Materials and Design 31 (2010) 1887–1899

Contents lists available at ScienceDirect

Materials and Design

journal homepage: www.elsevier .com/locate /matdes

The influence of core height and face plate thickness on the responseof honeycomb sandwich panels subjected to blast loading

Y. Chi, G.S. Langdon *, G.N. NurickBlast Impact and Survivability Research Unit (BISRU), Department of Mechanical Engineering, University of Cape Town, Private Box, Rondebosch 7701, South Africa

a r t i c l e i n f o a b s t r a c t

Article history:Received 15 September 2009Accepted 29 October 2009Available online 31 October 2009

Keywords:HoneycombSandwichPanelsBlast loading

0261-3069/$ - see front matter � 2009 Elsevier Ltd. Adoi:10.1016/j.matdes.2009.10.058

* Corresponding author.E-mail address: genevieve.langdon@uct.ac.za (G.S.

This paper reports on an investigation into the behaviour of circular sandwich panels with aluminiumhoneycomb cores subjected to air blast loading. Explosive tests were performed on sandwich panels con-sisting of mild steel face plates and aluminium honeycomb cores. The loading was generated by detonat-ing plastic explosives at a pre-determined stand-off distance. Core height and face plate thickness werevaried and the results are compared with previous experiments. It was observed that the panels exhibitedpermanent face plate deflection and tearing, and the honeycomb core exhibited crushing and densifica-tion. It was found that increasing the core thickness delayed the onset of core densification and decreasedback plate deflection. Increasing the plate thickness was also found to decrease back plate deflection,although the panels then had a substantially higher overall mass.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

Sandwich panels, which consist of face plates separated by acore, are finding increasing use in structural applications becauseof their high specific stiffness and strength. While the responseof monolithic structures subjected to blast loading has been stud-ied for many years [1–6], it has only be in recent years thatresearchers have begun to focus on the blast response of sandwichpanels [7–12]. Fleck and Deshpande [7] developed analytical for-mulae to characterise the structural response of clamped metallicsandwich beams subjected to uniformly distributed air and waterblasting loads. The study reviewed and compared the blast perfor-mance of a monolithic plate and various core topologies. Qui et al.[8] extended the analytical model [7] for clamped circular sand-wich plates. In addition, Xue and Hutchinson [9,10] performed fi-nite element calculations to compare the blast response of amonolithic plate and a metal sandwich plate, of the same materialand total mass. In addition to these studies, Zhu et al. [11] pre-sented a review on impact and blasting of metallic and sandwichstructures, and Kim Yuen et al. [12] presented an overview onsandwich panels subjected to blast loading. A variety of core topol-ogies were investigated in these studies and the results haveshown that metal sandwich panels have the potential to performbetter than monolithic plates under certain impact and blastsituations.

Metal honeycombs have been used as cores for blast, impactand quasi-statically loaded sandwich panels since they are light-

ll rights reserved.

Langdon).

weight, have good energy absorption properties and exhibit highstrains to failure [13–19]. Dharmasena et al. [20] and Zhu et al.[21,22] have studied the air-blast response of square metal honey-comb sandwich panels. Dharmasena et al. [20] identified threestages of the response of the sandwich components to blast load-ing. It was found that the sandwich panel had lower back platedeflection than the solid plate. However, after complete core crush-ing, the advantages of the sandwich panel (as compared to the so-lid plate) are diminished. Furthermore, Zhu et al. [22] investigatedthe effect of foil thickness, cell size, mass of charge, relative densityof the core, and the face-sheet thickness. Zhu and Lu [22] foundthat there is a compromise between strength and weight. The re-ported results [22] focused on the back plate response and showedthat increasing the face-sheet thickness increased the mass of thepanel but decreased the back plate deflection. Increasing the foilthickness and increasing the honeycomb relative density both de-creased the back plate deflection. Conversely, increasing the cellsize increased the back plate deflection; this effect was especiallynoteworthy for panels with thinner face-sheet [22].

Despite a wealth of literature on the response of sandwichstructures with various core topologies, and on the quasi-staticand impact loading of honeycomb sandwich panels, there are rel-atively few studies that investigate the air-blast response of circu-lar aluminium honeycomb sandwich panels [23–26]. Nurick et al.[25] investigated the behaviour of a particular configuration of cir-cular aluminium honeycomb sandwich panels subjected to intenseair blasts. The face plates were made of 1.6 mm thick low carbonsteel and the core was an aluminium honeycomb of thickness13 mm. This paper presents experimental results for other config-urations of fully clamped circular sandwich panels with aluminium

Fig. 1. A typical engineering stress–strain graph showing the compression stages ofa honeycomb [13].

Fig. 2. Hexagonal honeycomb cell geometry.

1888 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

honeycomb core subjected to blast loads detonated at a pre-deter-mined stand-off distance. The influence of plate thickness and coreheight were studied and compared with the results reported in Ref.[25].

