research article numerical analysis of dynamic...
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Research ArticleNumerical Analysis of Dynamic Response of Corrugated CoreSandwich Panels Subjected to Near-Field Air Blast Loading
Pan Zhang Yuansheng Cheng and Jun Liu
School of Naval Architecture and Ocean Engineering Huazhong University of Science and Technology Wuhan 430074 China
Correspondence should be addressed to Jun Liu hustljhusteducn
Received 7 May 2014 Accepted 24 August 2014 Published 27 October 2014
Academic Editor Jeong-Hoi Koo
Copyright copy 2014 Pan Zhang et al This is an open access article distributed under the Creative Commons Attribution Licensewhich permits unrestricted use distribution and reproduction in any medium provided the original work is properly cited
Three-dimensional fully coupled simulation is conducted to analyze the dynamic response of sandwich panels comprising equalthicknesses face sheets sandwiching a corrugated core when subjected to localized impulse created by the detonation of cylindricalexplosive A large number of computational cases have been calculated to comprehensively investigate the performance of sandwichpanels under near-field air blast loading Results show that the deformationfailure modes of panels depend strongly on stand-offdistance The beneficial FSI effect can be enhanced by decreasing the thickness of front face sheet The core configuration has anegligible influence on the peak reflected pressure but it has an effect on the deflection of a panel It is found that the benefits of asandwich panel over an equivalent weight solid plate to withstand near-field air blast loading are more evident at lower stand-offdistance
1 Introduction
Sandwich panels constructed from light face sheets and rela-tively low density cores are famous for the powerful ability tosupport bending load and to save structural weight Thepotential advantages of sandwich structures for shockmitiga-tion in both water and air blast have been exploited by someresearchers [1ndash3] but most works were focused on far-fieldblast and a less extent study is upon near-field air blast Inthe case of far-field blast shock fronts can be generally treatedwith acousticweak shock limit as the shockwaves attenuate toa soundwave and the impulse transmitted to an infinite rigidplate is twice that of incident shock wave due to the completeshock reflection A classic solution has been built by Taylor[4] to analyze the interaction between acoustic blast wave andfree-standing plate whereas the shock waves are character-ized with high intensity and spatially localization in the caseof near-field blast Moreover the nonlinear compressibilityeffects of air can further enhance the beneficial effects of FSIto reduce the impulse transmitted to structure [5ndash7]
Under high intensity blast loading the mechanical per-formance of metallic sandwich structures with several topol-ogy cores (such as square honeycomb [8 9] triangular hon-eycomb [10] hexagonal honeycomb [11 12] and pyramidal
lattice [13]) has been investigated Zhu et al [8] investigatedthe failure behaviors of honeycomb core sandwich panelunder either uniform or localized air blast loading Severaldistinct failure modes were identified Recent experimentalresults have revealed that the square honeycomb and pyra-midal lattice core sandwich panels suffer significantly smallerback face deflections than solid plates with identical masswhen subjected to high intensity air blast loading It is foundthat the benefits of square honeycomb core [9] and pyramidallattice core [13] sandwich panels over monolithic platesdiminish when the impulse intensity of shock wave isextremely highThis is due to the complete crush of core websat the center of panel and the dynamic fracture of front faces(as a result of the high strength of inertially stabilized trusses)Rimoli et al [14] applied a decoupled wet sand loadingmodeldeveloped by Deshpande et al [15] to simulate the dynamicresponse of edge-clamped corrugated core sandwich panelsand equivalent weight monolithic plates (made of 6061-T6aluminum alloy) subjected to wet sand blast loading A phe-nomenon of additional face sheet stretching and deflectionbetween corrugation web nodes was observed at high impul-sively blast loaded experiments Then Wadley et al [16]used a modified Johnson-Cook constitutive relation and
Hindawi Publishing CorporationShock and VibrationVolume 2014 Article ID 180674 16 pageshttpdxdoiorg1011552014180674
2 Shock and Vibration
140
150
372
175
Securely clampedDetonation point
Symmetry plane
All dimensions (mm)Symmetry plane
60∘
tf
tf
l
Hc
R
tc
Figure 1 Geometric models of the one-quarter corrugated sandwich panel and the one-quarter cylindrical explosive
Cockcroft-Latham failure criterion [17] to model the consti-tutive behavior of 6061-T6 aluminum alloy and adopted afully coupled simulation approach based on discrete particle-based method to make a detailed investigation into FSIeffects and fracturemechanisms of this complicatedmechan-ical problem The simulation results rationalized the localldquostreaksrdquo phenomenon which was arose from a local strongFSI effect and explained how dimples formed at the cornersof the test panels and well predicted panel failure at theclamped edges
Despite a great deal of work on blast performance of sand-wich panels mostly about far-field blast partly on the local-ized blast loading there are a few reports about the dynamicresponse of corrugated core sandwich panels under near-fieldair blast loadingThe aim of present study is to investigate thedynamic behavior of corrugated core sandwich panels sub-jected to highly intense near-field air blast loading and revealthe relationship between the structural performance metricsand geometry parameters by conducting numerical simula-tions
2 Geometry and Materials
The examined sandwich panels consist of two equal thicknessface sheets and corrugated core made of 304 stainless steelThe panels hold a given exposure area of 280mm times 300mmand the web of corrugated core keeps a constant inclinationangle of 60∘ Meanwhile the cylindrical TNT explosive is ofa radius of 175mm and a height of 372mm The geometricmodel of corrugated sandwich panels is depicted in Figure 1The geometric parameters of all sandwich panels consideredin this study are listed in Table 1 The calculated equivalentsolid plate thickness is listed in Table 2
For the face sheets and core webs of sandwich panelsmade of the same material the relative density 120588
119888
of core canbe expressed as
120588119888
=
119905119888
119905119888+ 119867119888cos120572
(1)
where 119905119888is the core web thickness 119867
119888is the core thickness
and 120572 is the inclination angle of core webThe material used to fabricate the corrugated sandwich
panels is annealed 304 stainless steel which has high workhardening potential under explosion loading Therefore theJohnson-Cook plasticity formulation which defines the flowstress as a function of equivalent plastic strain strain rateand temperature is employed in all simulationsThe dynamicflow stress is expressed by the following equation
120590119910= [119860 + 119861(120576
eq119901
)
119899
] [1 + 119888 ln(120576eq119901
1205760
)][1 minus (
119879 minus 119879119903
119879119898minus 119879119903
)
119898
]
(2)
where 120576eq119901and 120576
eq119901are equivalent plastic strain and equivalent
plastic strain rate respectively 119879 is the material temperature119879119903is the room temperature and 119879
119898is the melting temper-
ature of the material 119860 119861 119899 119888 1205760 and 119898 are Johnson-Cook
parameters determined by fitting to the experimental curvesThe material properties and the identified Johnson-Cookparameters [18] for the annealed 304 stainless steel are listedin Table 3 Although 304 stainless steel holds a desirable duc-tile performance to withstand the deformation produced byblast the sandwich panel will fail under severely loaded testscenariosTherefore the failure criterion of 304 stainless steelbased on the effective plastic strain is incorporated into num-erical model The rupture strain of material under impact isequal to 042 according to Ahn et alrsquos work [19]
Generally the surrounding air and the product of TNTexplosion are assumed to behave as ideal gasThe equation ofstate specified for ideal gas can be expressed as follows
119901 = (120574 minus 1) 1205881198921198900 where 120574 =
119862119901
119862V 1198900= 119862V119879 (3)
where 120574 is the adiabatic exponent 120588119892is the density of the gas
1198900is the internal energy 119862
119901and 119862V are the specific heat at
constant pressure and volume respectively and 119879 is the gas
Shock and Vibration 3
Table1Geometric
parametersa
ndsim
ulationresults
ofcompu
tatio
nalcases
considered
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S20-1
3020
1732
1020
1035
579
53021times10
6546
mdash1619
Pitting
Pitting
S20-2
5020
1732
1020
1035
579
1300
0times10
6440
mdash1189
Pitting
Indenting
S20-3
8020
1732
1020
1035
579
29786times10
5284
760
377
Indenting
Mod
eIS20-4
100
20
1732
1020
1035
579
92518times10
4240
352
208
Indenting
Mod
eIS20-5
3020
1732
07
2074
8530
52770times10
6565
mdashmdash
Pitting
Pitting
S20-6
5020
1732
07
2074
8530
13002times10
6479
mdash1304
Pitting
Indenting
S20-7
8020
1732
07
2074
8530
29583times10
5312
1018
494
Indenting
Mod
eIS20-8
100
20
1732
07
2074
8530
92578times10
4274
411
267
Indenting
Mod
eIS20-9
3016
1732
1020
1035
500
47094times10
6495
mdashmdash
Pitting
Pitting
S20-10
5016
1732
1020
1035
500
12140times10
6460
mdash1354
Pitting
Indenting
S20-11
8016
1732
1020
1035
500
23078times10
5318
997
505
Indenting
Mod
eIS20-12
100
161732
1020
1035
500
85525times10
4246
509
281
Indenting
Mod
eIS20-13
3016
1732
07
2074
8450
46998times10
6518
mdashmdash
Pitting
Petalling
S20-14
5016
1732
07
2074
8450
12163times10
6495
mdash1576
Pitting
Indenting
S20-15
8016
1732
07
2074
8450
22944times10
5346
1285
678
Indenting
Mod
eIS20-16
100
161732
07
2074
8450
85567times10
4265
574
347
Indenting
Mod
eIS20-17
3012
1732
1020
1035
419
33315times10
6443
mdashmdash
Pitting
Petalling
S20-18
5012
1732
1020
1035
419
1000
4times10
6470
mdash1869
Pitting
Indenting
S20-19
8012
1732
1020
1035
419
21489times10
5345
1400
720
Indenting
Mod
eIS20-20
100
121732
1020
1035
419
83993times10
4283
771
427
Indenting
Mod
eIS20-21
3025
1732
1020
1035
679
55797times10
6554
mdashmdash
Pitting
Pitting
S20-22
5025
1732
1020
1035
679
13328times10
6423
2026
720
Indenting
