CONCURRENT BENDING AND LOCALIZED IMPACT ON SANDWICH
PANELS
J.A. Kepler and P. Bull
Department of Mechanical Engineering
Aalborg University
Denmark
[email protected]
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 Mikkor et al. and [5] by Khalili et al.) whereas similar
combined loading on sandwich structures has apparently been largely neglected.
In [6], Malekzadeh 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.
a)
v2
c)
y
m
α
M
α
M
x
v1
b)
m
z
Fig. 1: A panel a), subjected to uniform cylindrical bending by moment M, 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 v1 (initial,
position b)) and v2 (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.
1
2
3
4
5
6
Fig. 2: Bending rig with specimen, laid out on table for clarity. The parts are:
1: Test specimen
2: Moment 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
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 bc d e
g
f
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: Mirror
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:
2
Lamina: UD prepreg, T700 carbon fibre with SE84LV epoxy resin, fibre mass 300 g/m , fibre volume fraction 60%, total mass
2
476 g/m , effective thickness 0.25 mm per lamina. Face-sheets: layup sequence [0°/90°/(core)/90°/0°], total effective thickness
tf = 0.50 mm per face-sheet (as verified by microscopy).
Core: Divinycell HP80, thickness 10 mm
Core end inserts: Mild 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
10
b)
10
15
a)
z
150
y
x
250
410
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
My = (F/2)⋅a
where
F: total transverse force [N]
-3
a: distance between outer and inner roller (a = 47.5⋅10 m)
Figure 6 shows the moment-strain development until failure.
Face-sheet strain
[ strain]
xx
Bending moment vs. strain
5000
4000
3000
2000
1000
0
0
100
200
300
400
Moment M xx [Nm]
Figure 6: Static bending test. The specimen failed symmetrically at a load of 300.5 Nm, and the maximum face sheet
strain was 4050 μstrain.
This corresponds to a nominal compressive failure stress of -364 MPa, 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.
z
y
M
x
M
Fig. 7: Outline of test specimen, subjected to uniform moment load M. 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
800
1600
2500
Impactor initial
velocity
v1 [m/s]
421
417
423
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
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 )
crushing line
x (0 )
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.
b)
d)
e)
y
z
M
a)
M
x
v
c)
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.
References
1.
Bull, P and Hallström, S., Journal of Sandwich Structures and Materials, Vol. 6, No 2, 2004.
2.
Kepler, J., “Localized Impact on Sandwich Structures – an Experimental Study”, Ph.D. thesis, Department of Mechanical
Engineering, Aalborg University, Denmark, ISSN 0905-2305, 2001.
3.
Rosenberg, Z., Mironi, 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. 827-831, 1994
4.
Mikkor, K. M., Thomson, R. S., Herszberg, I., Weller, T. and Mouritz, A. P., “Finite Element Modeling of Impact on
Preloaded Composite Panels”, Composite Structures, Vol. 75 p. 501-513, 2006
5.
Khalili, S. M. R., Mittal, R. K. and Mohammad Panah, N., “Analysis of Fiber Reinforced Composite Plates Subjected to
Transverse Impact in the Presence of Initial Stress”, Composite Structures, Vol. 77 p. 263-268, 2007
6.
Malekzadeh, K., Khalili, M. R. and Mittal, R. K., “Response of In-plane Linearly Prestressed Composite Sandwich Panels
with Transversely Flexible Core to Low-velocity Impact”, Journal of Sandwich Structures and Materials, Vol.8 p. 157-181,
2006