Original Article
Development of numerical realistic model for predicting low-velocity impact response of aluminium honeycomb sandwich structures
Journal of Sandwich Structures and Materials 2016, Vol. 18(1) 95–112 ! The Author(s) 2015 Reprints and permissions: sagepub.co.uk/journalsPermissions.nav DOI: 10.1177/1099636215603047 jsm.sagepub.com
Recep Gunes1 and Kemal Arslan2
Abstract This paper presents the results of an experimental and numerical study on aluminium honeycomb sandwich structures under low-velocity impact loadings. In order to investigate the impact behavior of honeycomb sandwich structures, which is consisted of two identical aluminium facesheets and an aluminium honeycomb core, an experimental study was carried out by using a drop-weight impact test system. Using this system, the contact forces and absorbed energies were measured to determine the influence of impact energy for one configuration of the sandwich structure. According to these results, a numerical model by finite element method of sandwich structures was developed, which is in good agreement with experimental results in terms of contact forces and deformations. Later, the effect of the cell size and the height variation of aluminum honeycomb core on the impact response of sandwich structures were investigated using the improved numerical model. The obtained numerical and experimental results were interpreted in detail. Keywords Honeycomb sandwich structure, low-velocity impact, explicit finite element analysis
1 2
Department of Mechanical Engineering, Erciyes University, Kayseri, Turkey Graduate School of Natural and Applied Sciences, Erciyes University, Kayseri, Turkey
Corresponding author: Recep Gunes, Department of Mechanical Engineering, Erciyes University, Kayseri 38039, Turkey. Email:
[email protected]
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Introduction Sandwich structures are the oldest known form of composite materials. A sandwich structure has basically three important components, such as relatively thin top and bottom facesheets, relatively thick lightweight core material, and adhesive layers for providing the connection in the intermediate layers. The core material has relatively low strength and higher damping properties, while the facesheets have more resistant structure. In this way, the composite structure has been designed as both high mechanical strength and quite lightweight. The main task of the core material is to maintain the distance between the facesheets since the distance provides both higher bending rigidity and higher moment of inertia of cross-sectional area of the sandwich structure. Sandwich structures are used widely in industrial applications due to the properties such as high bending strength and stiffness, lightweight, low cost with suitable materials combination, thermal insulating property, sound-insulating property, creation of appropriate aerodynamic surfaces, and resistance to high speed. Especially, sandwich structures play an important role in aerospace, automotive and shipping industries, and military applications in which lightness and rigidity are very important. Honeycomb sandwich structures can be used for high mechanical strength structures that absorb the energy resulting from impact loads. Therefore, it is very important to know the low-velocity impact responses of this type structures. Hazizan and Cantwell [1] used a simple energy balance model to predict the low-velocity impact response of an aluminium honeycomb sandwich structure and investigated the accuracy of the model taking into account the changes in the energy of the falling impactor as well as the different dimensions of the sandwich structures. They found that the model successfully predicts the impact response of the sandwich structures and also the partitioning of energy strongly depends on the geometry and the configuration of the sandwich structure. Lin and Fat [2] developed an analytical model based on set of experimental results for the perforation of the composite sandwich panels subjected to quasi-static and low-velocity impact loading by hemispherical indenters/projectiles. They investigated the effect of the indenter/projectile geometry on the deformation and perforation, which involve consecutive failures of top facesheet, core, and bottom facesheet of the sandwich panels. Foo et al. [3] investigated the failure response of aluminium sandwich panels under low-velocity impact. For this purpose, they developed a three-dimensional geometrically correct finite element model of the honeycomb sandwich plate in order to exhibit further understanding of the parameters affecting the initiation and propagation of impact damage. Foo et al. [4] extended their previous work to include a progressive damage model for composites in order to accurately predict the damage mechanisms and failure of the composite sandwich structures. They showed that the model predictions are in good agreement with the experimental results. Lee et al. [5] presented some experimental and numerical results for the impact and indentation damages on the honeycomb sandwich panels subjected to low-velocity impact. They compared the experimental and numerical damage areas and depth created by impact loading, and obtained good agreement between the numerical and experimental results.