Indenting
S20-23
8025
1732
1020
1035
679
30174times10
5290
547
278
Indenting
Mod
eIS20-24
100
25
1732
1020
1035
679
92833times10
4240
227
151
Mod
eIMod
eIS20-25
3020
1732
05
20546
495
52755times10
6597
mdashmdash
Pitting
Petalling
S20-26
5020
1732
05
20546
495
12955times10
6507
mdash1118
Pitting
Indenting
S20-27
8020
1732
05
20546
495
29245times10
5340
1174
583
Indenting
Mod
eIS20-28
100
20
1732
05
20546
495
89711times10
4305
615
383
Indenting
Mod
eIS28-1
3020
2425
1028
763
585
52730times10
6626
mdashmdash
Pitting
Pitting
S28-2
5020
2425
1028
763
585
12972times10
6542
mdash864
Pitting
Indenting
S28-3
8020
2425
1028
763
585
29478times10
5386
905
333
Indenting
Mod
eIS28-4
100
20
2425
1028
763
585
90553times10
4408
457
170
Indenting
Mod
eIS28-5
3020
2425
07
28546
532
52839times10
6585
mdashmdash
Pitting
Pitting
S28-6
5020
2425
07
28546
532
12950times10
6575
mdash1018
Pitting
Indenting
S28-7
8020
2425
07
28546
532
29364times10
5412
1115
456
Indenting
Mod
eIS28-8
100
20
2425
07
28546
532
90679times10
4327
516
240
Indenting
Mod
eI
4 Shock and Vibration
Table1Con
tinued
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S28-9
3016
2425
1028
763
505
47042times10
6587
mdashmdash
Pitting
Pitting
S28-10
5016
2425
1028
763
505
12155times10
6555
mdashmdash
Pitting
Pitting
S28-11
8016
2425
1028
763
505
23130times10
5409
1156
429
Indenting
Mod
eIS28-12
100
162425
1028
763
505
85910times10
4339
614
232
Indenting
Mod
eIS28-13
3016
2425
07
28546
452
47099times10
6619
mdashmdash
Pitting
Petalling
S28-14
5016
2425
07
28546
452
12130times10
6583
mdashmdash
Pitting
Pitting
S28-15
8016
2425
07
28546
452
23342times10
5435
1405
607
Indenting
Mod
eIS28-16
100
162425
07
28546
452
85914times10
4355
682
302
Indenting
Mod
eIS40-1
3020
3464
1040
546
589
52807times10
6714
mdashmdash
Pitting
Pitting
S40-2
5020
3464
1040
546
589
12993times10
6629
mdash77
9Pitting
Indenting
S40-3
8020
3464
1040
546
589
29633times10
5481
1120
270
Indenting
Mod
eIS40-4
100
20
3464
1040
546
589
91120times10
4408
620
144
Indenting
Mod
eIS40-5
3020
3464
07
40388
534
52818times10
674
1mdash
mdashPitting
Pitting
S40-6
5020
3464
07
40388
534
12993times10
6664
mdash614
Pitting
Indenting
S40-7
8020
3464
07
40388
534
29029times10
5504
1311
373
Indenting
Mod
eIS40-8
100
20
3464
07
40388
534
91133times10
4421
673
206
Indenting
Mod
eIS40-9
3016
3464
1040
546
509
46923times10
6686
mdashmdash
Pitting
Pitting
S40-10
5016
3464
1040
546
509
12123times10
6643
mdashmdash
Pitting
Pitting
S40-11
8016
3464
1040
546
509
23034times10
5493
1407
341
Indenting
Mod
eIS40-12
100
163464
1040
546
509
85560times10
4425
785
183
Indenting
Mod
eIS40-13
3016
3464
07
40388
454
47152times10
670
3mdash
mdashPitting
Pitting
S40-14
5016
3464
07
40388
454
12134times10
6665
mdashmdash
Pitting
Pitting
S40-15
8016
3464
07
40388
454
23274times10
5518
1659
482
Indenting
Mod
eIS40-16
100
163464
07
40388
454
85552times10
4438
846
263
Indenting
Mod
eI
Shock and Vibration 5
Table 2 Simulation results of equivalent solid plates
Case number SOD Thickness Peak reflectedpressure Momentum Maximum
deflection Failure mode
119877
(mm)119905119904
(mm)119901119903
(KPa)119875
(kgtimesms)120575119904
(mm) mdash
SP-1 30 58 64101 times 106 1034 3034 Mode ISP-2 50 58 14147 times 106 891 1636 Mode IbSP-3 80 58 28021 times 105 726 775 Mode IbSP-4 100 58 92619 times 104 630 593 Mode IbSP-5 30 53 63730 times 106 1035 3357 Mode ISP-6 50 53 14122 times 106 893 1835 Mode IbSP-7 80 53 26578 times 105 730 860 Mode IbSP-8 100 53 92532 times 104 634 671 Mode IbSP-9 30 50 63287 times 106 1035 3573 Mode ISP-10 50 50 14073 times 106 894 1966 Mode IbSP-11 80 50 25159 times 105 732 928 Mode IbSP-12 100 50 92506 times 104 636 726 Mode IbSP-13 30 45 62330 times 106 1036 mdash PetallingSP-14 50 45 13990 times 106 897 2209 Mode IbSP-15 80 45 23856 times 105 737 1057 Mode IbSP-16 100 45 92358 times 104 641 824 Mode IbSP-17 30 42 61957 times 106 1037 mdash PetallingSP-18 50 42 13891 times 106 899 2374 Mode IbSP-19 80 42 22945 times 105 740 1141 Mode IbSP-20 100 42 92216 times 104 645 884 Mode IbSP-21 30 68 65412 times 106 1032 2524 Mode ISP-22 50 68 14334 times 106 886 1309 Mode IbSP-23 80 68 30678 times 105 719 614 Mode IbSP-24 100 68 93192 times 104 622 462 Mode Ib
Table 3 Material properties and Johnson-Cook parameters for the annealed 304 stainless steel [18]
120588
(kgm3)119864
(GPa) ] 119879119898
(K)119879119903
(K)119862
(JkgK)119860
(MPa)119861
(MPa) 119899 119888 1205760
119898
7900 200 03 1673 293 440 310 1000 065 007 100 100
temperature The material properties of air adopted are fromthe ANSYS-AUTODYNmaterial library shown in Table 4
Additionally the Jones-Wilkins-Lee (JWL) EOS which iscurrently in favor for hydrodynamic calculations of detona-tion product expansions is used to model the behavior ofTNT explosive This equation defines the pressure of det-onation product as a function of relative volume 120588
0120588 and
internal energy per initial volume 1198641198980 as shown in (4)
119901 = 1198601015840
(1 minus
120596120588119890
11987711205880
) 119890minus1198771(1205880120588119890)
+ 1198611015840
(1 minus
120596120588119890
11987721205880
) 119890minus1198772(1205880120588119890)
+
120596120588119890
1205880
1198641198980
(4)
where 119901 is the blast pressure 120588119890and 120588
0are the density of
the explosive and explosive products respectively and 1198641198980
isspecific internal energy of the explosive The parameters
Table 4 Material properties of air [20]
120588
(kgm3)119879
(K)119862
(JkgK) 1205741198900
(kJkg)1225 2882 7176 14 2068 times 105
1198601015840
1198611015840
1198771 1198772 and 120596 are material constants which are related
to the type of explosive The material properties adopted forthe TNT explosive are also from the ANSYS-AUTODYNmaterial library shown in Table 5
3 Numerical Model and Validation
31 Numerical Model The sandwich panels are peripherallyclamped Due to the symmetry of structure and loading con-dition only one-quarter of the corrugated sandwich panel
6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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2 Shock and Vibration
140
150
372
175
Securely clampedDetonation point
Symmetry plane
All dimensions (mm)Symmetry plane
60∘
tf
tf
l
Hc
R
tc
Figure 1 Geometric models of the one-quarter corrugated sandwich panel and the one-quarter cylindrical explosive
Cockcroft-Latham failure criterion [17] to model the consti-tutive behavior of 6061-T6 aluminum alloy and adopted afully coupled simulation approach based on discrete particle-based method to make a detailed investigation into FSIeffects and fracturemechanisms of this complicatedmechan-ical problem The simulation results rationalized the localldquostreaksrdquo phenomenon which was arose from a local strongFSI effect and explained how dimples formed at the cornersof the test panels and well predicted panel failure at theclamped edges
Despite a great deal of work on blast performance of sand-wich panels mostly about far-field blast partly on the local-ized blast loading there are a few reports about the dynamicresponse of corrugated core sandwich panels under near-fieldair blast loadingThe aim of present study is to investigate thedynamic behavior of corrugated core sandwich panels sub-jected to highly intense near-field air blast loading and revealthe relationship between the structural performance metricsand geometry parameters by conducting numerical simula-tions
2 Geometry and Materials
The examined sandwich panels consist of two equal thicknessface sheets and corrugated core made of 304 stainless steelThe panels hold a given exposure area of 280mm times 300mmand the web of corrugated core keeps a constant inclinationangle of 60∘ Meanwhile the cylindrical TNT explosive is ofa radius of 175mm and a height of 372mm The geometricmodel of corrugated sandwich panels is depicted in Figure 1The geometric parameters of all sandwich panels consideredin this study are listed in Table 1 The calculated equivalentsolid plate thickness is listed in Table 2
For the face sheets and core webs of sandwich panelsmade of the same material the relative density 120588
119888
of core canbe expressed as
120588119888
=
119905119888
119905119888+ 119867119888cos120572
(1)
where 119905119888is the core web thickness 119867
119888is the core thickness
and 120572 is the inclination angle of core webThe material used to fabricate the corrugated sandwich
panels is annealed 304 stainless steel which has high workhardening potential under explosion loading Therefore theJohnson-Cook plasticity formulation which defines the flowstress as a function of equivalent plastic strain strain rateand temperature is employed in all simulationsThe dynamicflow stress is expressed by the following equation
120590119910= [119860 + 119861(120576
eq119901
)
119899
] [1 + 119888 ln(120576eq119901
1205760
)][1 minus (
119879 minus 119879119903
119879119898minus 119879119903
)
119898
]
(2)
where 120576eq119901and 120576
eq119901are equivalent plastic strain and equivalent
plastic strain rate respectively 119879 is the material temperature119879119903is the room temperature and 119879
119898is the melting temper-
ature of the material 119860 119861 119899 119888 1205760 and 119898 are Johnson-Cook
parameters determined by fitting to the experimental curvesThe material properties and the identified Johnson-Cookparameters [18] for the annealed 304 stainless steel are listedin Table 3 Although 304 stainless steel holds a desirable duc-tile performance to withstand the deformation produced byblast the sandwich panel will fail under severely loaded testscenariosTherefore the failure criterion of 304 stainless steelbased on the effective plastic strain is incorporated into num-erical model The rupture strain of material under impact isequal to 042 according to Ahn et alrsquos work [19]
Generally the surrounding air and the product of TNTexplosion are assumed to behave as ideal gasThe equation ofstate specified for ideal gas can be expressed as follows
119901 = (120574 minus 1) 1205881198921198900 where 120574 =
119862119901
119862V 1198900= 119862V119879 (3)
where 120574 is the adiabatic exponent 120588119892is the density of the gas
1198900is the internal energy 119862
119901and 119862V are the specific heat at
constant pressure and volume respectively and 119879 is the gas
Shock and Vibration 3
Table1Geometric
parametersa
ndsim
ulationresults
ofcompu
tatio
nalcases
considered
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S20-1
3020
1732
1020
1035
579
