CONCURRENT BENDING AND LOCALIZED IMPACT ON SANDWICH PANELS

Transcription

1 CONCURRENT BENDING AND LOCALIZED IPACT ON SANDWICH PANELS J.A. Kepler and P. Bull Department of echanical Engineering Aalborg University Denmark ABSTRACT This paper describes impact/bending testing of sandwich panels, with special emphasis on test procedures, registration of results and description of the specialized test equipment developed. The sandwich panels used in these tests were composed of 10 mm PVC foam core and carbon fiber reinforced polymer (CFRP) face-sheets. The sandwich panels were subjected to a cylindrical bending load of varying magnitude, while impacted at approximately 420 m/s. The impactor body was a steel sphere, diameter 10mm, with a mass of 4.1 g. A high-speed camera was used for qualitative registration of panel response. Introduction Sandwich panels consists of two stiff, thin face-sheets separated by and glued to the surfaces of a lightweight core plate. In this manner, a bending moment is largely carried by membrane stresses in the face sheets, while the core plate transfers shear loads. This structural layout may be employed to produce panels with a particularly good ratio between stiffness/strength and mass for a bending load (in much the same way as an I-profile steel beam). Some complications may arise, however; sandwich structures are sensitive to imperfections in load application, in particular in the form of concentrated loads, and as fully stressed structures, there is little reserve capacity once one structural component has failed. Furthermore, the multitude of damage modes make prediction of damage morphology a complex matter. The subject of localized, penetrating impact on sandwich structures has been the focus of some attention in recent years, see e.g. [1] by Bull and [2] by Kepler. Typically, the primary topics of interest have been energy absorption by penetration, and mechanical properties after impact (residual stiffness/strength). While these matters may be addressed separately, they do not necessarily give any fair indication of what happens when a stressed sandwich panel is penetrated. In particular, the risk of sudden catastrophic deterioration initiated by a localized impact is worth investigating. Simultaneous localized impact and structural preload has been investigated for pressurized tubes (airplanes, oil pipes etc.), see e.g. [3] by Rosenberg et al., and monolithic composite structures under such load combinations has received increasing attention in recent years (for some recent advances, see e.g. [4] by ikkor et al. and [5] by Khalili et al.) whereas similar combined loading on sandwich structures has apparently been largely neglected. In [6], alekzadeh et al. describes a model for predicting the contact force and panel response when subjecting an in-plane prestressed sandwich panel to low velocity, nonpenetrating impact. It was, among other things, demonstrated that the peak contact force would increase and the deflection decrease with increasing tensile preload. These tendencies were qualitatively identical to those described by Khalili et al. in [5]. However, the matters of overall dynamic structural response and possible catastrophic failure following penetrating impact were not focal points of that investigation. In the present paper, a series of tests are described, where a panel under cylindrical bending was penetrated by a spherical impactor. Figure 1 shows an outline of the preload and penetration situation.

2 a) y c) m v 2 α α x b) z m v 1 Fig. 1: A panel a), subjected to uniform cylindrical bending by moment, causing a slope α at the ends, relative to the undeformed state. A rigid impactor, traveling along the z-axis, is characterized by mass m, geometry and velocities v 1 (initial, position b)) and v 2 (exit, position c)). Test equipment and procedures For applying the bending moment, a test rig was designed and manufactured. The bending rig is shown in figure 2, with a sandwich panel specimen inserted Fig. 2: Bending rig with specimen, laid out on table for clarity. The parts are: 1: Test specimen 2: oment yoke 3: Clamp bars 4: Load yoke 5: Load screw 6: Yoke supports A test specimen may, as shown in figure 2, be inserted between the clamp bars (3). By rotating the load screws (5), the load yokes (4) is moved inwards or outwards, hereby acting on the moment yokes (2) and bending the test specimen. Ball-joints