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Grisˇ kevicˇiusˇ et al. [6] investigated the impact behavior of sandwich composite made from woven glass fiber and polyvinylester resin composite facesheets and polypropylene honeycomb core. They presented results about the use in safety important structures depending on the geometric properties of the sandwich structure. Lee et al. [7] carried out experimental and numerical investigation on impact behavior of aluminum honeycomb sandwich panels. They developed a numerical model based on continuum damage mechanics, which is fully coupled with elastoviscoplasticity. Sharma et al. [8] investigated low-velocity impact response of aluminium honeycomb core sandwich panels by varying core height using a flat impactor. They focused on the influence of core height on energy absorption capacity of honeycomb panel and found that the variation of the core height does not significantly affect the energy absorbing capacity of the sandwich panel. Butukuri et al. [9] carried out experimental studies which are factorial-based design to determine the low-velocity impact response of sandwich panels. In order to determine the effect of different design factors on the low-velocity impact response of the honeycomb sandwich structures, response surface analyses of data from factorial experiments were performed by them. Manes et al. [10] performed experimental and numerical investigations on sandwich panels subjected to low-velocity impact. They built a highly detailed finite element model using data obtained from flatwise compressive tests of honeycomb cores and a calibration of the constitutive law and ductile failure of aluminium skins. Menna et al. [11] presented a numerical strategy in order to proper simulate the dynamic behavior of Nomex core sandwich structures combined with E-glass phenolic facesheet under impact load. The numerical model has been calibrated on a series of experimental outcomes and satisfactory agreement was obtained between numerical predictions and experimental results. In order to establish links between the sandwich structure performance and the material and geometrical parameters, a numerical model was developed for metallic sandwich structures as armor for aeronautical applications by Kolopp et al. [12]. Ivan˜ez and Sanchez-Saez [13] presented a numerical model for the low-velocity impact response of composite sandwich beams with honeycomb core. They carried out experimental tests to verify the numerical model. Through this model, the knowledge has been gained about the contribution of each component of the composite sandwich beams to its energy absorption capacity. Crupi et al. [14] performed the numerical investigation of aluminium honeycomb sandwiches subjected to low-velocity impact tests which were reported in their previous experimental study [15]. The finite element model was developed with the capture of geometric details of the honeycomb sandwich by using three-dimensional X-ray computed tomography (XCT). By comparing the finite element results with experimental data, the numerical model was also validated. Although numerous studies have been performed on the low-velocity impact of sandwich structures till now, there are few studies on the development of numerical realistic models for understanding low-velocity impact response of aluminium honeycomb sandwich structures. In this study, a numerical realistic model is developed in order to properly simulate the low-velocity impact behavior of aluminium
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honeycomb sandwich structures through the explicit finite element code LS-DYNAÕ .
Experimental procedure In order to validate the finite element model, an experimental investigation on aluminium sandwich plates subjected to low-velocity impact loadings was carried out. Tests specimens consisted of aluminium alloy Al 3003-H19 foils for the honeycomb core with aluminium alloy Al 2024-T3 for the facesheets. Each plate measured 80 mm 80 mm, with a core thickness of 18 mm and a thickness of 1 mm for each top and bottom facesheet. The thickness of honeycomb wall was equal to 0.07 mm, and the cell size was 6.35 mm. All sandwich plates were fabricated by bonding the facesheets to the core material with Araldite 2015 adhesive film, a room temperature curing epoxy resin. The bonded sandwich plates were cured at room temperature under low pressure for 72 h to achieve good adhesion. After curing process of sandwich plate was completed, 80 80 mm test specimens were prepared to low-velocity impact test (Figure 1). The experiments were repeated for each energy levels, and very good agreements were obtained between the tests. However, the figures are not shown to avoid increasing the number of figures. Impact tests were performed using the instrumented impact testing system, CEAST Fractovis, which consists of a drop tower equipped with an impactor and a variable cross-head weight arrangement, a highspeed data acquisition system, a spring-assisted velocity system, and a load transducer mounted in the impactor as shown in Figure 2(a). The impactor has a hemispherical nose of 20 mm in diameter and connected to a force transducer with a maximum loading capacity of 40 kN. The total impacting mass of impactor nose, force transducer, and crosshead were kept constant at 5.045 kg for all tests. The specimens were fixed in clamped condition by an annular support with an internal diameter of 40 mm by means of a pneumatic system during the lowvelocity impact event (Figure 2(b)). After initial contact, secondary impacts on the
Figure 1. An aluminium honeycomb sandwich plate.
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sample were prevented by means of pneumatic anti-rebounding system (Figure 2(c)). The impact energy was changed using the spring-assisted velocity system, which accelerates the crosshead to increase the maximum impact energy. The sandwich plates were tested at impact energies of 41.95, 59.78, and 81.07 J corresponding to impactor velocities of 4.078, 4.868, and 5.669 m/s, respectively.