Table 1A summary of the honeycomb geometric information.

Core thickness(mm)

Cell size,S (mm)

Single walllength, l (mm)

Double walllength, c (mm)

Single wathickness

13 5.69 4 2.75 0.0829 6.43 3.83 3.39 0.07

150 10.22 6.6 4.5 0.11

Table 2Results of aluminium honeycomb specimens subjected to quasi-static compression.

Honeycombthickness (mm)

Peak stress(MPa)

Plateau stress(MPa)

Ons(% s

13 4.48 2.54 6829 4.21 1.73 72

150 5.67 2.23 74

1.1. Failure mode definitions for honeycomb sandwich panels

The failure modes of the mild steel face plates were expected tobe similar to those observed by Menkes and Opat for beams [1] andNurick and co-worker for plates [2]. This was also confirmed by theresults from tests on sandwich panels described in [25]. The fol-lowing failure modes are used to describe the failure of the frontand back face plates: large inelastic deformation (Mode I), partialtearing of the boundary (Mode II*), and tensile tearing around thefull boundary (Mode II). The failure of the core is described interms of compression and crushing of the honeycomb cellularstructure.

2. Material characterisation

The honeycomb core exhibits similar compression stages duringblasting to those observed in quasi-static axial compression tests. Atypical stress–strain curve of a honeycomb panel subjected to qua-si-static axial compression is shown in Fig. 1. When the honey-comb is loaded, a large proportion of the energy is absorbed bythe plastic buckling of the cells. At the same time, the stress trans-fer is limited by the characteristics of the stress–strain curve,which is usually described in terms of a relatively constant plateaustress [13] over a wide range of plastic strains (approximately 70–75%) as indicated in Fig. 1. The combination of the two character-istics (i.e. energy absorption and limited stress transfer) makesthe honeycomb a potential core material for sandwich panels usedto provide resistance against impact or blast loading. After thisstage of limited stress transfer, with an increase in strain, there isa steeply rising stress. This is called the densification region, asshown in Fig. 1, which starts at approximately 70–75% strain,regardless of the core thickness [13,15].

In this study, hexagonal aluminium honeycombs of three coreheights were used: 13 mm, 29 mm and 150 mm. The dimensionsof the honeycombs are defined in Fig. 2. A summary of the honey-comb geometric information is given in Table 1. The honeycombswere made of aluminium alloy AA3003, which has low rate sensi-tivity. Quasi-static compression tests were performed at a cross-head speed of 1 mm/min to characterise the core material. Thespecimens were kept at the same dimensions as those used inthe blast tests (i.e. 106 mm in diameter). The strain rates involvedin the blast tests were unknown so crushing at higher strain rateswas not attempted, although it is recognised that the energyabsorption obtained from the quasi-static crush tests will providea lower bound. The results of the quasi-static compression are

ll, h (mm)

Double wallthickness, 2h (mm)

Branchangle, a (�)

Relativedensity (%)

Density(kg/m3)

0.18 90 3.59 96.210.16 106 3.09 82.810.25 96 2.95 79.06

et of densificationtrain)

Energyabsorbed (MJ)

Energy absorbed perunit volume (MJ/m3)

171 1.52318 1.25

2193 1.66

Fig. 3. Graph of engineering stress–strain curves of compressed aluminium honeycombs of different thickness.

(b)

(a)

Fig. 4. Schematic of sandwich panel loading arrangement. (a) Core thickness13 mm (test series S13) and 29 mm (test series S29 and S29-1). (b) Core thickness150 mm (test series S150).

Table 3A summary of blast tests performed.

Honeycomb core thickness (mm)

13 29 150

Plate thickness (mm)1.0 S29-1 (9)1.6 S13 (16) S29 (25) S150 (7)

Experiment information Previous [25] Current investigation

Note: The numbers of experiments performed for each test series are shown in thebrackets.

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1889

summarised in Table 2. Typical engineering stress–strain curvesare shown in Fig. 3 for each honeycomb type.

3. Blast testing

The test specimens were constructed using 244 mm by 244 mmsquare, mild steel plates, having an exposed area of 106 mm diam-eter. Core thicknesses of 29 mm and 150 mm were sandwiched be-tween 1.6 mm thick faceplates. Tests were performed on these twoconfigurations (referred to as series S29 and S150, respectively).Blast tests were also performed on panels with 1 mm thick faceplates and 29 mm thick cores (referred to as series S29-1). TheseS29-1 panels were 34% lighter than the S29 panels (159 g versus242 g approximately). These results were compared to those avail-able in [25] for similar panels with 13 mm thick cores and 1.6 mmthick plates (referred to as series S13). The schematics of the exper-imental arrangements of these sandwich panels are shown inFig. 4. A test matrix is shown in Table 3. A total of 41 experimentswere performed for this study. The number of experiments per-formed for each test series is indicated in the brackets.