53021times10
6546
mdash1619
Pitting
Pitting
S20-2
5020
1732
1020
1035
579
1300
0times10
6440
mdash1189
Pitting
Indenting
S20-3
8020
1732
1020
1035
579
29786times10
5284
760
377
Indenting
Mod
eIS20-4
100
20
1732
1020
1035
579
92518times10
4240
352
208
Indenting
Mod
eIS20-5
3020
1732
07
2074
8530
52770times10
6565
mdashmdash
Pitting
Pitting
S20-6
5020
1732
07
2074
8530
13002times10
6479
mdash1304
Pitting
Indenting
S20-7
8020
1732
07
2074
8530
29583times10
5312
1018
494
Indenting
Mod
eIS20-8
100
20
1732
07
2074
8530
92578times10
4274
411
267
Indenting
Mod
eIS20-9
3016
1732
1020
1035
500
47094times10
6495
mdashmdash
Pitting
Pitting
S20-10
5016
1732
1020
1035
500
12140times10
6460
mdash1354
Pitting
Indenting
S20-11
8016
1732
1020
1035
500
23078times10
5318
997
505
Indenting
Mod
eIS20-12
100
161732
1020
1035
500
85525times10
4246
509
281
Indenting
Mod
eIS20-13
3016
1732
07
2074
8450
46998times10
6518
mdashmdash
Pitting
Petalling
S20-14
5016
1732
07
2074
8450
12163times10
6495
mdash1576
Pitting
Indenting
S20-15
8016
1732
07
2074
8450
22944times10
5346
1285
678
Indenting
Mod
eIS20-16
100
161732
07
2074
8450
85567times10
4265
574
347
Indenting
Mod
eIS20-17
3012
1732
1020
1035
419
33315times10
6443
mdashmdash
Pitting
Petalling
S20-18
5012
1732
1020
1035
419
1000
4times10
6470
mdash1869
Pitting
Indenting
S20-19
8012
1732
1020
1035
419
21489times10
5345
1400
720
Indenting
Mod
eIS20-20
100
121732
1020
1035
419
83993times10
4283
771
427
Indenting
Mod
eIS20-21
3025
1732
1020
1035
679
55797times10
6554
mdashmdash
Pitting
Pitting
S20-22
5025
1732
1020
1035
679
13328times10
6423
2026
720
Indenting
Indenting
S20-23
8025
1732
1020
1035
679
30174times10
5290
547
278
Indenting
Mod
eIS20-24
100
25
1732
1020
1035
679
92833times10
4240
227
151
Mod
eIMod
eIS20-25
3020
1732
05
20546
495
52755times10
6597
mdashmdash
Pitting
Petalling
S20-26
5020
1732
05
20546
495
12955times10
6507
mdash1118
Pitting
Indenting
S20-27
8020
1732
05
20546
495
29245times10
5340
1174
583
Indenting
Mod
eIS20-28
100
20
1732
05
20546
495
89711times10
4305
615
383
Indenting
Mod
eIS28-1
3020
2425
1028
763
585
52730times10
6626
mdashmdash
Pitting
Pitting
S28-2
5020
2425
1028
763
585
12972times10
6542
mdash864
Pitting
Indenting
S28-3
8020
2425
1028
763
585
29478times10
5386
905
333
Indenting
Mod
eIS28-4
100
20
2425
1028
763
585
90553times10
4408
457
170
Indenting
Mod
eIS28-5
3020
2425
07
28546
532
52839times10
6585
mdashmdash
Pitting
Pitting
S28-6
5020
2425
07
28546
532
12950times10
6575
mdash1018
Pitting
Indenting
S28-7
8020
2425
07
28546
532
29364times10
5412
1115
456
Indenting
Mod
eIS28-8
100
20
2425
07
28546
532
90679times10
4327
516
240
Indenting
Mod
eI
4 Shock and Vibration
Table1Con
tinued
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S28-9
3016
2425
1028
763
505
47042times10
6587
mdashmdash
Pitting
Pitting
S28-10
5016
2425
1028
763
505
12155times10
6555
mdashmdash
Pitting
Pitting
S28-11
8016
2425
1028
763
505
23130times10
5409
1156
429
Indenting
Mod
eIS28-12
100
162425
1028
763
505
85910times10
4339
614
232
Indenting
Mod
eIS28-13
3016
2425
07
28546
452
47099times10
6619
mdashmdash
Pitting
Petalling
S28-14
5016
2425
07
28546
452
12130times10
6583
mdashmdash
Pitting
Pitting
S28-15
8016
2425
07
28546
452
23342times10
5435
1405
607
Indenting
Mod
eIS28-16
100
162425
07
28546
452
85914times10
4355
682
302
Indenting
Mod
eIS40-1
3020
3464
1040
546
589
52807times10
6714
mdashmdash
Pitting
Pitting
S40-2
5020
3464
1040
546
589
12993times10
6629
mdash77
9Pitting
Indenting
S40-3
8020
3464
1040
546
589
29633times10
5481
1120
270
Indenting
Mod
eIS40-4
100
20
3464
1040
546
589
91120times10
4408
620
144
Indenting
Mod
eIS40-5
3020
3464
07
40388
534
52818times10
674
1mdash
mdashPitting
Pitting
S40-6
5020
3464
07
40388
534
12993times10
6664
mdash614
Pitting
Indenting
S40-7
8020
3464
07
40388
534
29029times10
5504
1311
373
Indenting
Mod
eIS40-8
100
20
3464
07
40388
534
91133times10
4421
673
206
Indenting
Mod
eIS40-9
3016
3464
1040
546
509
46923times10
6686
mdashmdash
Pitting
Pitting
S40-10
5016
3464
1040
546
509
12123times10
6643
mdashmdash
Pitting
Pitting
S40-11
8016
3464
1040
546
509
23034times10
5493
1407
341
Indenting
Mod
eIS40-12
100
163464
1040
546
509
85560times10
4425
785
183
Indenting
Mod
eIS40-13
3016
3464
07
40388
454
47152times10
670
3mdash
mdashPitting
Pitting
S40-14
5016
3464
07
40388
454
12134times10
6665
mdashmdash
Pitting
Pitting
S40-15
8016
3464
07
40388
454
23274times10
5518
1659
482
Indenting
Mod
eIS40-16
100
163464
07
40388
454
85552times10
4438
846
263
Indenting
Mod
eI
Shock and Vibration 5
Table 2 Simulation results of equivalent solid plates
Case number SOD Thickness Peak reflectedpressure Momentum Maximum
deflection Failure mode
119877
(mm)119905119904
(mm)119901119903
(KPa)119875
(kgtimesms)120575119904
(mm) mdash
SP-1 30 58 64101 times 106 1034 3034 Mode ISP-2 50 58 14147 times 106 891 1636 Mode IbSP-3 80 58 28021 times 105 726 775 Mode IbSP-4 100 58 92619 times 104 630 593 Mode IbSP-5 30 53 63730 times 106 1035 3357 Mode ISP-6 50 53 14122 times 106 893 1835 Mode IbSP-7 80 53 26578 times 105 730 860 Mode IbSP-8 100 53 92532 times 104 634 671 Mode IbSP-9 30 50 63287 times 106 1035 3573 Mode ISP-10 50 50 14073 times 106 894 1966 Mode IbSP-11 80 50 25159 times 105 732 928 Mode IbSP-12 100 50 92506 times 104 636 726 Mode IbSP-13 30 45 62330 times 106 1036 mdash PetallingSP-14 50 45 13990 times 106 897 2209 Mode IbSP-15 80 45 23856 times 105 737 1057 Mode IbSP-16 100 45 92358 times 104 641 824 Mode IbSP-17 30 42 61957 times 106 1037 mdash PetallingSP-18 50 42 13891 times 106 899 2374 Mode IbSP-19 80 42 22945 times 105 740 1141 Mode IbSP-20 100 42 92216 times 104 645 884 Mode IbSP-21 30 68 65412 times 106 1032 2524 Mode ISP-22 50 68 14334 times 106 886 1309 Mode IbSP-23 80 68 30678 times 105 719 614 Mode IbSP-24 100 68 93192 times 104 622 462 Mode Ib
Table 3 Material properties and Johnson-Cook parameters for the annealed 304 stainless steel [18]
120588
(kgm3)119864
(GPa) ] 119879119898
(K)119879119903
(K)119862
(JkgK)119860
(MPa)119861
(MPa) 119899 119888 1205760
119898
7900 200 03 1673 293 440 310 1000 065 007 100 100
temperature The material properties of air adopted are fromthe ANSYS-AUTODYNmaterial library shown in Table 4
Additionally the Jones-Wilkins-Lee (JWL) EOS which iscurrently in favor for hydrodynamic calculations of detona-tion product expansions is used to model the behavior ofTNT explosive This equation defines the pressure of det-onation product as a function of relative volume 120588
0120588 and
internal energy per initial volume 1198641198980 as shown in (4)
119901 = 1198601015840
(1 minus
120596120588119890
11987711205880
) 119890minus1198771(1205880120588119890)
+ 1198611015840
(1 minus
120596120588119890
11987721205880
) 119890minus1198772(1205880120588119890)
+
120596120588119890
1205880
1198641198980
(4)
where 119901 is the blast pressure 120588119890and 120588
0are the density of
the explosive and explosive products respectively and 1198641198980
isspecific internal energy of the explosive The parameters
Table 4 Material properties of air [20]
120588
(kgm3)119879
(K)119862
(JkgK) 1205741198900
(kJkg)1225 2882 7176 14 2068 times 105
1198601015840
1198611015840
1198771 1198772 and 120596 are material constants which are related
to the type of explosive The material properties adopted forthe TNT explosive are also from the ANSYS-AUTODYNmaterial library shown in Table 5
3 Numerical Model and Validation
31 Numerical Model The sandwich panels are peripherallyclamped Due to the symmetry of structure and loading con-dition only one-quarter of the corrugated sandwich panel
6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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Navigation and Observation
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International Journal of
Shock and Vibration 3
Table1Geometric
parametersa
ndsim
ulationresults
ofcompu
tatio
nalcases
considered
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S20-1
3020
1732
1020
1035
579
53021times10
6546
mdash1619
Pitting
Pitting
S20-2
5020
1732
1020
1035
579
1300
0times10
6440
mdash1189
Pitting
Indenting
S20-3
8020
1732
1020
1035
579
29786times10
5284
760
377
Indenting
Mod
eIS20-4
100
20
1732
1020
1035
579
92518times10
4240
352
208
Indenting
Mod
eIS20-5
3020
1732
07
2074
8530
52770times10
6565
mdashmdash
Pitting
Pitting
S20-6
5020
1732
07
2074
8530
13002times10
6479
mdash1304
Pitting
Indenting
S20-7
8020
1732
07
2074
8530
29583times10
5312
1018
494
Indenting
Mod
eIS20-8
100
20
1732
07
2074
8530
92578times10
4274
411
267
Indenting
Mod
eIS20-9
3016
1732
1020
1035
500
47094times10
6495
mdashmdash
Pitting
Pitting
S20-10
5016
1732
1020
1035
500
12140times10
6460
mdash1354
Pitting
Indenting
S20-11
8016
1732
1020
1035
500
23078times10
5318
997
505
Indenting
Mod
eIS20-12
100
161732
1020
1035
500
85525times10
4246
509
281
Indenting
Mod
eIS20-13
3016
1732
07
2074
8450
46998times10
6518
mdashmdash
Pitting
Petalling
S20-14
5016
1732
07
2074
8450
12163times10
6495
mdash1576
Pitting
Indenting
S20-15
8016
1732
07
2074
8450
22944times10
5346
1285
678
Indenting
Mod
eIS20-16
100
161732
07
2074
8450
85567times10
4265
574
347
Indenting
Mod
eIS20-17
3012
1732
1020
1035
419
33315times10
6443
mdashmdash
Pitting
Petalling
S20-18
5012
1732
1020
1035
419
1000
4times10
6470
mdash1869
Pitting
Indenting
S20-19
8012
1732
1020
1035
419
21489times10
5345
1400
720
Indenting
Mod
eIS20-20
100
121732
1020
1035
419
83993times10
4283
771
427
Indenting
Mod
eIS20-21
3025
1732
1020
1035
679
55797times10
6554
mdashmdash
Pitting
Pitting
S20-22
5025
1732
1020
1035
679
13328times10
6423
2026
720
Indenting
Indenting
S20-23
8025
1732
1020
1035
679
30174times10
5290
547
278
Indenting
Mod
eIS20-24
100
25
1732
1020
1035
679
92833times10
4240
227
151
Mod
eIMod
eIS20-25
3020
1732
05
20546
495
52755times10
6597
mdashmdash
Pitting
Petalling
S20-26
5020
1732
05
20546
495
12955times10
6507
mdash1118
Pitting
Indenting
S20-27
8020
1732
05
20546
495
29245times10
5340
1174
583
Indenting
Mod
eIS20-28
100
20
1732
05
20546
495
89711times10
4305
615
383
Indenting
Mod
eIS28-1
3020
2425
1028
763
585
52730times10
6626
mdashmdash
Pitting
Pitting
S28-2
5020
2425
1028
763
585
12972times10
6542
mdash864
Pitting
Indenting
S28-3
8020
2425
1028
763
585
29478times10
5386
905
333
Indenting
Mod
eIS28-4
100
20
2425
1028
763
585
90553times10
4408
457
170
Indenting
Mod
eIS28-5
3020
2425
07
28546
532
52839times10
6585
mdashmdash
Pitting
Pitting
S28-6
5020
2425
07
28546
532
12950times10
6575
mdash1018
Pitting
Indenting
S28-7
8020
2425
07
28546
532