Numerical model In order to design a monitoring system for the impact diagnosis and obtain a lowcost preliminary simulated experience, the use of numerical realistic models has become a necessity. Thus, experimental studies can be reduced to just a validation assessment. Low-velocity impact on aluminium honeycomb sandwich structure was modeled taking into account the actual geometry of the experimental sandwich plate using LS-DYNAÕ [16] explicit dynamic product. The top and bottom facesheets (Al 2024-T3) and the honeycomb core (Al 3003-H19) were assumed an elastoplastic material and the material behavior is based upon the piecewise linear plasticity material model. Damage is also coupled with this material law based on the failure strain. Figure 3 shows the true stress–strain curves of facesheet (Al 2024-T3) and honeycomb core (Al 3003-H19) materials, and the mechanical properties of them are listed in Table 1. Impactor and clamped rings were
Figure 2. (a) Drop weight test system, (b) impact test fixture system and specimen, and (c) anti-rebounding system.
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500
Stress (MPa)
400 300 200 100 0 0
Al 2024-T3 Al 3003-H19 0.025 0.05 0.075
0.1 0.125 0.15 0.175 Strain
0.2
Figure 3. The true stress–strain diagrams of facesheet Al-2024 T3 and honeycomb core Al3003 H19 materials.
Table 1. Material properties of sandwich structure components. Materials
Young’s modulus (GPa)
Poisson’s ratio
Density (kg/m3)
Yield stress (MPa)
Failure strain
Al 2024-T3 Al 3003-H19
73.1 70
0.33 0.33
2780 2730
350 183
0.62 0.52
considered as rigid. The top and bottom facesheets and honeycomb core were meshed with four-noded shell elements that use the Belytschko–Leviathan algorithm [16]; however, the impactor and clamping rings were meshed with eightnoded solid elements. Boundary conditions were applied with rigid clamping rings according to the nature of experiments. Namely, the motion of bottom ring was constrained in all directions, while the motion of impactor and top ring was only allowed in the z-direction. The rotations in all directions for impactor, top, and bottom rings were constrained. Moreover, compressive force that was measured experimentally from a load cell for top ring and has a value of 2403.45 N was applied to top ring with LOAD_SEGMENT option in LS-DYNAÕ . Contact modeling is critical for predicting the low-velocity impact response of sandwich structures. In this study, an AUTOMATIC_SINGLE_SURFACE contact algorithm was defined in order to take into account self-contacting interfaces during the projectile impact process. The contact between the projectile and sandwich structure was defined with ERODING_SURFACE_TO_SURFACE algorithm. The static and dynamic friction coefficients were set to 0.47 and 0.38, respectively. An AUTOMATIC_SURFACE_TO_SURFACE contact definition was defined between the clamping rings and sandwich panel with the static and dynamic friction coefficients of 0.47 and 0.38, respectively.
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Moreover, TIED_NODES_TO_SURFACE_TIEBREAK (TNTS) contact algorithm was defined between honeycomb core and aluminium facesheets to model separations, which may occur at the adherence interface due to the impact load. TNTS contact algorithm is based on failure forces (normal and shear failure forces) and failure of connection which will occur when m1 m2 fn fs þ 1 fn,fail fs,fail
ð1Þ
where fn is the normal force, fn,fail is the normal failure force, m1 is the normal force exponent, fs is the shear force, fs,fail is the shear failure force, and m2 is the shear force exponent. Aluminium facesheets and honeycomb core were bonded with epoxy adhesive (Araldite 2015). Normal and shear failure strengths of Araldite 2015 are given in Table 2 [17]. Normal and shear failure forces were calculated from normal and shear failure strengths and were applied at the adherence interface with equation (1). Increased mesh density for the sandwich plates showed practically no improvement in the finite element method solutions while significantly increasing computational expense. Therefore, the finite element mesh for the sandwich plate, impactor, and clamping rings was used as shown in Figure 4. All simulations were performed in a workstation with eight CPUs in 2.5 GHz and 32 GB RAM. In order to calibrate the numerical model, the sandwich plates were tested at impact energies of 41.95, 59.78, and 81.07 J corresponding to impactor velocities of 4.078, 4.868, and 5.669 m/s, respectively, in accordance with experiments. Later, the effects of the cell size and core height on the low-velocity impact response of the sandwich structure were investigated using calibrated and validated numerical model. For this purpose, sandwich structures were analyzed using three different cell sizes (3.175, 6.350, and 9.525 mm) for core height of 18 mm and three different core heights (10, 18, and 25 mm) for cell size of 6.350 mm, while the top and bottom facesheets thickness of 1 mm was kept constant through the analysis.