The face plates were not adhered to the honeycomb, so that theenergy absorption due to debonding of the components could bediscounted. Very little testing has been performed on unbondedsandwich panels, so it has not been possible to determine thechange in performance resulting from the bonding. It is anticipatedthat bonded panels would outperform unbonded ones, as the faceplates may be able to sustain larger stretching. However, it is alsopossible that the transfer of membrane stretching forces from theface plates to the core (via the bond-line) may result in prematurefailure of the core as honeycomb cores are significantly stronger in

Fig. 5. Photograph of the blast experimental arrangement.

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50

Impulse (Ns)

Def

lect

ion

(mm

)

S13 Front S13 Back S13 Front S13 Back

Onset of core densification Modes * and failure

Fig. 6. Graph of face plate deflection versus impulse for blast tested sandwich panels with a 13 mm thick core (S13), data from [25].

1890 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

through-thickness compression than under other loading condi-tions. A full treatment of the effect of adhesion is beyond the scopeof this paper, but it is anticipated that future work may examinethis topic. Herein, the study is restricted to unbonded panels.

The panels were clamped in a frame with a hole of 106 mmdiameter in the centre. A 150 mm long tube with the same internaldiameter as the clamping frame was attached to the front of theclamping rig. The PE4 was moulded into a 34 mm diameter flatcylindrical disc and sited at the open end of the tube. The detonatorwas attached to the centre of the disc using 1 g of explosive, re-ferred to as a ‘leader’. The tube was employed between the explo-sive and target plate to improve the spatial uniformity of the blastloading incident on the target plate, as described by Jacob et al.[27]. A photograph of the complete experimental arrangement onthe ballistic pendulum is shown in Fig. 5.

The experimental procedure was similar to that used by Nuricket al. [6,13]. The blast loading is created from the pressure wavegenerated by the explosion, which is produced by detonating thePE4. The impulse was obtained using the measured swing of thependulum. All of the S29 and S29-1 panels were mounted ontothe pendulum. Due to space constraints, the S150 panels werenot mounted, but were tested in a free-standing position and the

impulse estimated using an impulse–charge mass relationshipestablished from the other test work.

4. Results and analysis

In all test series, it was observed that the face plates behavedsimilarly to a clamped monolithic circular mild steel plate sub-jected to uniformly distributed blast loading [2,23]. It was pre-dicted using the experimental results reported by Teeling-Smithand Nurick [2] that a 1.6 mm thick, single clamped circular platesubjected to uniformly distributed blast loading will have a maxi-mum mid-point deflection of 30 mm prior to the initiation ofboundary tearing. This prediction was used as a starting point foranalysing the panel response, as it suggests that for sandwich pan-els with a core thickness lower than 30 mm, contact between thefront plate and back plate would occur prior to front plate tearing.

In all test series, the front plate of the sandwich panel exhibitedlower deflections than the single plate because of the structuralsupport provided by the sandwich core. The back plate alsoexhibited lower deflections, since the force magnitude and hencedeflections of the back plates are regulated by the core. However,

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1891

after the onset of core densification, higher stresses are transferredto the back plate which exhibited steeply increasing deflections.

The test series are examined individually in the following sec-tions. To fully explain the effect of core densification, analysis isperformed in the order of test series S13, S150 and S29. SeriesS13 shows the effect of densification and eliminates front platetearing until much higher impulse. Series S150 shows the effectsof front plate tearing with no core densification. Series S29 exhibitsboth the front plate tearing and densification phenomena. Obser-vations from series S13 and S150 are used to interpret the behav-iour of the S29 panels. S29-1 is compared to S29 to describe theinfluence of face plate thickness on panel response.

4.1. Test series S13

These tests results were obtained from work reported by Nuricket al. [25] using the same test procedure implemented herein.Graphs of face plate deflection verses impulse and core crushingverses impulse are shown in Figs. 6 and 7, respectively. The onsetof core densification obtained from quasi-static compression test issuperimposed on both graphs.

0

5

10

15

20

25

30

0 5 10 15 20 2

Impulse (N

Cru

sh d

ista

nce

(mm

)

Centre MidCentre Mid

Onset of core densification

Fig. 7. Graph of core crush distance versus imp

Fig. 8. Photographs of the face plat

The core exhibited complete densification at 23.1 N s, as shownin Fig. 7. Due to the relatively low core thickness (compared to themaximum 30 mm prediction for tearing of a 1.6 mm thick mono-lithic steel plate [2]), the front plate deflection is obstructed bythe back plate, as visible in Fig. 8 for an impulse of 30.3 N s. Withfurther increases in the applied impulse, the front plate was foundto exhibit tearing at 32.1 N s, as shown in Fig. 6. The obstruction re-duced the front plate deflection when compared to a single plate ofthe same thickness subjected to the same blasting.