29364times10
5412
1115
456
Indenting
Mod
eIS28-8
100
20
2425
07
28546
532
90679times10
4327
516
240
Indenting
Mod
eI
4 Shock and Vibration
Table1Con
tinued
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S28-9
3016
2425
1028
763
505
47042times10
6587
mdashmdash
Pitting
Pitting
S28-10
5016
2425
1028
763
505
12155times10
6555
mdashmdash
Pitting
Pitting
S28-11
8016
2425
1028
763
505
23130times10
5409
1156
429
Indenting
Mod
eIS28-12
100
162425
1028
763
505
85910times10
4339
614
232
Indenting
Mod
eIS28-13
3016
2425
07
28546
452
47099times10
6619
mdashmdash
Pitting
Petalling
S28-14
5016
2425
07
28546
452
12130times10
6583
mdashmdash
Pitting
Pitting
S28-15
8016
2425
07
28546
452
23342times10
5435
1405
607
Indenting
Mod
eIS28-16
100
162425
07
28546
452
85914times10
4355
682
302
Indenting
Mod
eIS40-1
3020
3464
1040
546
589
52807times10
6714
mdashmdash
Pitting
Pitting
S40-2
5020
3464
1040
546
589
12993times10
6629
mdash77
9Pitting
Indenting
S40-3
8020
3464
1040
546
589
29633times10
5481
1120
270
Indenting
Mod
eIS40-4
100
20
3464
1040
546
589
91120times10
4408
620
144
Indenting
Mod
eIS40-5
3020
3464
07
40388
534
52818times10
674
1mdash
mdashPitting
Pitting
S40-6
5020
3464
07
40388
534
12993times10
6664
mdash614
Pitting
Indenting
S40-7
8020
3464
07
40388
534
29029times10
5504
1311
373
Indenting
Mod
eIS40-8
100
20
3464
07
40388
534
91133times10
4421
673
206
Indenting
Mod
eIS40-9
3016
3464
1040
546
509
46923times10
6686
mdashmdash
Pitting
Pitting
S40-10
5016
3464
1040
546
509
12123times10
6643
mdashmdash
Pitting
Pitting
S40-11
8016
3464
1040
546
509
23034times10
5493
1407
341
Indenting
Mod
eIS40-12
100
163464
1040
546
509
85560times10
4425
785
183
Indenting
Mod
eIS40-13
3016
3464
07
40388
454
47152times10
670
3mdash
mdashPitting
Pitting
S40-14
5016
3464
07
40388
454
12134times10
6665
mdashmdash
Pitting
Pitting
S40-15
8016
3464
07
40388
454
23274times10
5518
1659
482
Indenting
Mod
eIS40-16
100
163464
07
40388
454
85552times10
4438
846
263
Indenting
Mod
eI
Shock and Vibration 5
Table 2 Simulation results of equivalent solid plates
Case number SOD Thickness Peak reflectedpressure Momentum Maximum
deflection Failure mode
119877
(mm)119905119904
(mm)119901119903
(KPa)119875
(kgtimesms)120575119904
(mm) mdash
SP-1 30 58 64101 times 106 1034 3034 Mode ISP-2 50 58 14147 times 106 891 1636 Mode IbSP-3 80 58 28021 times 105 726 775 Mode IbSP-4 100 58 92619 times 104 630 593 Mode IbSP-5 30 53 63730 times 106 1035 3357 Mode ISP-6 50 53 14122 times 106 893 1835 Mode IbSP-7 80 53 26578 times 105 730 860 Mode IbSP-8 100 53 92532 times 104 634 671 Mode IbSP-9 30 50 63287 times 106 1035 3573 Mode ISP-10 50 50 14073 times 106 894 1966 Mode IbSP-11 80 50 25159 times 105 732 928 Mode IbSP-12 100 50 92506 times 104 636 726 Mode IbSP-13 30 45 62330 times 106 1036 mdash PetallingSP-14 50 45 13990 times 106 897 2209 Mode IbSP-15 80 45 23856 times 105 737 1057 Mode IbSP-16 100 45 92358 times 104 641 824 Mode IbSP-17 30 42 61957 times 106 1037 mdash PetallingSP-18 50 42 13891 times 106 899 2374 Mode IbSP-19 80 42 22945 times 105 740 1141 Mode IbSP-20 100 42 92216 times 104 645 884 Mode IbSP-21 30 68 65412 times 106 1032 2524 Mode ISP-22 50 68 14334 times 106 886 1309 Mode IbSP-23 80 68 30678 times 105 719 614 Mode IbSP-24 100 68 93192 times 104 622 462 Mode Ib
Table 3 Material properties and Johnson-Cook parameters for the annealed 304 stainless steel [18]
120588
(kgm3)119864
(GPa) ] 119879119898
(K)119879119903
(K)119862
(JkgK)119860
(MPa)119861
(MPa) 119899 119888 1205760
119898
7900 200 03 1673 293 440 310 1000 065 007 100 100
temperature The material properties of air adopted are fromthe ANSYS-AUTODYNmaterial library shown in Table 4
Additionally the Jones-Wilkins-Lee (JWL) EOS which iscurrently in favor for hydrodynamic calculations of detona-tion product expansions is used to model the behavior ofTNT explosive This equation defines the pressure of det-onation product as a function of relative volume 120588
0120588 and
internal energy per initial volume 1198641198980 as shown in (4)
119901 = 1198601015840
(1 minus
120596120588119890
11987711205880
) 119890minus1198771(1205880120588119890)
+ 1198611015840
(1 minus
120596120588119890
11987721205880
) 119890minus1198772(1205880120588119890)
+
120596120588119890
1205880
1198641198980
(4)
where 119901 is the blast pressure 120588119890and 120588
0are the density of
the explosive and explosive products respectively and 1198641198980
isspecific internal energy of the explosive The parameters
Table 4 Material properties of air [20]
120588
(kgm3)119879
(K)119862
(JkgK) 1205741198900
(kJkg)1225 2882 7176 14 2068 times 105
1198601015840
1198611015840
1198771 1198772 and 120596 are material constants which are related
to the type of explosive The material properties adopted forthe TNT explosive are also from the ANSYS-AUTODYNmaterial library shown in Table 5
3 Numerical Model and Validation
31 Numerical Model The sandwich panels are peripherallyclamped Due to the symmetry of structure and loading con-dition only one-quarter of the corrugated sandwich panel
6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
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International Journal of
4 Shock and Vibration
Table1Con
tinued
Case
number
SOD
Thickn
ess
Cellsize
Relativ
edensity
ofcore
Equivalent
solid
plate
thickn
ess
Peak
reflected
pressure
Maxim
ummom
entum
offro
ntface
Maxim
umdeflection
Failu
remod
e
119877 (mm)
Face
sheets
Core
Cell
wall
119897
(mm)
120588119888 ()
119905119904
(mm)
119901119903
(KPa)
119875
(kgtimes
ms)
Fron
tface
Back
face
Fron
tface
Back
face
119905119891
(mm)
119867119888
(mm)
119905119888
(mm)
120575119891
(mm)
120575119887
(mm)
mdashmdash
S28-9
3016
2425
1028
763
505
47042times10
6587
mdashmdash
Pitting
Pitting
S28-10
5016
2425
1028
763
505
12155times10
6555
mdashmdash
Pitting
Pitting
S28-11
8016
2425
1028
763
505
23130times10
5409
1156
429
Indenting
Mod
eIS28-12
100
162425
1028
763
505
85910times10
4339
614
232
Indenting
Mod
eIS28-13
3016
2425
07
28546
452
47099times10
6619
mdashmdash
Pitting
Petalling
S28-14
5016
2425
07
28546
452
12130times10
6583
mdashmdash
Pitting
Pitting
S28-15
8016
2425
07
28546
452
23342times10
5435
1405
607
Indenting
Mod
eIS28-16
100
162425
07
28546
452
85914times10
4355
682
302
Indenting
Mod
eIS40-1
3020
3464
1040
546
589
52807times10
6714
mdashmdash
Pitting
Pitting
S40-2
5020
3464
1040
546
589
12993times10
6629
mdash77
9Pitting
Indenting
S40-3
8020
3464
1040
546
589
29633times10
5481
1120
270
Indenting
Mod
eIS40-4
100
20
3464
1040
546
589
91120times10
4408
620
144
Indenting
Mod
eIS40-5
3020
3464
07
40388
534
52818times10
674
1mdash
mdashPitting
Pitting
S40-6
5020
3464
07
40388
534
12993times10
6664
mdash614
Pitting
Indenting
S40-7
8020
3464
07
40388
534
29029times10
5504
1311
373
Indenting
Mod
eIS40-8
100
20
3464
07
40388
534
91133times10
4421
673
206
Indenting
Mod
eIS40-9
3016
3464
1040
546
509
46923times10
6686
mdashmdash
Pitting
Pitting
S40-10
5016
3464
1040
546
509
12123times10
6643
mdashmdash
Pitting
Pitting
S40-11
8016
3464
1040
546
509
23034times10
5493
1407
341
Indenting
Mod
eIS40-12
100
163464
1040
546
509
85560times10
4425
785
183
Indenting
Mod
eIS40-13
3016
3464
07
40388
454
47152times10
670
3mdash
mdashPitting
Pitting
S40-14
5016
3464
07
40388
454
12134times10
6665
mdashmdash
Pitting
Pitting
S40-15
8016
3464
07
40388
454
23274times10
5518
1659
482
Indenting
Mod
eIS40-16
100
163464
07
40388
454
85552times10
4438
846
263
Indenting
Mod
eI
Shock and Vibration 5
Table 2 Simulation results of equivalent solid plates
Case number SOD Thickness Peak reflectedpressure Momentum Maximum
deflection Failure mode
119877
(mm)119905119904
(mm)119901119903
(KPa)119875
(kgtimesms)120575119904
(mm) mdash
SP-1 30 58 64101 times 106 1034 3034 Mode ISP-2 50 58 14147 times 106 891 1636 Mode IbSP-3 80 58 28021 times 105 726 775 Mode IbSP-4 100 58 92619 times 104 630 593 Mode IbSP-5 30 53 63730 times 106 1035 3357 Mode ISP-6 50 53 14122 times 106 893 1835 Mode IbSP-7 80 53 26578 times 105 730 860 Mode IbSP-8 100 53 92532 times 104 634 671 Mode IbSP-9 30 50 63287 times 106 1035 3573 Mode ISP-10 50 50 14073 times 106 894 1966 Mode IbSP-11 80 50 25159 times 105 732 928 Mode IbSP-12 100 50 92506 times 104 636 726 Mode IbSP-13 30 45 62330 times 106 1036 mdash PetallingSP-14 50 45 13990 times 106 897 2209 Mode IbSP-15 80 45 23856 times 105 737 1057 Mode IbSP-16 100 45 92358 times 104 641 824 Mode IbSP-17 30 42 61957 times 106 1037 mdash PetallingSP-18 50 42 13891 times 106 899 2374 Mode IbSP-19 80 42 22945 times 105 740 1141 Mode IbSP-20 100 42 92216 times 104 645 884 Mode IbSP-21 30 68 65412 times 106 1032 2524 Mode ISP-22 50 68 14334 times 106 886 1309 Mode IbSP-23 80 68 30678 times 105 719 614 Mode IbSP-24 100 68 93192 times 104 622 462 Mode Ib
Table 3 Material properties and Johnson-Cook parameters for the annealed 304 stainless steel [18]
120588
(kgm3)119864
(GPa) ] 119879119898
(K)119879119903
(K)119862
(JkgK)119860
(MPa)119861
(MPa) 119899 119888 1205760
119898
7900 200 03 1673 293 440 310 1000 065 007 100 100
temperature The material properties of air adopted are fromthe ANSYS-AUTODYNmaterial library shown in Table 4
Additionally the Jones-Wilkins-Lee (JWL) EOS which iscurrently in favor for hydrodynamic calculations of detona-tion product expansions is used to model the behavior ofTNT explosive This equation defines the pressure of det-onation product as a function of relative volume 120588
0120588 and
internal energy per initial volume 1198641198980 as shown in (4)
119901 = 1198601015840
(1 minus
120596120588119890
11987711205880
) 119890minus1198771(1205880120588119890)
+ 1198611015840
(1 minus
120596120588119890
11987721205880
) 119890minus1198772(1205880120588119890)
+
120596120588119890
1205880
1198641198980
(4)
where 119901 is the blast pressure 120588119890and 120588
0are the density of
the explosive and explosive products respectively and 1198641198980
isspecific internal energy of the explosive The parameters
Table 4 Material properties of air [20]
120588
(kgm3)119879
(K)119862
(JkgK) 1205741198900
(kJkg)1225 2882 7176 14 2068 times 105
1198601015840
1198611015840
1198771 1198772 and 120596 are material constants which are related
to the type of explosive The material properties adopted forthe TNT explosive are also from the ANSYS-AUTODYNmaterial library shown in Table 5
3 Numerical Model and Validation
31 Numerical Model The sandwich panels are peripherallyclamped Due to the symmetry of structure and loading con-dition only one-quarter of the corrugated sandwich panel