Results and discussion In order to validate the numerical model, the predicted results were compared with the experimental results. The contact–force history curves and the global
Table 2. Normal and shear failure strength of Araldite 2015 [17]. Normal failure strength, f (MPa)
21.63 1.61
Shear failure strength, f (MPa)
17.9 1.8
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Figure 4. Finite element models of the sandwich plate, impactor, and clamping rings.
deformations of the sandwich structures were used in the validation for the impact energies of 41.95, 59.78, and 81.07 J, while the cell size of 6.350 mm and the core height of 18 mm. Figure 5 shows the comparison between the experimental and numerical contact–force histories for the impact energies of 41.95, 59.78, and 81.07 J, respectively. As shown in the comparisons, there exists a good agreement between the predicted and experimental results in terms of overall trend for all impact energies. For increasing impact energy, the predicted peak contact–force values are 9.29 kN, 10.46 kN, and 11.49 kN, whereas the experimental peak contact–force values were obtained as 9.88 kN, 10.99 kN, and 11.57 kN. Thus, the differences between the predicted and experimental peak contact–force values are about 5.97%, 4.82%, and 0.69% for the impact energies of 41.95, 59.78, and 81.07 J, respectively. In addition, the differences between the predicted and experimental contact durations are about 0.61 ms, 0.50 ms, and 0.96 ms, respectively. After the impact tests, the predicted and experimental through-the-thickness impact damages of honeycomb sandwich plates are compared in Figure 6 for the impact energies of 41.95, 59.78, and 81.07 J. The predicted permanent displacements are 7.00 mm, 8.30 mm, and 10.70 mm, whereas the experimental permanent displacements were measured as 6.69 mm, 8.17 mm, and 10.21 mm for the impact energies of 41.95, 59.78, and 81.07 J, respectively. Accordingly, the obtained deviations between the numerical and experimental permanent displacements are less than 5% for all corresponding energy levels. The failure mode observed during the numerical and experimental tests was the core buckling that concentrated under
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Contact Force (kN)
(a) 10 8
Experimental LS-DYNA
6 4 2
0
1
2
3 4 Time (ms)
5
6
7
3 4 Time (ms)
5
6
7
3 4 Time (ms)
5
6
7
(b) 12
Contact Force (kN)
10
Experimental LS-DYNA
8 6 4 2 0
1
2
(c) 12
Contact Force (kN)
10
Experimental LS-DYNA
8 6 4 2 0
1
2
Figure 5. Comparison of experimental and numerical contact–force histories of aluminium honeycomb sandwich plate for the impact energies of (a) 41.95 J, (b) 59.78 J, and (c) 81.07 J (d ¼ 6.35 mm, h ¼ 18 mm).
the point of impact for all energy levels, which did not lead to perforation of the sandwich plates. As can be seen from Figure 6, the finite element model was able to predict correctly the deformed shape of the honeycomb sandwich plates after the impact tests.
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Figure 6. Comparison of experimental and numerical after-impact deformations of aluminium honeycomb sandwich plate for the impact energies of (a) 41.95 J, (b) 59.78 J, and (c) 81.07 J.
Kinetic Energy (J)
100 41.95 J 59.78 J 81.07 J
80 60 40 20 0 0
1
2
3 4 Time (ms)
5
6
7
Figure 7. Kinetic energy histories of the impactor.
By using the validated numerical model, the predicted kinetic energy–time and contact force–central displacement variations are shown in Figures 7 and 8, respectively, for the impact energies of 41.95, 59.78, and 81.07 J. The absorbed energy could be found in forms of damages on the specimen and heat dissipation. By neglecting small losses, the absorbed energy by the specimen can be accepted equal to the difference between the impact energy and the rebound energy of the impactor as follows 1 E ¼ EI ER ¼ mðv2I v2R Þ ð2Þ 2
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12
Contact Force (kN)
10
41.95 J 59.78 J 81.07 J
8 6 4 2 0 0
2
4 6 8 Displacement (mm)
10
12
Figure 8. The effect of impact energy on the contact force–central displacement variation of honeycomb sandwich plate.