Observations have also been made on the changes in the gradi-ent of the crushing of the core. At approximately 23 N s the centre ofthe core experiences a gradient decrease, as shown in Fig. 7. This isbecause the centre of the core reaches the maximum crushing limitand no further crushing is possible with increasing impulse. Fur-thermore, at approximately 31 N s, the outer-span of the core expe-riences an increase in gradient. This is due to the front plate tearing.

4.2. Test series S150

The results from blast tests on panels with a 150 mm core thick-ness are shown in Table 4. The face plate deflection–impulse graph

5 30 35 40 45 50

s)

-span Outer-span-span Outer-span

ulse of S13, replotted using data from [25].

es cross-section from S13 [25].

Table 4Blast results of blast tested sandwich panels with a 150 mm honeycomb core and 1.6 mm face plates.

Test number Impulse (N s) Plate mid-point deflection Plate failure mode

Front plate Back plate Front plate Back plate

(mm) d/H (mm) d/H

S150M34 40.45 20.23 12.64 4.62 2.89 II IS150M24 36.45 21.04 13.15 4.38 2.74 II IS150M20 33.17 25.01 15.63 3.52 2.20 II IS150R18 31.17 23.18 14.49 3.89 2.43 II IS150M18 31.17 23.00 14.38 1.70 1.06 I (thinning) IS150M16 28.93 20.70 12.94 1.76 1.10 I (thinning) IS150M12 23.73 15.94 9.96 1.13 0.71 I I

1892 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

is shown in Fig. 9. The core crushing–impulse graph is shown inFig. 10, with the onset of core densification obtained from quasi-static compression data superimposed on the graph.

The front plate deflection increased with increasing impulse upto approximately 31 N s, thereafter tearing occurred and a smalldecrease in deflection is observed with increasing impulse, asshown in Fig. 9. The onset of core densification in the quasi-staticaxial compression tests occurred at 112 mm crush distance (74%strain). From Fig. 10, it is evident that none of the S150 tests pro-duced densification of the core for impulses up to 40.5 N s. A sharp

0

5

10

15

20

25

30

0 5 10 15 20

Impu

Def

lect

ion

(mm

)

S150 Front S150 B

Two points overlap, one is Modeother is Mode

Fig. 9. Graph of face plate deflec

0

30

60

90

120

150

0 5 10 15 20 2

Impuls

Cru

sh d

ista

nce

(mm

)

CentreCentre

Onset of core densification

Fig. 10. Graph of core crushing dis

increase in the core crushing gradients is observed at 31 N s, asshown in Fig. 10. Photographs of core crushing for impulses ofand greater than 31 N s are shown in Fig. 11. Because no densifica-tion takes place, the only factor that influenced the gradient changeis the front plate tearing. The torn front plate has kinetic energywhich causes further crushing of the core. Higher impulses causethe residual velocity of the torn fragment to increase [28] which in-creases the percentage crush. The core crushing increased at a fas-ter rate with increasing impulse after tearing. The load transferredto the back plate is regulated by the core. In this case, no core den-

25 30 35 40 45 50

lse (Ns)ack S150 Front S150 Back

Mode failure (thinning) and the

tion versus impulse of S150.

5 30 35 40 45 50

e (Ns)Mid-span Outer-spanMid-span Outer-span

tance versus impulse of S150.

Fig. 11. Photographs of the honeycomb cores from S150.

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1893

sification occurs therefore no dramatic increase in the back platedeflection gradient is observed, since the magnitude of stresstransfer through the core is limited to the plateau stress. In alltests, back plate deflections did not exceed 5 mm.

4.3. Test series S29

The experimental results for test series S29 are listed in Table 5.The face plate deflection–impulse graph and the core crushing–im-pulse graph are shown in Figs. 12 and 13, respectively. The onset ofcore densification obtained from the quasi-static compression testsis superimposed on both figures. The onset of core densification oc-curred at approximately 30 N s, as shown in Fig. 13. This allowedhigher stresses to be transferred to the back plate, and the backplate deflection increases rapidly, as shown in Fig. 12. At approxi-

Table 5Blast results of blast tested sandwich panels with a 29 mm honeycomb core and 1.6 mm

Test number Impulse (N s) Plate mid-point deflection

Front plate

(mm) d/H

S29M24 36.75 18.83 11.77S29M22 36.79 23.53 14.71S29M20 34.36 23.31 14.57S29M18 31.30 26.35 16.47S29RM17 31.16 23.45 14.66S29RM16 31.69 23.22 14.51S29M16 29.15 22.52 14.08S29RM15.5 29.73 23.45 14.65S29M15 28.06 21.03 13.14S29RM15 25.48 21.23 13.27S29M14 27.77 19.89 12.43S29RM14 24.85 20.67 12.92S29M12 25.50 18.23 11.39S29RM12 25.37 19.00 11.88S29M10 21.50 16.58 10.36S29RM10 19.85 15.60 9.75S29RM09 19.59 14.80 9.25S29M09 18.97 15.42 9.64S29M08 17.55 12.18 7.61S29RM07 15.41 12.08 7.55S29M06 14.90 8.59 5.37S29RM05 10.60 8.37 5.23S29M04 11.31 5.80 3.63S29M03 8.70 4.13 2.58S29M02 5.34 0.42 0.26

mately 31 N s, the front plate is torn (as shown in Fig. 12) and thecore crushing increased at a faster rate (as shown in Fig. 13). This isbecause the torn front plate has kinetic energy which causes fur-ther core crushing. As a result, after the front plate tearing, a gra-dient increase is observed in the crushing of the mid-span andouter-span of the core, as shown in Fig. 13.