6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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Shock and Vibration 5
Table 2 Simulation results of equivalent solid plates
Case number SOD Thickness Peak reflectedpressure Momentum Maximum
deflection Failure mode
119877
(mm)119905119904
(mm)119901119903
(KPa)119875
(kgtimesms)120575119904
(mm) mdash
SP-1 30 58 64101 times 106 1034 3034 Mode ISP-2 50 58 14147 times 106 891 1636 Mode IbSP-3 80 58 28021 times 105 726 775 Mode IbSP-4 100 58 92619 times 104 630 593 Mode IbSP-5 30 53 63730 times 106 1035 3357 Mode ISP-6 50 53 14122 times 106 893 1835 Mode IbSP-7 80 53 26578 times 105 730 860 Mode IbSP-8 100 53 92532 times 104 634 671 Mode IbSP-9 30 50 63287 times 106 1035 3573 Mode ISP-10 50 50 14073 times 106 894 1966 Mode IbSP-11 80 50 25159 times 105 732 928 Mode IbSP-12 100 50 92506 times 104 636 726 Mode IbSP-13 30 45 62330 times 106 1036 mdash PetallingSP-14 50 45 13990 times 106 897 2209 Mode IbSP-15 80 45 23856 times 105 737 1057 Mode IbSP-16 100 45 92358 times 104 641 824 Mode IbSP-17 30 42 61957 times 106 1037 mdash PetallingSP-18 50 42 13891 times 106 899 2374 Mode IbSP-19 80 42 22945 times 105 740 1141 Mode IbSP-20 100 42 92216 times 104 645 884 Mode IbSP-21 30 68 65412 times 106 1032 2524 Mode ISP-22 50 68 14334 times 106 886 1309 Mode IbSP-23 80 68 30678 times 105 719 614 Mode IbSP-24 100 68 93192 times 104 622 462 Mode Ib
Table 3 Material properties and Johnson-Cook parameters for the annealed 304 stainless steel [18]
120588
(kgm3)119864
(GPa) ] 119879119898
(K)119879119903
(K)119862
(JkgK)119860
(MPa)119861
(MPa) 119899 119888 1205760
119898
7900 200 03 1673 293 440 310 1000 065 007 100 100
temperature The material properties of air adopted are fromthe ANSYS-AUTODYNmaterial library shown in Table 4
Additionally the Jones-Wilkins-Lee (JWL) EOS which iscurrently in favor for hydrodynamic calculations of detona-tion product expansions is used to model the behavior ofTNT explosive This equation defines the pressure of det-onation product as a function of relative volume 120588
0120588 and
internal energy per initial volume 1198641198980 as shown in (4)
119901 = 1198601015840
(1 minus
120596120588119890
11987711205880
) 119890minus1198771(1205880120588119890)
+ 1198611015840
(1 minus
120596120588119890
11987721205880
) 119890minus1198772(1205880120588119890)
+
120596120588119890
1205880
1198641198980
(4)
where 119901 is the blast pressure 120588119890and 120588
0are the density of
the explosive and explosive products respectively and 1198641198980
isspecific internal energy of the explosive The parameters
Table 4 Material properties of air [20]
120588
(kgm3)119879
(K)119862
(JkgK) 1205741198900
(kJkg)1225 2882 7176 14 2068 times 105
1198601015840
1198611015840
1198771 1198772 and 120596 are material constants which are related
to the type of explosive The material properties adopted forthe TNT explosive are also from the ANSYS-AUTODYNmaterial library shown in Table 5
3 Numerical Model and Validation
31 Numerical Model The sandwich panels are peripherallyclamped Due to the symmetry of structure and loading con-dition only one-quarter of the corrugated sandwich panel
6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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6 Shock and Vibration
Table 5 Material properties of TNT explosive [20]
120588
(kgm3)1198601015840
(GPa)1198611015840
(GPa) 1198771
1198772
120596
C-Jdetonationvelocity(ms)
C-Jenergyunitvolume(kJm3)
C-J pressure(kPa)
1630 37377 375 415 09 035 6930 60 times 106 21 times 107
Flow out boundaryAir
TNT explosivePanel
Figure 2 View of three-dimensional finite element model andhigh-resolution view of the field around sandwich panel (Forinterpretation of the references to colour in this figure legend thereader is referred to the web version of this paper)
and cylindrical explosive are modeled in simulations Thethree-dimensional numerical model used in this study isshown in Figure 2
Both the face sheets and core are meshed using 1mmBelytschko-Tsay shell element based on Mindlin plate theory[20] Meanwhile the surrounding air is modeled using Euler-Godunov solver which is a second-order multimaterial Eulersolver with the default SLIC (simplified line interface cal-culation) transport scheme in ANSYS-AUTODYN and thecylindrical charge fills in the part of the surrounding air Boththe contact between the core cell wall and the face sheets dueto the plastic buckling and the self-contact of the core walldue to cell wall folding are taken into account in simulationsIn addition the Eulerian meshes fully couple with the shellelements During the process of coupling the Euler cellsintersected by the Lagrange interface define a stress profilefor the Lagrange boundary vertices In return the Lagrangeinterface defines a geometric constraint to the flowofmaterialin the Euler gridThe parameter named ldquocover fraction limitrdquois used to determine when a partially covered Euler cell isblended to a neighbor cellThe value of cover fraction limit isset to 05 During the discrete process the size of the smallestcell in the surrounding Euler grid should be at most one halfof the artificial thickness of shell element to guarantee theFSI effect In addition the mesh of surrounding air is biasedtowards the field near to sandwich panels (shown in Figure 2)The mesh optimization has been executed to make a trade-off between the results of computation and the computationefficiency
Numerical oscillation
Exponential decay
Time of arrival
times104
4
3
2
1
0
Pmax
000 005 010 015 020
Time (ms)
Pres
sure
(kPa
)
Figure 3 History of the air-blast induced pressure at the stand-offdistance of 30mm
32 Validation
321 Free Field Air Blast Pressure The 2D axial symmetrymodel with a very fine Eulerian mesh is used to simulate thedetonation and expanding process of a 55 g spherical explo-sive placed in an infinity atmosphere field Figure 3 showsthe pressure-time history of a gauge point positioned at astand-offdistance of 30mm It is found that a slight numericaloscillation of pressure appeared after the pressure reaches thepeak value and followed by exponential decay The impulsecarried by shock wave as a key factor to determinate thedeflection of blast loaded plates can be calculated by thetime integration of pressure It is believed that the error dueto numerical oscillation exerts a weak influence on the areaunder the pressure line
A comparison is made between the simulation resultsand the predicting values of the fitting equation developedby Kinney and Graham [21] in terms of the peak pressureof stand-off distances with range from 30mm to 200mm asshown in Figure 4 It is clear that the numerical results arevery close to the predicting values
322 StructureBlast Interaction A series of experiments ofclamped mild steel quadrangular plates to localized air blastloading were conducted by Jacob et al [22] The test plateswere made of mild steel and cut from the shelf cold rolledplates The blast loading was generated by the detonation of aplastic explosive with circular dish shape The plastic explo-sive was placed at the center of a polystyrene foam pad withthickness of 12mm and detonated by a short stub detonatorwithmass of 1 gThemass of charge is in the range of 35ndash45 g
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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International Journal of
Shock and Vibration 7Pe
ak p
ress
ure (
kPa)
Simulation resultsKinney and Graham
Stand-off distance (mm)40 80 120 160 200
times104
4
3
2
1
0
Figure 4 Peak pressure versus stand-off distance for simulationresults and the fitting equation developed by Kinney and Graham[21]
Therefore the damage power of detonator should be takeninto account in simulations But the outline of the detonatorwas not given The charge height was changed to ensurethat the weight of the charge modeled in simulation wasequal to the summation of the weight of detonator and thecharge used in experiments However this would result in anoverprediction of the damage power of detonator as a resultof that the stand-off distance between the plate and detonatoris smaller
To intuitively validate the numerical approach usedin this study five experimental cases possessing detailedpostmortem profiles are selected to simulate the dynamicresponse process using ANSYS-AUTODYN Exposed area ofthose impacted square plates is 190mm times 129mm and thethickness of the plates is 16mm The material coefficientsused to model the dynamic mechanical behavior of impactedplates are given by Jacob et al [22] The one-quarter finiteelement models of a test plate and air field are shown inFigure 5 The finite element model of plate consists of 5000quadrilateral elements while the finite element model of airfield consists of 1000000 hexahedral elements
Figure 6 shows the cross-sectional profiles of experimentsand numerical simulations in each case It is observed thatpartial tearing appeared in the central area both in theexperiment and simulation for the test case of N01100122and the predicted profiles are closely similar to experimentalresults The comparison of predicted center point deflectionsand measured values is presented in Table 6 The simulationresults fit well with the experimental results except smalloverestimation observed in simulations The main reason isthat the damage power of detonator is overpredicted It isconcluded that the numerical method adopted in the simu-lations is accurate enough which is expected to be applied inthe dynamic analysis of complex constructions to near-fieldair blast loading such as sandwich panels
Figure 5 One-quarter finite element models of quadrangular plateand air field
(a)
(b)
Figure 6 Comparison of cross sectional profiles of (a) experiments[22] and (b) proposed simulations The test numbers from bottomto top are N021100124 N01100123 N22100145 N01100121 andN01100122
4 Numerical Simulation Results
Using sandwich panel for impact mitigation the protectingconstruction designers generally assess the performance ofsandwich panel from following aspects impulse on front
8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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8 Shock and Vibration
Table 6 Experimental and numerical results for selected cases
Test number Mass of charge(g)
Experimental displacement(mm)
Numerical displacement(mm)
Relative error()
N02100124 35 204 228 118N01100123 38 214 252 178N01100121 40 237 267 127N22100145 39 234 259 107N01100122 45 Tearing Tearing mdash