Table 3. Absorbed energy rates by sandwich plate components for different values of impact energy, cell size, and core height.
Parameters Sandwich plate components Upper plate Honeycomb Lower plate Rebound energy of the impactor
Impact energy (J) (d ¼ 6.35 mm, h ¼ 18 mm)
Cell size (mm) (E ¼ 59.78 J, h ¼ 18 mm)
Core height (mm) (E ¼ 59.78 J, d ¼ 6.35 mm)
41.95
59.78
81.07
3.175
6.35
9.525
10
18
25
59.2% 32.9% 0.5%
58.6% 35% 0.5%
52.6% 42.8% 0.5%
63.4% 31.2% 0.8%
58.6% 35% 0.5%
14.5% 81.4% 3.2%
57.4% 35.5% 1.4%
58.6% 35% 0.5%
59% 35.4% 0.3%
7.4%
5.9%
4.1%
4.6%
5.9%
0.9%
5.7%
5.9%
5.3%
where E, EI, and ER, are the absorbed energy, the impact energy of the impactor, and the elastic energy due to rebound, respectively; vI, vR, and m are velocity of the impactor at impact (impact velocity), rebound velocity of the impactor (residual velocity), and the mass of the impactor, respectively. A large part of the impactor energy is absorbed by sandwich structure; thus, 92.6%, 94.1%, and 95.9% of the kinetic energy of the impactor were absorbed as plastic deformation by the sandwich plates for the impact energies of 41.95, 59.78, and 81.07 J, respectively. The residual kinetic energies of the impactor and the contact durations are approximately the same in all energy levels (Figure 7). It is evident that increasing impactor energy causes an increase in both the peak contact force and the central displacement on the sandwich plate (Figure 8). Table 3 shows the absorbed energy rates by sandwich plate components for different values of
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impact energy, cell size, and core height. The cell size of the honeycomb is very effective on the energy absorption of the structure. Thus, the absorbed energy by the honeycomb core increases with increasing cell size of the honeycomb. However, the effect of core height of the honeycomb has a minor effect on the energy absorption of the sandwich structure. This is because the after-impact central displacements and deformation geometries remain almost the same levels and the deformations have not reached the bottom facesheets. In thinner sandwich structures, the deformation after-impact event can reach the bottom facesheet; therefore, the low-velocity impact response of the thinner sandwich plate can be different from the thick ones. The effect of the cell size and the height variation of aluminum honeycomb core on the impact response of sandwich structures were investigated using the improved numerical model.
The effect of cell size In order to determine the effect of the cell size on the low-velocity impact response of the aluminium honeycomb sandwich structure, three different cell sizes (3.175, 6.350, and 9.525 mm) were used in the numerical analysis for the impact energy of 59.78 J, while the honeycomb core height is 18 mm. Figure 9 shows the after-impact central displacements and deformation geometries of the honeycomb sandwich plates for the cell sizes of 3.175, 6.350, and 9.525 mm, respectively. It is evident that the plastic deformation and central displacement value of the sandwich structures increase with increasing the cell size of the honeycomb. In addition, the buckling on the honeycomb core is observed to progress from center toward edges of the sandwich plate with increasing cell size. Especially, all cells of the sandwich plate having 9.525-mm cell size buckled and upper facesheet shifted toward down after the impact loading. Consequently, the stability of the structure is disturbed by increased cell size. The peak contact force increases with decreasing cell size of the honeycomb, while the contact time shortens (Figure 10). The peak contact forces are 12.96 kN, 10.46 kN, and 5.79 kN for the cell sizes of 3.175, 6.350, and 9.525 mm, respectively. It is obvious that the impact resistance of the sandwich structure decreases with increasing cell size. This is because the buckling occurs in the cell walls for the sandwich structure having honeycomb with larger cell size. Figure 11 shows the kinetic energy histories for the cell sizes of 3.175, 6.350, and 9.525 mm (E ¼ 59.78 J, h ¼ 18 mm). The absorbed energies are the same level, which are about 94.8% for the cell sizes of 3.175 and 6.350 mm. However, for the cell size of 9.525 mm, impact energy of 99.1% is absorbed as plastic deformation by the sandwich structure. As observed from Figure 11 that the contact duration decreases with decreasing cell size of the honeycomb. Figure 12 shows the contact force–central deflection variations during the low-velocity impact event for the aluminium honeycomb sandwich plate having cell sizes of 3.175, 6.350, and 9.525 mm (E ¼ 59.78 J, h ¼ 18 mm). It is evident that the stiffness of sandwich structure is increased with
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Figure 9. The effect of cell size on the after-impact central displacements of aluminium honeycomb sandwich plate.
