan for all the specimens and impact energies presented in Table 1
is shown in Fig. 8. The shapes of the projected delamination areas at various impact energies are shown in Fig. 9. The projected delamination area, Fig. 8, showed similar values as the impact energy increased from 5 J to 10 J. For impact energies higher than 10 J, the projected delamination area of the laminate A1 was very high in comparison to the other laminates. The laminate A2 started to have the same trend at 45 J impact energy as compared to BL. The reason for this response could be justified by the shape of the damage area, Fig. 9. For all the laminates at 5, 7.5, and 10 J impact energies, the projected delamination area was very concentrated at the impact location. For higher energy values, the projected delamination area of the BL laminate extended in the traditional manner to form the well-known circular profile which is a projection of individual peanut shaped delaminations [32]. For the A1 laminate, the damage started in the traditional plies by normal delaminations forming the central continuous damage area. As the damage increased in the traditional plies, the stiffness decreased which led to higher deformation by which the thin plies started to crack forming these discontinuations in the projected damage area. The discontinuity in the damage area implies that these delaminations were formed at different interfaces, and they are, most probably, not connected. Although they resulted in larger area, these unconnected threads might lead to improvement in the damage tolerance due to the discontinuity of the damage. For the A2 laminate, the same phenomenon existed. However, due to the lower deformation (dmax is smaller which led to smaller bending strain) at higher impact energies, which resulted from the distribution of the thin-plies surrounding the normal ones, Fig. 6(b), the discontinuity decreased, and the formation of the unconnected thread was delayed up to 45 J impact energy. The pattern obtained for delamination in the laminate A1 can be justified by the potential of the thin-plies to decrease the energy release rates, delaying the propagation of both intra- and interlaminar cracks [15]. Jalalvand et al. [33,34] and Czél and Wisnom [15] presented analytical models for different damage mechanisms that can be predicted in laminates made of hybrid thin and traditional plies. For such configurations, the authors of these papers suggested that the low strain to failure plies are expected to fail first with extra matrix cracking. This cracking is followed by small delaminations, and instead of delamination propagation, the analytical models showed the possibility to have multiple small delaminations at the same interface up to the final saturation. Although it was developed for UD hybrid laminates, the phenomena can be used to justify the unconnected threads observed for the laminate A1 and A2, as shown in Fig. 9.
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
35
6 5 4 3 2 1 0
0
1
2
3
0
1
2
3
4
5
6
8 7 6 5 4 3 2 1 0
4
5
6 Fig. 5. Maximum load and the delamination threshold load as functions of the impact energy.
16 14 12 10 8 6 4 2 0 0
Fig. 4. Sample thresholds).
1
load-time
2
plots
3
(centerlines
4
represents
5
6
the
delamination
4.3. CAI results The failure load and stress of the specimens under compression after impact (CAI) are shown in Fig. 10. As noticed on the plots there is a higher compressive strength of the non-impacted specimens of the second alternative (A2) as compared to the other specimens. Although they have the same in-plane stiffness, an improvement of 10% is recorded for the laminate A2 over the other two laminates. This improvement can be justified from the in situ strength of each individual ply. Taking into account that the in situ phenomenon is a fracture based property rather than a strength related one [35,36], the current results reveal that using the mix of thin and normal plies in the same panel highly changes the fracture toughness properties. It can also be noted that all the configurations start the CAI strength degradation at the same impact
Fig. 6. Energy and displacement history as functions of the impact energy for the three laminates.
36
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 9. Projected delamination area at some impact energies as measured by the cscan.
Fig. 7. Load, displacement and energy history for the three laminates at 45 J impact energy.
Fig. 8. Summary of the projected delamination area for the three laminates at different impact energy.
energy (around 10 J). This result shows that the effect of adding thin-plies in the two proposed schemes is insignificant in terms of the impact energy required to initiate the strength degradation. The laminate A2 kept its improved strength for the whole range of the examined impact energy. The compression after impact of the two alternatives A1 and A2 as normalized to the strength of the BL laminate at the same impact energy is shown in Fig. 11. The result showed that the CAI strength of the A1 laminate is either lower than or equivalent to the baseline laminate for most of the examined impact energies and even for the non-impacted specimens. The maximum reduction of 10% in the CAI strength was measured at 7.5 J impact energy. This result revealed that clustering of thin plies at the middle of the specimen has insignificant effect on the compression
Fig. 10. Compression after impact load and strength.
after impact strength. On the other hand, for the A2 laminate, the compression after impact strength is higher than that for the BL for all the examined impact energies. An improvement of up to 15% is recorded in the compression after impact strength in comparison to the traditional normal ply laminates.