(a) 119905 = 0 120583s (b) 119905 = 43 120583s (c) 119905 = 70 120583s
(d) 119905 = 121 120583s (e) 119905 = 270 120583s (f) 119905 = 454 120583s
Figure 7 AUTODYN screenshots showing the transient response of the detonation products and the sandwich panel (case number S20-1)
plate permanent deflections history of pressure near or onthe FSI surface and the deformationfailure patterns
Figure 7 shows typical results of the fully coupled cal-culation revealing the interaction between the explosivegas products and the sandwich panel (Case number S20-1) Figure 7(a) gives the initial state of the explosive andsandwich panel at t = 0 The cylindrical explosive is instantlyignited at this moment by the detonator placed at the centerpoint of the top face of charge At 119905 = 43 120583s the explosivegas products abruptly expand outward to compress thesurrounding air and then begin to interact with the front
surface of sandwich panel at 119905 = 70 120583s Once interactingthe sandwich panel acts as flow boundary for the expandedexplosive products reversely which exert pressure load onstructure At the initial stage of fluid-structure interactionthe sandwich panel nearly keeps undeformed for two reasons(1) the impartedmomentum to sandwich panel is not enoughlarge and (2) the structural response behaves a degree of timehysteresis A dent failure is firstly formed at the central areaof sandwich front face at 119905 = 121 120583s which has been alsofound by Zhu et al [23] After that the deformation extendsoutwards and downwards with the transferring of impulse
Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration 9
DelaminationCreaseCrease
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(a) Front face
Tearing failure
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(b) Back face
Tearing failure
Plastic bucklingElastic buckling
Stretching
First cellSecond cell Third cell
4200e minus 01
3780e minus 01
3360e minus 01
2940e minus 01
2520e minus 01
2100e minus 01
1680e minus 01
1260e minus 01
8400e minus 02
4200e minus 02
0000e + 00
EFFPLSTN [All]
(c) Corrugated core
Figure 8 Deformation and failure patterns of a sandwich panel (case number S20-1) (For interpretation of the references to colour in thisfigure legend the reader is referred to the web version of this paper)
The front face deforms continually to slap into the back face at119905 = 270 120583s In this process the middle core webs provide theresistance force to the front face and begin to crush Due tothe large stretching and bending deformation the front facefails at the joint between the front face sheet and middle coreweb and then the explosive products permeate into the innerspace of cell As the simulation proceeding the face sheetsand the core continue to deform under their inertia shownin Figure 7(f) and the contact detected between them doesits work in subsequent process till the contact force reducesto 0
After the slight oscillation of the sandwich panel dueto the occurrence of spring back effect all kinetic energyof the structure is dissipated by the plastic bending andstretching of face sheets and the crushing of corrugated coreThe final material status and deformationfailure patternsof a panel are shown in Figure 8 Overall most part of thesandwich panel sinks into plastic status In the central portionof front and back face sheets the pitting failure which ischaracterized by a localized pit and fracture on the surface
occurs and significant tearing failure is observed on both thefront and back faces along the middle ridge of corrugatedcore shown in Figures 8(a) and 8(b) Careful examinationshows that some creases are formed on front face along thoseridges of corrugated core owing to the localized supportingforce of core web and delamination failure between the frontface and core also occurs as indicated in Figure 8(a) Underthe loading force of pressured explosive products the frontface gains considerable momentum away from blast locationand undergoes extremely large stretching deformation whichleads central part material to failing and the eroding effectbased on geometric strain is applied to those excessivelydistorted elements Figure 8(c) illustrates the failure modeof a corrugate core under near-field air blast loading Thecrushing strains of corrugated core webs in the centralportion are greatest enough to fail because of the greatestapplied pressure associated with the charge location Thewebs of the second (sorted from center to outskirts) cellundergo remarkable plastic buckling deformation mean-while the webs of the third cell undergo elastic buckling
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
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Shock and Vibration
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DistributedSensor Networks
International Journal of
10 Shock and Vibration
Solid platePanel front facePanel back face
Material failed
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
00 20 40 60 80 100 120 140
(a)
Solid platePanel front facePanel back face
Defl
ectio
n (m
m)
Distance from center of plate (mm)
35
30
25
15
20
10
5
0
0 20 40 60 80 100 120 140
Material failed
(b)
Figure 9 Sandwich panel (case number S20-1) and solid plate (case number SP-1) sectioned profiles predicted along (a) the directionperpendicular to corrugations and (b) the direction parallel to corrugations
deformation whereas the others nearly remain undeformedThe failure modes of the panels and solid plates tested hereinare identified and listed in Tables 1 and 2 respectively
The predicted sectioned half profiles along the directionsparallel to corrugations and perpendicular to corrugationsare plotted in Figure 9 for the sandwich panel and equivalentweight monolithic plate under the same air blast loadingThe midpoint of back face of sandwich panel only undergoeshalf deflection of solid plate midpoint The difference valuebetween the front face and back face deflections reveals thatthe core crushing effect is outstanding in the center area butvanishes in the outskirts field under near-field air blast Asignificant local bending deformation superimposes on theglobal bending deformation of front face along the directionperpendicular to corrugations in the center field The globalinelastic deformation plays a dominative role in the exteriorzone of front face The local bending deformation is not clearon the back face and front face (along the direction parallel tocorrugations) It is also found that the material of center fieldof front face sheet fails under this high-intensity loading andthe tearing failure occurs on the back face between the firstcell core webs (Figure 8(b))
In order to better understand FSI effect the distributionsof pressure are investigated both temporally and spatiallySeveral pressure gauges are intentionally placed in air fieldabove the shock impacted plates along the axis of cylindricalcharge The exact locations of those pressure gauges areshown in Figure 10The location of Gauge 1 is firstly occupiedby the undeformed test plate at initial state and Gauge 2is placed near the FSI surface to measure the pressure ofthe reflected shock wave The distances between the adjacentgauges are equal to 2mm The pressure-time histories of fivepressure gauges for a sandwich panel (case number S20-1)and an equivalent weightmonolithic plate (case number SP-1)
Shock impacted plate Neutral plane
Detonation point
Cylindrical charge
Axis of cylindricalcharge
Gauge 2Gauge 1
Gauge 3Gauge 4Gauge 5
Figure 10 Locations of pressure gauges placed
under the same explosion loading are shown in Figure 11Theincidentwave travels forward viaGauge 5 firstly It is observedfromGauges 2ndash5 that the pressure jump phenomenon for theincident wave is not evident But this phenomenon ordinarilyexists in far-field explosion This is due to the fact that theimpacted plate restricts the expansion of explosive productsThose air particles close to FSI surface like those aroundGauges 2ndash5 have been rapidly compressed due to the expan-sion of explosive production leading to increase of pressureTherefore at the initial phase the increase of pressure ofthe locations Gauges 2ndash5 is induced not only by the high-pressure incident shock front but also by the highly nonlinearcompression of air medium The pressure of reflected wavedeserves more attention because it is the effective loadingon structure Using a solid plate (case number SP-1) as abenchmark the sandwich panel (case number S20-1) givesa decrease of peak reflected pressure by 173 as a result ofthe beneficial FSI effect It is noticeable that the pressure-time
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
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Journal ofEngineeringVolume 2014
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VLSI Design
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Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
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Electrical and Computer Engineering
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Chemical EngineeringInternational Journal of Antennas and
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International Journal of
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DistributedSensor Networks
International Journal of
Shock and Vibration 11
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
times106
6
5
4
3
2
1
0
000 001 002 003
Time (ms)
Pres
sure
(kPa
)
(a)
times106
6
5
4
3
2
1
0
Pres
sure
(kPa
)
000 001 002 003
Time (ms)
Gauge 1Gauge 2Gauge 3
Gauge 4Gauge 5
(b)
Figure 11 Pressure-time history of gauges placed in air field along the axis of cylindrical charge for the test plates (a) Sandwich panel (casenumber S20-1) and (b) solid plate (case number SP-1)
curves of Gauges 3ndash5 hold a double-humped feature Gauge 3is characterized by a high hump for reflected shock and a lowhump for incident shock As the decay of the reflected shockwave the characteristics of Gauges 4-5 are just reverse to thatof Gauge 3 And the pressure of Gauge 1 increases abruptly asthe explosive products flow into the field occupied by the testplates initially
Figure 12 shows that the peak reflected pressure decaysvery fast from center to outside It conforms that sandwichpanel construction with light face sheet can reduce the reflec-ted pressure However this benefit of the sandwich panelconstruction over an equivalent weight solid plate is evidentonly in the center field of plate under near-field air blastloading It is shown that the locality characteristic inferredfrom the deformation pattern is not exhibited in pressurefield