4.4. Test series S29-1

The experimental results of test series S29-1 are shown in Table6. These panels had thinner (1 mm) face plates. The graph of faceplate deflection–impulse and the graph of core crushing–impulseare shown in Figs. 14 and 15, respectively. The onset of core densi-fication is superimposed on both graphs.

The graphs in this test series have similar shapes and stages tothose in test series S29; however the thinner face plates have lower

face plates.

Plate failure mode

Back plate Front plate Back plate

(mm) d/H

21.47 13.42 II II*

12.65 7.91 II I12.22 7.64 II I

3.88 2.43 II* I3.25 2.03 I (thinning) I3.10 1.94 I (thinning) I2.73 1.71 I (thinning) I3.31 2.07 I (thinning) I2.66 1.66 I (thinning) I3.18 1.99 I (thinning) I2.56 1.60 I I2.97 1.85 I (thinning) I2.45 1.53 I I2.91 1.82 I I2.37 1.48 I I2.41 1.51 I I2.40 1.50 I I2.25 1.41 I I1.96 1.23 I I1.95 1.22 I I1.70 1.06 I I1.56 0.98 I I1.56 0.98 I I1.10 0.69 I I0.50 0.31 I I

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50

Impulse (Ns)

Def

lect

ion

(mm

)

S29 Front S29 Back S29 Front S29 Back

Onset of core densification Modes and failure

Fig. 12. Graph of face plate mid-point deflection versus impulse of S29.

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50

Impulse (Ns)

Cru

sh d

ista

nce

(mm

)

Centre Mid-span Outer-spanCentre Mid-span Outer-span

Onset of core densification

Fig. 13. Graph of honeycomb crush distance versus impulse of S29.

Table 6Blast results of blast tested sandwich panels with a 29 mm honeycomb core and 1 mm face plates.

Test number Impulse (N s) Plate mid-point deflection Plate failure mode

Front plate Back plate Front plate Back plate

(mm) d/H (mm) d/H

S29-1M14 25.59 24.41 24.41 20.54 20.54 II II*

S29-1M13 24.93 21.25 21.25 14.40 14.40 II IS29-1M12 23.70 25.95 25.95 5.83 5.83 II* IS29-1M10 20.86 22.88 22.88 5.19 5.19 I (thinning) IS29-1M08 18.56 19.83 19.83 4.70 4.70 I (thinning) IS29-1M06 14.18 15.81 15.81 4.25 4.25 I IS29-1M04 9.76 10.51 10.51 3.60 3.60 I IS29-1M03 7.57 6.53 6.53 3.19 3.19 I IS29-1M02 5.93 2.88 2.88 2.38 2.38 I I

1894 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

tearing threshold impulses. The gradient increases in the crushing ofthe mid-span and outer-span of the core (shown in Fig. 15) are due tothe front plate tearing at approximately 24 N s. In addition, the gra-

dient increase in the back plate deflection (shown in Fig. 14) is due tothe onset of densification at the centre region of the core. The steeprise in the back plate deflection can also be seen in Fig. 16.

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50

Impulse (Ns)

Def

lect

ion

(mm

)

S29-1 Front S29-1 Back S29-1 Front S29-1 Back

Onset of core densificationModes and failure

Fig. 14. Graph of face plate mid-point deflection versus impulse of S29-1.

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50

Impulse (Ns)

Cru

sh d

ista

nce

(mm

)

Centre Mid-span Outer-spanCentre Mid-span Outer-span

Onset of core densification

Fig. 15. Graph of honeycomb crush distance versus impulse of S29-1.

A dramatic increase in the back plate deflection (8.57mm) due to the onset of core densification at the centre region

Fig. 16. Photograph of plate deflection profile of the back plates from S29-1.

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1895

5. Discussion

5.1. Effect of core thickness

The effect of core thickness is studied by comparing test serieswhich used the same face plate thickness but different core thick-nesses, namely S13, S29 and S150.

Jacob et al. [27] proposed a dimensionless damage number for amonolithic circular plate subjected to blast loading:

/c ¼I 1þ ln Rp

R0

� �� �pRpH2 ffiffiffiffiffiffiffiffiffir0q

p1

1þ lnðr=R0Þ

� �ð1Þ

where I is impulse, Rp is the plate radius, R0 is load radius, H is platethickness, r0 is static yield stress of the plate, q is plate materialdensity, and r is stand-off distance.