3 and edge bearings are employed to allow some freedom of deformation, as may e.g. be caused by bending-twisting coupling effects in the test specimen For the impact tests, the bending rig is fitted to a rigid steel frame, permitting transverse impact by a steel sphere propelled by a compressed-air gun. The test setup is shown in figure 3. a b c d e f g h i j k Fig. 3: Test setup, impact test with bending rig and high-speed camera, as seen from above. a: Target frame b: Impactor capture device c: High-speed camera d: Bending rig e: Test specimen f: Speed trap g: irror h: Impactor trajectory i: Blast shield j: Barrel tube k: Gun chamber Testing proceeds as follows: The bending rig (d) is mounted on the main target frame (a). An impactor is placed in the gun chamber (k, initially disconnected from the high-pressure compressor, not shown). The high-speed camera is placed in a steel box (c) with a small viewport. A specimen (e) is placed in the bending rig (d) and subjected to the desired bending preload. Safety gates (not shown) are closed, and the high-pressure compressor is connected to the gun chamber. The gun chamber (k) is pressurized to the desired pressure (approximately 100 atm. in these tests). Finally, two 650 W photo-lamps are turned on, whereupon the gun is fired. The impactor travels through the barrel tube (j) along the trajectory line (h). The blast shield (i) deflects most of the air jet following the impactor. The speed trap (f) records the incident velocity of the impactor body. After penetration, the impactor is captured in the capture device (b). The high-speed camera is trigged manually. After opening the safety gates, the gun chamber is partially dismantled to ensure that it cannot hold any pressure, whereupon the specimen may be retrieved. The high speed camera (Olympus i-speed 2) settings were: Recording speed: 8000 frames/second Frame size: 256 x 192 pixels Shutter (exposure time): 63 μs Test specimens materials and geometry The test specimens were fabricated from carbon fibre / epoxy prepreg, with PVC foam core. Full width steel core inserts were used in the regions of load application. The materials used were: Lamina: UD prepreg, T700 carbon fibre with SE84LV epoxy resin, fibre mass 300 g/m 2, fibre volume fraction 60%, total mass 476 g/m 2, effective thickness 0.25 mm per lamina. Face-sheets: layup sequence [0 /90 /(core)/90 /0 ], total effective thickness t f = 0.50 mm per face-sheet (as verified by microscopy). Core: Divinycell HP80, thickness 10 mm Core end inserts: ild steel, thickness 10 mm The effective face-sheet thickness in the finished specimens was somewhat less than the nominal thickness (0.56 mm, according to the supplier). The most likely reason is that some of the resin bled into the core surface cavities and the breather cloth during vacuum curing. The specimens were fabricated in three stages: 1: Steel end inserts glued to the core, using epoxy glue. 2: Primary face-sheet prepregs added. The specimen was laid up with the compression face-sheet against a plane surface, to avoid excessive imperfections which might cause premature face-sheet buckling. 3: Reinforcement prepregs in the steel core / foam core transition zone added in a separate cure process. The prepreg formulation permitted curing without external pressure (only vacuum and elevated temperature). The final geometry is indicated in figure 4

4 10 b) a) y z x Fig. 4: Test specimen geometry. All dimensions in mm. The specimen is symmetric about the yz-plane. The magnified view shows the local reinforcement of the compressed face sheet ([0 /90 /0 /90 ] dropping to [90 /0 /90 ] dropping to [0 /90 ]) in the transition zone between steel core (b) and foam core (a). Static bending experiment One test specimen was subjected to a static 4-point (or 4-line, being a plate specimen) bending test. The purpose of the static test was to determine the maximum feasible structural load on the specimen. The line supports were rigid steel rollers, and the force was distributed to the upper rollers via a balance beam, as shown in figure 5. Fig. 5: 4-line static bending test, showing lower rollers (distance 255 mm), upper rollers (distance 350 mm) and balance beam. The lower-roller distance was set to 255 mm to ensure that the transverse force was transmitted to the steel core-inserts. From the indicated force, the uniform moment load was calculated as y = (F/2) a

5 where F: total transverse force [N] a: distance between outer and inner roller (a = m) Figure 6 shows the moment-strain development until failure. Bending moment vs. strain Face-sheet strain xx [ strain] oment xx [Nm] Figure 6: Static bending test. The specimen failed symmetrically at a load of Nm, and the maximum face sheet strain was 4050 μstrain. This corresponds to a nominal compressive failure stress of -364 Pa, as calculated from the failure moment (300.5 Nm). Considering the thin face-sheets and the layup sequence, this value is not unrealistically poor, although somewhat less than expected. Figure 7 shows the failure locations and magnified views of the failure zone in the compressive-stress face-sheet.

6 z y x Fig. 7: Outline of test specimen, subjected to uniform moment load. The magnifications show the failure of the compressive-stress face-sheet near the reinforcement ply drop-off. The failure mode shown in figure 7 indicates a combination of local compressive stress and shear stress causing the failure in the face-sheet. This can be caused by localized face-sheet buckling, which may in turn be initiated by the shift of the centre line of stiffness as the reinforcement plies are dropped off. Another contributing reason for the relatively moderate failure load may be found in the fibre quality of the face-sheet prepregs. Images from both optical and electron microscopy reveal a kidney-shaped fibre cross-section, which may be detrimental to the mechanical properties. For the impact tests, the maximum face sheet strain was set at approximately 60% of the ultimate strain in the static test. Bending and impact experiments Three specimens, similar to the one used in the static experiment, were prepared. A strain gauge was applied to the tensilestress face-sheet of each to provide an independent measure of the tensile strain. The specimens were placed in the bending rig shown in figure 2, whereupon bending load was applied to the specimen, prior to ballistic penetration by a spherical steel impactor, diameter 10 mm, mass 4.1 gram. Table 1 outlines the bending strains and impact velocities of the three specimens. Specimen Face-sheet strain εxx [μstrain] a b c Impactor initial velocity v1 [m/s] Table 1: Primary test parameters, combined bending and impact Catastrophic failure occurred in specimens b and c, while specimen a retained a significant load carrying capacity. Visual inspection of back face-sheet after test