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
Fig. 11. Compression after impact strength normalized to its value of BL.
5. Conclusions The current study aimed to propose a hybrid system of the traditional plies and the thin-plies in order to indicate the advantages of both of them (especially the thin-plies) in damage resistance and impact tolerance of laminated composites. Two laminates with thin-plies of TeXtreme fabrics were designed and tested in addition to the baseline which is made of all normal ply thickness. The specimens were then impacted using the drop-weight tower with impact energies ranging from 5 J to 45 J. After impact, the specimens were c-scanned and then compressed in order to define the residual strength. The result of the impact test showed that up to 30 J impact energy there is no difference between the three laminates in the contact time, displacement, maximum load, absorbed energy, etc. Only at 45 J, the laminate A2 experienced a lower value of the absorbed energy and the displacement which implies that, for this laminate, the fiber breakage was delayed in comparison to the other laminates. However, this idea still needs several investigations for impact energies higher than 45 J. The most significant difference in the impact result is the delamination threshold. As expected, the delamination threshold of the specimens containing thin-plies is higher than that of the normal plies due to the effect of the ply thickness on the in situ strength. The projected delamination area for the laminates containing thin-plies is the same as the normal plies up to a certain limit (depends on the laminate), and then, it showed higher values, compared to the BL. The larger damage area results from the formation of unconnected, narrow, and dispersed delaminations. The results of the compression after impact showed the importance of this study. The use of both the thin-plies and the normal ones in the same laminate guaranteed the benefits of the higher strength from the thin-plies and avoid the brittle nature of thin-ply composites. Improvement of up to 15% in the compression after impact was recorded to the laminate A2 which was designed to have each traditional ply surrounded by two thin-plies. Acknowledgement This paper was made possible by NPRP grant # NPRP05-12982-560 from the Qatar National Research Fund (a member of Qatar Foundation). The statements made herein are solely the responsibility of the authors. References [1] Mazumdar SK. Composites manufacturing, materials, product and process engineering. Florida, USA: CRC Press LLC; 2002. [2] Kawabe K, Tomoda S, Matsuo T. A pneumatic process for spreading reinforcing fiber tow. In: The 42nd international SAMPE symposium and exhibition. Anaheim, CA, May; 1997.
37
[3] Sasayama H, Kawabe K, Tomoda S, Kageyama K, Ogata N. Effect of lamina thickness on first ply failure in multidirectionally laminated composites. In: The 8th Japan international SAMPE symposium and exhibition (JISSE-8). Tokyo, Japan, November; 2003. [4] Tsai ST, Kawabe K. Thin ply laminates. Tech rep EP 1 793 989 B1. European patent; 2007. [5] Saito H, Morita M, Kawaba K, Kanesaki M, Takeuchi H, Tanaka M, et al. Effect of ply-thickness on impact damage morphology in CFRP laminates. J Compos Mater 2011;30:1097–106. [6] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Notched response of non-crimp fabric thin-ply laminates. Compos Sci Technol 2013;79:97–114. [7] Shin S, Kim RY, Kawabe K, Tsai SW. Experimental studies of thin-ply laminated composites. Compos Sci Technol 2007;67:699–1008. [8] Saito H, Takeuchi H, Kimpara I. A study of crack suppression mechanism of thin-ply carbon-fiber-reinforced polymer laminate with mesoscopic numerical simulation. J Compos Mater 2014;48:2085–96. [9] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Large damage capability of noncrimp fabric thin-ply laminates. Composites: Part A 2014;63:110–22. [10] Guillamet J, Turon A, Costa J, Renart J, Linde P, Mayugo JA. Damage occurrence at edges of non-crimp-fabric thin-ply laminates under off-axis uniaxial loading. Compos Sci Technol 2014;98:44–50. [11] Fuller JD, Wisnom MR. Pseudo-ductility and damage suppression in thin ply CFRP angle-ply laminates. Composites: Part A 2015;69:64–71. [12] Fuller JD, Wisnom MR. Exploration of the potential for