Jacob et al. [27] reported that for a monolithic plate blasted atdifferent stand-off distances (13–300 mm) the results are envel-

1896 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

oped within the 90% confidence lines of the Nurick and Martin [6]empirical relationship:

dH¼ 0:425/c ð2Þ

A graph of deflection–thickness ratio versus dimensionless im-pulse for the front plates of the test series S13, S29 and S150 isshown in Fig. 17. This data does not include the torn plates (ModesII* and II failure). The Nurick and Martin empirical relationship (Eq.(2)) and the 90% and 99.9% confidence lines are superimposed inFig. 17. Some data from S29 and all data from S13 and S150 fall be-low the lower 90% confidence line. It should be noted that Eq. (2)describes the deflection–impulse relationship for an impulsivelyloaded plate that is free to move in air. When backed by anotherstructure (such as a deformable core) it is expected that Eq. (2) willoverestimate the deflections, as shown in Fig. 17. The honeycombcore provides structural support for the front plate under blastingsituations. In addition, at all test impulses above 30 N s the S13data fall below the 99.9% line. This is when the front plate startsto deform into the back plate, due to the low core thickness.

A graph of core centre crushing–thickness ratio versus impulseis shown in Fig. 18, for the S13, S29 and S150 panels. The corecrushing is divided by the original thickness to normalise theeffect of core thickness. It can be seen that with increasing corethickness, the ratio of core crushing is dramatically decreased;

0

5

10

15

20

0 5 10 15 2

Dimensionl

Def

lect

ion-

thic

knes

s ra

tio

S13 [25]

90% confidence line

99.9% c

Fig. 17. Graph of deflection–thickness ratio versus dimensio

0

0.2

0.4

0.6

0.8

1

0 5 10 15 20

Impu

Cru

sh d

ista

nce

- thi

ckne

ss ra

tio

S13 S29 S150

Onset of core densification

Onset of fr

Fig. 18. Graph of crush distance–thickness ratio versus impu

as might be expected since front plate deformation controls corecrushing.

The core crushing gradient change is shown in Fig. 18. There aretwo scenarios – a gradient increase and a gradient decrease:

� In test series S13, as the centre of the core crushes to the maxi-mum limit, the crushing gradient decreases and the best fitapproaches a horizontal line. This shows that core crushing gra-dient decreases when the core reaches complete densification.

� In test series S150, the core was too thick for densification tooccur over the impulse range applied to the front plates. There-fore core crushing gradient increase is due to front plate tearing.

� In test series S29, both core densification and front plate tearingwere observed at a similar impulse range (above approximately31.3 N s) and it exhibited no apparent core crushing gradientchange.

A graph of back plate deflection–impulse of the three test seriesis shown in Fig. 19. At low impulses (less than 24 N s), it is difficultto distinguish by core height. Increasing the core height results indensification at a higher impulse and hence a more rapid increasein the back plate deflection. The load transfer through the core islarge enough to cause back plate tearing after the core densifica-tion. Hence, increasing the core height resulted in back plate tear-ing at a higher threshold impulse. This observation was supported

0 25 30 35 40

ess impulseS29 S150

onfidence line

δ/H = 0.425φc

nless impulse of the front plates of S13, S29 and S150.

25 30 35 40 45 50

lse (Ns)S13 S29 S150

ont plate tearing

lse of the honeycomb core centres of S13, S29 and S150.

Fig. 19. Graph of back plate deflection versus impulse of S13, S29 and S150.

0

5

10

15

20

25

30

0 10 20 30 40 50 60 70 80

Dimensionless impulse

Def

lect

ion-

thic

knes

s ra

tio

S29 S29-1

90% confidence lines

99.9% confidence lines

δ/H = 0.425φc

Fig. 20. Graph of deflection–thickness ratio versus dimensionless impulse of the front plates of S29 and S29-1.

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35 40 45 50Impulse (Ns)

Cru

sh d

ista

nce

(mm

)

S29 Centre S29 Mid-span S29 Outer-span S29-1 CentreS29-1 Mid-span S29-1 Outer-span S29 Centre S29 Mid-spanS29 Outer-span S29-1 Centre S29-1 Mid-span S29-1 Outer-span

Onset of core densif ication

S29-1 densification

S29 densification

Fig. 21. Graph of core crushing versus impulse of S29 and S29-1.

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1897

0

5

10

15

20

25

0 5 10 15 20 25 30 35 40

Impulse (Ns)

Def

lect

ion-

thic

knes

s ra

tio

S29 S29-1

Mode * failure

Onset of core densification for S29for S29-1

Fig. 22. Graph of deflection–thickness ratio versus impulse of the back plates of S29 and S29-1.