5 Discussions
The performance of corrugated core sandwich panels underexplosion loading is closely related to the geometric param-eters of panels In this research a parametric study isconducted to reveal the relationship between the key charac-teristics (deformationfailure impulse transmitted and peakpressure near FSI surface) of structural response and severalspecific geometry parameters (stand-off distance face sheetthickness cell size and core web thickness relative densityof core) The results of parametric study are presented in thissection
51 Effect of Stand-off Distance To investigate the effect ofstand-off distance on the deformationfailure pattern and thereflected pressure all panel configurations and solid plates
considered here are subjected to the air blast loading createdby the detonation of the TNT charge with a given mass but atstand-off distances of 30mm 50mm 80mm and 100mmrespectively
Figure 13 shows the final deformationfailure patterns ofthe sandwich panels with the same face sheet thickness (119905
119891=
20mm) and core web thickness (119905119888= 10mm) but different
cell sizes (119897 = 20 28 and 40mm resp) and equivalent weightsolid plates under different stand-off distances It is foundthat the failure modes of front face turn from the pitting fail-ure to indenting failure with the increase of stand-off distanceand that those of back face turn from the pitting failure andpetalling failure to the global dome failure (Mode I failure)Meanwhile the failure modes of equivalent solid plates turnfrom the global dome failure attached by an inner dome tojust the global dome failure Additionally the failure area ofstructures is larger at lower stand-off distance As expectedthe center deflections of sandwich front face back face andsolid plates decreasewith the increase of stand-off distance asshown in Figure 14The plot illustrates that the benefits of thesandwich panel construction over a solid plate to withstandblast loads are clearly evident with smaller back plate deflec-tions compared with the equivalent weight solid plates sub-jected to the same loads However the explosion loading fora given charge mass at lower stand-off distance exhibits highintensity and spatial localization And the back faces of somesandwich panel configurations are likely to fail under thisloadingTherefore the benefits of sandwich panel will vanishin this case Figure 13 shows that tearing damage occurs onthe back face of sandwich panels with cell sizes of 28mmand 40mm subjected to the explosion loading at the stand-off distance of 30mm Careful examination shows that somestreaks are formed on front face between the core webs owingto the effect of a locally strong FSI shown in Figure 13
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
12 Shock and Vibration
Distance from center of plate (mm)0 20 40 60 80 100 120 140
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Solid plateSandwich panel
(a)
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
Distance from center of plate (mm)0 20 40 60 80 100 120 140
Solid plateSandwich panel
(b)
Figure 12 Spatial distributions of peak reflected pressure near the FSI surface along (a) the direction perpendicular to corrugations and (b)the direction parallel to corrugations
SOD = 30 mm SOD = 50 mm SOD = 80 mm SOD = 100mm
Increasing stand-off distance
(a)
Increasing stand-off distance
(b)
Increasing stand-off distance
(c)
Increasing stand-off distance
(d)
Figure 13 Cross-sectional view of the deformationfailure modes of sandwich panels and equivalent solid plates under different stand-offdistances (a) 119905
119891
= 20mm 119905c = 10mm and 119897 = 20mm (b) 119905119891
= 20mm 119905119888
= 10mm and 119897 = 28mm (c) 119905119891
= 20mm 119905119888
= 10mm and119897 = 40mm (d) 119905
119904
= 58mm
The peak reflected pressure of the FSI surface is animportant index for quantitatively analyzing the benefits ofthe FSI effect Figure 15 shows the peak reflected pressure ofthe air particle placed at the center point of the sandwichpanel front faces and the solid plates Those sandwich panelswith the same core configuration (119905
119888= 10mm 119897 = 20mm)
but different face sheet thicknesses (119905119891= 12mm 16mm
20mm and 25mm) are chosen to contrast with equivalentsolid plates It is found that the sandwich panel with a lower
peak reflected pressure performs better than the equivalentweight solid plate as a result of the lower inertia of sandwichpanel front face This benefit of sandwich construction ismore remarkable for air blast loading at lower stand-offdistance and the influence rule of stand-off distance onpeak reflected pressure is the same for panels with differentface sheet thicknesses Using the equivalent solid plates asa benchmark these four sandwich panels with face sheetthicknesses of 12mm 16mm 20mm and 25mm lead to
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
Shock and Vibration 13M
axim
um d
eflec
tion
(mm
)
Stand-off distance (mm)
Solid plate
Cell size = 20 mmCell size = 28 mm
Cell size = 40 mm Equivalent solid plate
Front face
Back face
35
100755025
30
25
20
10
5
0
15
Figure 14 Comparison of the center deflection of the sandwichpanel front face back face and equivalent weight solid plate versusthe stand-off distance 119905
119891
= 20mm 119905119888
= 10mm 119905119904
= 58mm
a decrease in the peak reflected pressures by 4623 25591729 and 147 at the stand-off distance of 30mm 27981374 811 and 702 at the stand-off distance of 50mm635 827 071 and 164 at the stand-off distance of80mm and 892 755 011 and 039 at the stand-offdistance of 100mm respectively
52 Effect of Face Sheet Thickness The thickness of front facesheet is critical to the FSI effect Therefore the investigationof the effect of face sheet thickness is necessary to reveal someinherent laws to guide the design of corrugated core sandwichpanel
As indicated in Figure 15 the thickness of front facehas a significant influence on the peak reflected pressure Ifthe thickest front face (25mm) is adopted as a benchmarkthe other front faces (12mm 16mm and 20mm) give adecrease of peak reflected pressure by 4029 156 and498 at stand-off distance of 30mm 2494 891 and246 at stand-off distance of 50mm 2878 2352 and129 at stand-off distance of 80mm and 952 787and 034 at stand-off distance of 100mm respectively Thebenefit of FSI effect for sandwich construction is enhancedgreatly as the reduction of face sheet thicknessMoreover thiseffect of face sheet thickness is more outstanding under near-field explosion condition
To investigate the effect of face sheet thickness on thecenter deflection of front face and back face the results ofthe sandwich panels with the same core configuration (119905
119888=
10mm 119897 = 20mm) and varied face sheet thicknesses (119905119891=
12mm 16mm 20mm and 25mm) are shown in Figure 16With the increase of the face sheet thickness the deflectionsof midpoint reduce especially under the near-field explosionloading but it is an important issue for a designer to deal withthe confliction properly between the stiffness and weight
tf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
ts = 42mmts = 50mmts = 58mmts = 68mm
times106
2
1
48 49 50 51 52 53
Peak
refle
cted
pre
ssur
e (kP
a) Enlarged view
Stand-off distance (mm)
Enlarged field
times106
6
7
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Figure 15 Comparison of the peak reflected pressure (near the FSIsurface) of sandwich panels and equivalent weight solid plates versusthe stand-off distance For sandwich panels 119905
119888
= 10mm and 119897 =20mm
53 Effect of Core Configuration The core configuration con-tains the core web thickness cell size and relative density ofcore All of them play a notable role in the impact mitigationeffect of corrugated core sandwich panel
The peak reflected pressure of sandwich panels with thesame face sheet thickness (119905
119891= 20mm) but different core
web thicknesses (119905119888= 10mm 07mm and 05mm) and dif-
ferent cell sizes (119897 = 20mm 28mm and 40mm) is plotted inFigure 17 At first glance there is a complete overlap amongthose peak reflected pressure stand-off distance curves It isconcluded that the variations of the core web thickness andcell size result in a negligible change of peak reflected pressureof the midpoint of front face However the core works as afoundation for the front face and simultaneously works as aloading transporter for the back face The configuration of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
14 Shock and VibrationM
axim
um d
eflec
tion
(mm
)
30
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Front facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Back facetf = 12 mmtf = 16 mmtf = 20 mmtf = 25 mm
Figure 16Maximumdeflections of sandwich panels with same coreconfiguration (119905
119888
= 10mm 119897 = 20mm)
core can considerably affect the deformation of front faceand back face It can be observed from Figure 14 that forgiven face sheet thickness and core web thickness largercell sizes (larger core thicknesses for the given inclinationangle of 60∘) result in smaller back face deflections and largerfront face deflections The failure mode of back face is a typeof global deformation the deformation of back face mostlydepends on the global bending resistance of sandwich panelThe larger cell sizes have high rigidity in flexure resulting ina smaller back face deflection under blast loading Howeverthe failure pattern of front face deflection is a type of centrallocalized failure and is dominated by localized deflectionThemaximum deflections of front face are related to the localstiffness of the central field As the increase of cell sizesthe central field between two core webs becomes weak instiffness This is the reason that larger cell sizes result inlarger front face deflections For a given cell size the effectof core web thickness on deformation is shown in Figure 18As expected the deflections of front face and back face reducewith the increase of core web thickness which results in theenhancement of global bending resistance and local stiffnessof sandwich panel
As illustrated previously the crushing mechanism of coremedia is mostly dependent on the relative density of coreThe relative density of core is determined by the core webthickness and the cell size However the relative densityof core has no significant influence on the peak reflectedpressure as inferred from Figure 17