pseudo-ductility in thin ply CFRP angle-ply laminates via an analytical method. Compos Sci Technol 2015;112:8–15. [13] Moon JB, Kim MG, Kim CG, Bhowmik S. Improvement of tensile properties of CFRP composites under LEO space environment by applying MWNTs and thinply. Composites: Part A 2011;42:694–701. [14] Arteiro A, Catalanotti G, Xavier J, Camanho PP. Notched response of non-crimp fabric thin-ply laminates: analysis methods. Compos Sci Technol 2013;88:165–71. [15] Czél G, Wisnom MR. Demonstration of pseudo-ductility in high performance glass/epoxy composites by hybridisation with thin-ply carbon prepreg. Composites: Part A 2013;52:23–30. [16] Yokozeki T, Aoki Y, Ogasawara T. Experimental characterization of strength and damage resistance properties of thin-ply carbon fiber/toughened epoxy laminates. Compos Struct 2008;82:382–9. [17] Yokozeki T, Kuroda A, Yoshimura A, Ogasawara T, Aoki T. Damage characterization in thin-ply composite laminates under out-of-plane transverse loadings. Compos Struct 2010;93:49–57. [18] González EV, Maimí P, Camanho PP, Lopes CS, Blanco N. Effect of ply clustering in laminated composite plates under low-velocity impact loading. Compos Sci Technol 2011;71:805–17. [19] González EV, Maimí P, Camanho PP, Turon A, Mayugo JA. Simulation of dropweight impact and compression after impact tests on composite laminates. Compos Struct 2012;94:3364–78. [20] Amacher R, Cugnoni J, Botsis J, Sorensen L, Smith W, Dransfeld C. Thin ply composites: experimental characterization and modeling of size-effects. Compos Sci Technol 2014;101:121–32. [21] Sebaey TA, González EV, Lopes CS, Blanco N, Maimi P, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: effect of the mismatch angle between plies. Compos Struct 2013;101:255–64. [22] AITM1-0010. Fiber reinforced plastics: determination of compression strength after impact. Airbus test method. Blagnac Cedex, France; 2005. [23] ASTM-D7136/D7136M. Standard test method for measuring the damage resistance of a fiber-reinforced polymer matrix composite to a drop-weight impact event. West Conshohocken, PA, USA: ASTM International, American Society for Testing and Materials; 2007. [24] ASTM-D7137/D7137M. Standard test method for compressive residual strength properties of damaged polymer matrix composite plates. West Conshohocken, PA, USA: ASTM International, American Society for Testing and Materials; 2007. [25] Lopes CS, Seresta O, Coquet Y, Gürdal Z, Camanho PP, Thuis B. Low-velocity impact damage on dispersed stacking sequence laminates. Part I: experiments. Compos Sci Technol 2009;69:926–36. [26] Olsson R, Donadon MV, Falzon BG. Delamination threshold load for dynamic impact on plates. Int J Solids Struct 2006;43:3124–41. [27] Davies GAO, Zhang X, Zhou G, Watson S. Numerical modeling of impact damage. Composites 1994;25:342–50. [28] Sebaey TA, González EV, Lopes CS, Blanco N, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: design and optimization. Compos Struct 2013;95:569–76. [29] Sebaey TA, González EV, Lopes CS, Blanco N, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: the bending stiffness effect. Compos Struct 2013;106:30–2. [30] Lopes CS, Camanho PP, Gürdal Z, Maimí P, González EV. Low-velocity impact damage on dispersed stacking sequence laminates. Part II: numerical simulations. Compos Sci Technol 2009;69:937–47. [31] Sebaey TA, González EV, Lopes CS, Blanco P, Costa J. Damage resistance and damage tolerance of dispersed CFRP laminates: effect of ply clustering. Compos Struct 2013;106:96–103. [32] Abrate S. Impact on composite structures. second ed. Cambridge, UK: Cambridge University Press; 1998.
38
T.A. Sebaey, E. Mahdi / Composites: Part A 86 (2016) 31–38
[33] Jalalvand M, Czél G, Wisnom MR. Damage analysis of pseudo-ductile thin-ply UD hybrid composites – a new analytical method. Composites: Part A 2015;69:83–93. [34] Jalalvand M, Czél G, Wisnom MR. Parametric study of failure mechanisms and optimal configurations of pseudo-ductile thin-ply UD hybrid composites. Composites: Part A 2015;74:123–31.
[35] Nairn JA, Hu S, Bark S. A critical evaluation of theories for predicting microcracking in composite laminates. J Mater Sci 1993;28:5099–111. [36] Nairn JA. Matrix microcracking in composites. In: Talreja R, Månson JA, editors. Polymer matrix composites. Elsevier Science; 2000. p. 403–32.