1898 Y. Chi et al. / Materials and Design 31 (2010) 1887–1899

by the results of the S150 panel tests, where the core did not reachdensification due to the large core height; hence no back plate tear-ing occurred for charge masses in excess of those causing tearing inthe panels with 13 mm and 29 mm core heights. However, the fullenergy absorption capability of the core was not utilised and hencethe core height was far in excess of that required. This increasedthe panel mass and the volume it occupied. It should also be notedthat there is a variation in the densities of the different thicknesscores, varying by up to approximately 18%. This variation in den-sity is much smaller than the core thickness variation. The densityvariation will have a small influence on the stress transfer charac-teristics of the core but the compressive strain in the honeycomb islower in the thicker cores and thus the stress transfer characteristicis greatly influenced by the core thickness as described above.

5.2. Influence of face plate thickness

The influence of plate thickness is investigated by comparingtest series which used the same core but different plate thick-nesses, namely S29 and S29-1. A graph of deflection–thickness ra-tio versus dimensionless impulse for the front plates of S29 andS29-1 is shown in Fig. 20, with the Nurick and Martin empiricalrelationship (Eq. (2)) superimposed on the graph. All front platedeflections from panels with thinner face plates (S29-1) are belowthe prediction as the stiffness of the honeycomb became moredominant compared to the stiffness of the front plate – this de-creased the normalised front plate deflections. As a result, the hon-eycomb becomes more effective in decreasing the dimensionlessfront plate deflection of thinner face plates.

Since the (dimensional) deflections of the thinner front plateswere found to be higher for a given impulse, it followed that the corecrushing would also be greater for a given impulse. Densificationwas expected to occur at a lower impulse for the panels with thinnerface-plate. (that is, S29-1 compared to S29). This was observedexperimentally, as illustrated in Fig. 21 where the S29-1 cores beganto densify at 21 N s, compared to 30 N s for the S29 cores.

A graph of deflection–thickness ratio versus impulse for theback plates of S29 and S29-1 is shown in Fig. 22. At the same im-pulse, S29-1 exhibited higher back plate deformations, as the frontplate displacements are higher. The onset of core densification andfront plate tearing took place at lower impulses for S29-1 and sothe back plate deflection increased at a faster rate at lower im-pulses. Therefore, tearing of the back plate began at lower impulsesfor the thinner face plate panels: 25 N s for S29-1compared to36 N s for S29. It should be noted that the S29-1 panels are 34%

lighter than the S29 panels and have a 31% decrease in tearingthreshold impulse.

6. Concluding comments

The results of an investigation into the response of honeycombsandwich panels to blast loading have been reported. From post-test analyses, it was shown that prior to core densification, the coreprovided structural support to the front plate and regulated thestress transferred to the back plate. It is observed that the sand-wich panels exhibited four inter-related phenomena, which donot necessarily occur in a fixed order or in every experiment. Thesedepended on the material properties of the sandwich componentsand the panel configuration, such as core material, core height, andface plate thickness.

The front plate deflection responses were similar to those of aclamped monolithic circular plate subjected to uniformly distrib-uted blast loading [23], but with lower plate deflections since thecore provided some structural support. Thinner front plates exhib-ited lower normalised deflections, indicating that the influence ofthe core support was more significant when the face plates weremore flexible. The front plate deflection–impulse curve was usuallycharacterised by a linearly increasing trend-line that ended withthe onset of front plate tearing. The deflection at onset of tearingwas usually the maximum front plate deflection in Mode I. There-after the front plate exhibited Mode II failure (complete tearing)and the deflection started to decrease with increasing impulse.

The impulse required for onset of core densification increased withincreasing core thickness. Increasing the core height delayed theonset of densification, as observed in particular during the testingof S150 panels, as the front plate deflections had to exceed 70–75%of the core height (that is, a minimum of 105 mm) for densificationto occur.

The honeycomb core experienced changes in the rate of crushingwith increasing impulse. An increasing rate was due to front platetearing. A decreasing crush rate was due to the honeycomb reach-ing its physical crushing limit (i.e. onset of core densification at thecentre).

The back plate deflection was usually characterised by two re-gions, with low plate deflections initially because the load transferis limited by the honeycomb core plateau stress. When the corestarted to densify, higher loads were transferred to the back plate,causing it to deflect at a faster rate. This higher gradient region, atsufficiently high impulses, ended with the onset of the back platetearing. The panels with thinner face plate exhibited back plate

Y. Chi et al. / Materials and Design 31 (2010) 1887–1899 1899

tearing at lower impulses due to the higher stress transfer throughthe core and the lower deflection required for strain at the plateboundary to reach critical tearing values.

Acknowledgements

The authors wish to thank Mrs. P. Park-Ross, Mr. G. Newins andMr. P. Jacob at the University of Cape Town for their technical assis-tance. Dr. S. Chung Kim Yuen, Mr. A. Vara, Mr. M. Pitterman and Mr.L.S. Bbosa are acknowledged for their experimental assistance.Prof. D. Karagiozova is also acknowledged for her advice on this re-search. The financial support of the National Research Foundationand Denel Aerospace is gratefully acknowledged.