Figure 19 shows the maximum deflections of sandwichpanels (S20-25simS20-28 S28-5simS28-8 and S40-1simS40-4) withdifferent cell sizes and core web thicknesses but the samerelative density of core (546) The variation tendency offront face deflections is not as clear as that of back face For a
times106
6
5
4
3
2
1
0
25 50 75 100
Stand-off distance (mm)
Peak
refle
cted
pre
ssur
e (kP
a)
tc = 10 mmtc = 07 mmtc = 05 mm
(a)
times106
6
5
4
3
2
1
0
Peak
refle
cted
pre
ssur
e (kP
a)
25 50 75 100
Stand-off distance (mm)
Cell size = 20mmCell size = 28mmCell size = 40mm
(b)
Figure 17 Peak reflected pressure of midpoint of sandwich panelswith (a) different core web thicknesses and with (b) different cellsizes
given relative density the cell size plays a leading role in theinfluence of back face deflections
Figure 20 shows the maximum deflections of back face ofthree sandwich panels with similar weight (ie the equivalentsolid plates thicknesses are 579mm 585mm and 589mmresp) but different relative density of core (ie 1035 763and 546 resp) subjected to the same blast loading It isfound that the back face deflections increase with the increaseof core relative density especially at lower stand-off distance
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
Shock and Vibration 15
25
20
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
tc = 10 mmtc = 07 mmtc = 05 mm
Front facetc = 10 mmtc = 07 mmtc = 05 mm
Back face
Figure 18 Maximum deflections of front face and back face ofsandwich panels with different core web thicknesses
30
20
25
15
10
5
025 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Front faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Back faceS20-25simS20-28S28-5simS28-8S40-1simS40-4
Figure 19 Maximum deflections of sandwich panels with the samerelative density (546)
6 Conclusions
A detailed numerical analysis based on fully coupled numer-ical method is presented aiming to reveal the interactionbetween the explosive gas product and sandwich panel Theimportant role of the nonlinear compression of air mediumplayed in beneficial FSI effect under near-field air blast isrevealed by investigating the distributions of shock wavepressure both temporally and spatially It is concluded thatthe nonlinear compression effect is significant under near-field air blast Under near-field air blast loading the frontface of sandwich panel undergoes a significant local bending
120588c = 1035
120588c = 763
120588c = 546
20
15
10
5
0
25 50 75 100
Stand-off distance (mm)
Max
imum
defl
ectio
n (m
m)
Figure 20 Maximum deflections of back face of sandwich panelswith different relative density of core
deformation which leads to the occurrence of streak typedeformation between core webs in the center field while thedeformation pattern of back face is determined by globalbending deformation
The parametric study shows that the failure modes ofsandwich panels depend strongly on stand-off distance Andthe failure area increases with the decrease of stand-off dis-tance The benefits of sandwich construction over solid plateto withstand blast loads are clearly evident with lower backplate deflections and lower peak reflected pressures comparedwith the equivalent weight solid plates subjected to the sameload It is found that those benefits are more evident at lowerstand-off distance As the reduction of face sheet thicknessthe benefit of FSI effect of sandwich construction is enhancedgreatly and the deflections of sandwich front face and backface increase The sandwich panels with a light face sheetare more likely to fail Therefore it is essential to optimizethe configuration of sandwich panel to keep back face fromdamage failure The core configuration has a negligible influ-ence on the peak reflected pressure By adopting larger cellsizes and thicker core webs the deflections of back faces canbe reduced For a given areal density of sandwich panel theback face deflections increase with the increase of corerelative density
Conflict of Interests
The authors declare that there is no conflict of interestsregarding the publication of this paper
Acknowledgments
The reported research is supported by the National Nat-ural Science Founding of China (under the Contract no51209099) and the Supporting Technology Funding of Ship-building Industry The financial contributions are herebygratefully acknowledged
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
16 Shock and Vibration
References
[1] Z Xue and J W Hutchinson ldquoPreliminary assessment of sand-wich plates subject to blast loadsrdquo International Journal ofMech-anical Sciences vol 45 no 4 pp 687ndash705 2003
[2] L Yueming A V Spuskanyuk S E Flores et al ldquoThe responseof metallic sandwich panels to water blastrdquo Journal of AppliedMechanics Transactions ASME vol 74 no 1 pp 81ndash99 2007
[3] Z Xue and J W Hutchinson ldquoA comparative study of impulse-resistant metal sandwich platesrdquo International Journal of ImpactEngineering vol 30 no 10 pp 1283ndash1305 2004
[4] G I Taylor ldquoThe pressure and impulse of submarine explosionwaves on platesrdquo in The Scientific Papers of Sir Geoffrey IngramTaylor Aerodynamics and the Mechanics of Projectiles andExplosions G K Batchelor Ed vol 3 pp 287ndash301 CambridgeUniversity Press Cambridge UK 1963
[5] N Kambouchev L Noels and R Radovitzky ldquoNonlinear com-pressibility effects in fluid-structure interaction and their impli-cations on the air-blast loading of structuresrdquo Journal of AppliedPhysics vol 100 no 6 Article ID 063519 2006
[6] N Kambouchev L Noels and R Radovitzky ldquoNumerical simu-lation of the fluid-structure interaction between air blast wavesand free-standing platesrdquo Computers and Structures vol 85no 11-14 pp 923ndash931 2007
[7] N KambouchevAnalysis of blast mitigation strategies exploitingfluid-structure interaction [PhD thesis]Massachusetts Instituteof Technology Cambridge Mass USA 2007
[8] F Zhu G Lu D Ruan and D Shu ldquoTearing of metallic sand-wich panels subjected to air shock loadingrdquo Structural Engineer-ing and Mechanics vol 32 no 2 pp 351ndash370 2009
[9] K P Dharmasena H N G Wadley Z Xue and J W Hutchin-son ldquoMechanical response of metallic honeycomb sandwichpanel structures to high-intensity dynamic loadingrdquo Interna-tional Journal of Impact Engineering vol 35 no 9 pp 1063ndash1074 2008
[10] Z Wei V S Deshpande A G Evans et al ldquoThe resistance ofmetallic plates to localized impulserdquo Journal of the Mechanicsand Physics of Solids vol 56 no 5 pp 2074ndash2091 2008
[11] G N Nurick G S Langdon Y Chi andN Jacob ldquoBehaviour ofsandwich panels subjected to intense air blast Part 1 Experi-mentsrdquo Composite Structures vol 91 no 4 pp 433ndash441 2009
[12] D Karagiozova G N Nurick andG S Langdon ldquoBehaviour ofsandwich panels subject to intense air blastsmdashpart 2 numericalsimulationrdquo Composite Structures vol 91 no 4 pp 442ndash4502009
[13] K P Dharmasena H N G Wadley K Williams Z Xue andJ W Hutchinson ldquoResponse of metallic pyramidal lattice coresandwich panels to high intensity impulsive loading in airrdquoInternational Journal of Impact Engineering vol 38 no 5 pp275ndash289 2011
[14] J J Rimoli B Talamini J J Wetzel K P Dharmasena RRadovitzky andH N GWadley ldquoWet-sand impulse loading ofmetallic plates and corrugated core sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 38 no 10 pp 837ndash8482011
[15] V S Deshpande R MMcMeeking H N GWadley and A GEvans ldquoConstitutive model for predicting dynamic interac-tions between soil ejecta and structural panelsrdquo Journal of theMechanics and Physics of Solids vol 57 no 8 pp 1139ndash11642009
[16] HNGWadley T Borvik L Olovsson et al ldquoDeformation andfracture of impulsively loaded sandwich panelsrdquo Journal of theMechanics and Physics of Solids vol 61 no 2 pp 674ndash699 2013
[17] M Cockcroft and D Latham ldquoDuctility and the workability ofmetalsrdquo Journal of the Institute ofMetals vol 96 no 1 pp 33ndash391968
[18] S Lee F Barthelat J W Hutchinson and H D EspinosaldquoDynamic failure of metallic pyramidal truss core materialsmdashexperiments and modelingrdquo International Journal of Plasticityvol 22 no 11 pp 2118ndash2145 2006
[19] D-G Ahn G-J Moon C-G Jung G-Y Han and D-Y YangldquoIMPACT behavior of A STS 304H sheet with a thickness of 07MMrdquo Arabian Journal for Science and Engineering vol 34 no1 pp 57ndash71 2009
[20] AUTODYN Theory Manual Revision 43 Century DynamicsConcord Mass USA 2005
[21] G F Kinney andK J Graham Explosive Shocks in Air SpringerNew York NY USA 2nd edition 1985
[22] N Jacob K Y Chung G N Nurick D Bonorchis S A Desaiand D Tait ldquoScaling aspects of quadrangular plates subjectedto localised blast loadsmdashexperiments and predictionsrdquo Interna-tional Journal of Impact Engineering vol 30 no 8-9 pp 1179ndash1208 2004
[23] F Zhu L Zhao G Lu and E Gad ldquoA numerical simulation ofthe blast impact of square metallic sandwich panelsrdquo Interna-tional Journal of Impact Engineering vol 36 no 5 pp 687ndash6992009
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of
International Journal of
AerospaceEngineeringHindawi Publishing Corporationhttpwwwhindawicom Volume 2014
RoboticsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Active and Passive Electronic Components
Control Scienceand Engineering
Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
International Journal of
RotatingMachinery
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporation httpwwwhindawicom
Journal ofEngineeringVolume 2014
Submit your manuscripts athttpwwwhindawicom
VLSI Design
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Shock and Vibration
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Civil EngineeringAdvances in
Acoustics and VibrationAdvances in
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Electrical and Computer Engineering
Journal of
Advances inOptoElectronics
Hindawi Publishing Corporation httpwwwhindawicom
Volume 2014
The Scientific World JournalHindawi Publishing Corporation httpwwwhindawicom Volume 2014
SensorsJournal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Modelling amp Simulation in EngineeringHindawi Publishing Corporation httpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Chemical EngineeringInternational Journal of Antennas and
Propagation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
Navigation and Observation
International Journal of
Hindawi Publishing Corporationhttpwwwhindawicom Volume 2014
DistributedSensor Networks
International Journal of