14 3.175 mm 6.350 mm 9.525 mm
Contact Force (kN)
12 10 8 6 4 2 0 0
2
4 6 Time (ms)
8
10
Figure 10. The effect of cell size on the contact force histories of aluminium honeycomb sandwich plate (E ¼ 59.78 J, h ¼ 18 mm).
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Figure 11. The effect of cell size on the kinetic energy histories of the impactor (E ¼ 59.78 J, h ¼ 18 mm).
Figure 12. The effect of cell size on the contact force–central deflections of aluminium honeycomb sandwich plate (E ¼ 59.78 J, h ¼ 18 mm).
decreasing cell size of the honeycomb; hence, this situation causes decreased plastic deformation through-the-thickness of the sandwich structure.
The effect of core height To investigate the core height on the low-velocity impact response of the aluminium honeycomb sandwich structure, three different core heights (10 mm, 18 mm, and 25 mm) were used in the numerical analysis for the impact energy of 59.78 J, while the honeycomb cell size is 6.350 mm. As shown in Figure 13, the after-impact central displacements and deformation geometries remain almost the same. Thus, the variation in the core height of honeycomb has insignificant effect on
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Figure 13. The effect of cell height on the after-impact central displacements of aluminium honeycomb sandwich plate (E ¼ 59.78 J, d ¼ 6.35 mm).
the after-impact deformation of the sandwich structure. Figure 14 shows the effect of core height on the contact force histories, the kinetic energy histories, and the contact force–central deflections of aluminium honeycomb sandwich plate for the impact energy of 59.78 J, while the honeycomb cell size 6.350 mm. It can be seen from Figure 14(a, b) that the variation of contact–forces and kinetic energies are the same values for all honeycomb core heights. The peak contact force and the after-impact central displacement values are almost the same for three honeycomb core heights i.e. h ¼ 10 mm, 18 mm, and 25 mm (Figure 14(c)). As a result, the honeycomb core height is not critical variable in terms of energy absorption of the aluminium honeycomb sandwich plates subjected to low-velocity impact [8, 18].
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(a) 12
10 mm 18 mm 25 mm
Contact Force (kN)
10 8 6 4 2 0 0
1
2
3 Time (ms)
4
5
6
(b) 60
10 mm 18 mm 25 mm
Kinetic Energy (J)
50 40 30 20 10 0 0
1
2
3 Time (ms)
4
5
6
(c) 12
Contact Force (kN)
10
10 mm 18 mm 25 mm
8 6 4 2 0 0
2
4 6 Displacement (mm)
8
10
Figure 14. The effect of core height on (a) the contact force histories, (b) the kinetic energy histories, and (c) the contact force–central deflections of aluminium honeycomb sandwich plate (E ¼ 59.78 J, d ¼ 6.35 mm).
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Conclusions This study focused on developing of numerical realistic model for predicting lowvelocity impact response of aluminium honeycomb sandwich structures. In the experimental part of this study, a series of impact tests were conducted with sandwich specimens fabricated with aluminium honeycomb core and aluminium facesheets to improve a numerical realistic model of them. The obtained results, demonstrate good agreement, not only for the contact force histories and the peak forces, but also with regard to the detailed comparisons between numerical and experimental deformations. In the numerical part, it is demonstrated that fine micromechanical models based on shell elements give good correlation with specimen impact tests. After verifying the numerical model, the effect of the cell size and the height variation of aluminum honeycomb core on the impact behavior of sandwich structures were numerically investigated. The cell size of the honeycomb core has a significant effect on the impact behavior of the aluminium honeycomb sandwich plates. Thus, the stiffness and stability of sandwich structure are increased with decreasing cell size of the honeycomb. However, the core height of the honeycomb has a minor effect on the impact response of the sandwich plates. Eventually, the change in core height does not have any influence on the absorbed energy by the sandwich structure. Declaration of Conflicting Interests The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.
Funding The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: The authors would like to acknowledge funding from the Scientific and Technological Research Council of Turkey (TUBITAK) under the research Grant No. 112M917 and Coordination Unit of Scientific Research Projects of Erciyes University under the research Grant No. FYL-2013-4729.
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