References

[1] Menkes SB, Opat HJ. Tearing and shear failures in explosively loaded clampedbeams. Exp Mech 1973;13:480–6.

[2] Teeling-Smith RG, Nurick GN. The deformation and tearing of circular platessubjected to impulsive loads. Int J Impact Eng 1991;11(1):77–92.

[3] Nurick GN, Shave GC. The deformation and tearing of thin square platessubjected to impulsive loads – an experimental study. Int J Impact Eng1996;18(1):99–116.

[4] Jones N. Structural Impact. 1st ed. Cambridge University Press; 1989.[5] Nurick GN, Martin JB. Deformation of thin panels subjected to impulsive

loading – a review. Part I: theoretical considerations. Int J Impact Eng1989;8(2):159–70.

[6] Nurick GN, Martin JB. Deformation of thin panels subjected to impulsiveloading – a review. Part II: experimental studies. Int J Impact Eng1989;8(2):171–86.

[7] Fleck NA, Deshpande VS. The resistance of clamped sandwich beams to shockloading. J Appl Mech 2004;71:386–401.

[8] Qui X, Deshpande VS, Fleck NA. Dynamic response of clampedcircular sandwich plate subject to shock loading. J Appl Mech 2004;71(5):637–45.

[9] Xue Z, Hutchinson JW. Preliminary assessment of sandwich plates subject toblast loads. Int J Mech Sci 2003;45(4):687–705.

[10] Xue Z, Hutchinson JW. A comparative study of impulse–resistance metalsandwich plates. Int J Impact Eng 2004;30(10):1283–305.

[11] Zhu F, Lu G. A review of blast and impact of metallic and sandwich structures.Electron J Struct Eng 2007;7:92–101. Special Issue: Loading on Structures.

[12] Yuen SCK, Nurick GN, Theobald MD, Langdon GS. Sandwich panels subjected toblast loading. In: Shukla A, Rajapakse Y, Ravichandran G, editors. Dynamicfailure of materials and structures. Springer; 2009.

[13] Gibson LJ, Ashby MF. Cellular solids, structure and properties. 2nded. Cambridge University Press; 1997.

[14] Lu G, Yu TX. Energy absorption of structures and materials. 2nd ed. WoodheadPublishing Ltd.; 2003.

[15] Ciba-Geigy Plastics and Additives Company. Aeroweb aluminium honeycombfor energy absorption. Duxford, AD-EA/1X, September; 1981.

[16] Hexcel Composites. HexWeb honeycomb sandwich design technology.Duxford, AGU 075b, December; 2000.

[17] Wierzbicki T. Crushing analysis of metal honeycombs. Int J Impact Eng1983;1(2):157–74.

[18] Yamashita M, Gotoh M. Impact behaviour of honeycomb structures withvarious cell specifications – numerical simulation and experiment. Int J ImpactEng 2005;32(1–4):618–30.

[19] Burton WS, Noor AK. Structural analysis of the adhesive bond in a honeycombcore sandwich panel. Finite Elem Anal Des 1997;26:213–27.

[20] Dharmasena KP, Wadley HNG, Xue Z, Hutchinson JW. Mechanical response ofmetallic honeycomb sandwich panel structures to high-intensity dynamicloading. Int J Impact Eng 2008;35:1063–74.

[21] Zhu F, Zhao L, Lu G, Gad E. Finite element analysis of square metallic sandwichpanels under blast loading. In: Proceedings of the 9th internationalsymposium on plasticity and impact mechanics (IMPLAST), Germany; 2007,p. 591–8.

[22] Zhu F, Zhao L, Lu G, Wang Z. Deformation and failure of blast-loaded metallicsandwich panels – experimental investigations. Int J Impact Eng2008;35:937–51.

[23] Broderick JW. Investigation into the response of honeycomb sandwichmaterial subjected to impulsive loading. Undergraduate dissertation,University of Cape Town; 1990.

[24] Farrow GH. The response of impulsively loaded sandwich plates. MScdissertation, University of Cape Town; 1995.

[25] Nurick GN, Langdon GS, Chi Y, Jacob N. Response of blast loaded honeycombsandwich plates – part 1: experiments. Compos Struct 2009:433–41.

[26] Karagiozova D, Nurick GN, Langdon GS. Behaviour of sandwich panels subjectto intense air blasts – part 2: numerical simulations. Compos Struct2009:442–50.

[27] Jacob N, Nurick GN, Langdon GS. The effect of stand-off distance on the failureof fully clamped circular mild steel plates subjected to blast loads. Eng Struct2007;29:2723–36.

[28] Balden VH, Nurick GN. Numerical simulation of the post-failure motion of steelplates subjected to blast loading. Int J Impact Eng 2005;32:14–34.