7 Post-impact analysis of the specimens show that catastrophic damage, i.e. near-total loss of load carrying capacity, is initiated by failure in the compression face-sheet, as a combination of localized convex buckling and residual crushing. Figure 8 shows the primary damage morphology of specimens a, b and c. a) face-core separation 0 /90 separation y (90 ) x (0 ) crushing line b) y (90 ) x (0 ) c) y (90 ) x (0 ) Fig. 8: Post-impact study of back face-sheet damages. a: 800 μstrain preload b: 1600 μstrain preload c: 2500 μstrain preload It is noted that specimens b and c do not fail symmetrically about a horizontal axis. This may be due to manufacturing quality or uneven bending. The following chain of reasoning is consistent with the observed phenomena: 1: The impactor punches through the specimen, causing localized face/core debonding before penetrating the back facesheet. 2: If the strain energy density is sufficient, the buckling induced by penetration may extend further across the specimen (as witnessed in specimens b and c). The load carrying capacity in the buckled region is reduced to a negligible fraction, and the remaining unbuckled part of the compressive face-sheet must carry the full load. 3: Eventually, the residual load-carrying capacity of the compressive face-sheet is exceeded, and final failure occurs (typically through a combination of microbuckling and crushing). This speculative failure pattern is outlined in figure 9.

8 z y a) b) x v c) d) e) Fig. 9: Failure progression. Upper left corner: Specimen, seen from the back side. v indicates the impactor velocity. a) to e): magnified view of the central area at different stages of penetration. The impactor is represented as a sphere with dashed outline (when hidden) or full outline (when exposed). a: onset of debonding between back face-sheet and core. b: maximum debonding size, beginning face-sheet failure c: tearoff of central strip of outer (0 ) ply d: delamination front progressing in y-direction e: crushing failure extending across the remaining width of the specimen For specimen a (preload strain 800 μstrain), the final stage is stage c in figure 9. For specimens b and c (preload strain 1600 and 2500 μstrain respectively), the failure progression continues through stages d and e in figure 9. Conclusions The present test series consisted of merely 4 specimens one for bending-only, and three for combined bending and penetration. Quantitative conclusions would be premature on this basis, but the following may be stated: It was demonstrated that localized, penetrating impact on a convexly preloaded sandwich panel may initiate catastrophic damage. It is speculated that the prime cause of final loss of load carrying capacity is propagation of the debonding between the back face-sheet and the core. It is furthermore assumed that the debonding is initiated as the impactor penetrates and partially punches off the compressive-stress back face-sheet, whereupon the debonding front propagates according to fracture-mechanics, i.e. release of structural bending energy. It was noted that the specimen with the least bending preload did not fail catastrophically this indicates a threshold preload level. Consequentially, an increased bonding strength/toughness between the back (compressed) face-sheet and the core may be beneficial to the damage tolerance of the panel. Alternatively, through-the-thickness stitching may be employed. In an earlier test, where the direction of bending had instead been convex (back face-sheet in tension), the damage progression described above did not occur. This supports the assumption that the debonding of the compressive face-sheet from the core facilitates the final failure of the specimen. As indicated in figure 3, the high-speed camera viewed the front side of the specimen. However, the primary mechanisms of failure occurred on the back side of the specimen. In order to verify the assumed damage progression, the high-speed camera should be placed so as to view the back side of the specimen. This will require some offset of the bending rig from the target frame. Acknowledgments The authors gratefully acknowledge the contributions from the Danish research and development consortium Komposand.

9 References 1. Bull, P and Hallström, S., Journal of Sandwich Structures and aterials, Vol. 6, No 2, Kepler, J., Localized Impact on Sandwich Structures an Experimental Study, Ph.D. thesis, Department of echanical Engineering, Aalborg University, Denmark, ISSN , Rosenberg, Z., ironi, J., Cohen, A. and Levy, P., On the Catastrophic Failure of High-pressure Vessels by Projectile Impact, International Journal of Impact Engineering, Vol. 15 p , ikkor, K.., Thomson, R. S., Herszberg, I., Weller, T. and ouritz, A. P., Finite Element odeling of Impact on Preloaded Composite Panels, Composite Structures, Vol. 75 p , Khalili, S.. R., ittal, R. K. and ohammad Panah, N., Analysis of Fiber Reinforced Composite Plates Subjected to Transverse Impact in the Presence of Initial Stress, Composite Structures, Vol. 77 p , alekzadeh, K., Khalili,. R. and ittal, R. K., Response of In-plane Linearly Prestressed Composite Sandwich Panels with Transversely Flexible Core to Low-velocity Impact, Journal of Sandwich Structures and aterials, Vol.8 p , 2006