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1

2 Numerical simulation and validation of impact response of axially-restrained

3 steelconcretesteel sandwich panels

4 Alex M. Remennikov , Sih Ying Kong

5 School of Civil, Mining and Environmental Engineering, University of Wollongong, Wollongong, NSW 2522, Australia

6

8a r t i c l e i n f o

9 Article history:0 Available online xxxx

1 Keywords:2 Steelconcretesteel sandwich panels3 Protective structures4 Impact load5

6

a b s t r a c t

Steelconcretesteel (SCS) sandwich panels are an effective means for protecting personnel and infra-structure facilities from the effects of external blast and high-speed vehicle impact. In conventional

SCS construction, the external steel plates are connected to the concrete infill by welded shear stud con-

nectors. This paper describes a programme of research in which the non-composite SCS panels with axi-

ally restrained connections were studied experimentally and numerically. High fidelity finite element

models for axially restrained steelconcretesteel panels subjected to impact loading conditions were

developed using LS-DYNA. The simulation results were validated against the dynamic testing experimen-

tal results. The numerical models were able to predict the initial flexural response of the panels followed

by the tensile membrane resistance at large deformation. It was found that the strain rate effects of the

materials and the concrete material model could have significant effect on the numerically predicted flex-

ural strength and tensile membrane resistance of the panels.

2012 Published by Elsevier Ltd.

0

1 1. Introduction

2 Composite steelconcretesteel (SCS) or double skin composite3 structures consist of a concrete core connected to two steel face-4 plates using mechanical shear connectors. This form of construc-5 tion was originally conceived during the initial design stages for6 the Convy River submerged tube tunnel in the UK [1]. It was sub-7 sequently used in building cores, gravity seawalls, nuclear struc-8 tures and defence structures.9 Shear resistance between the steel and concrete interfaces is of0 prime importance to achieve a full composite action. Current tech-1 niques for achieving composite action using mechanical shear con-2 nectors include headed studs, friction-welded bars and J-hooks.3 Oduyemi and Wrigth [2], Wright et al. [3] and Shanmugam and4 Kumar [4] carried out experimental investigations on the response

5 of SCS structural members with headed shear studs subjected to6 static loading. Corus UK have developed the Bi-steel panels with7 transverse steel bars friction-welded to both steel faceplates simul-8 taneously [5]. Liew and Sohel [6] presented double J-hook connec-9 tors to interlock the steel faceplates. The Bi-steel and J-hook are not0 only effective in resisting longitudinal shear between steel and1 concrete interface, they also provide transverse shear resistance.2 Steelconcretesteel panels are an effective means of protecting3 structures against extreme impact and blast loading due to

their high strength and high ductility characteristics. Young and

Coyle [7] tested Bi-steel panels subjected to contact and close-indetonations of high explosives. Liew et al. [8] performed low-

velocity impact tests on the J-hook panels filled with lightweight

concrete.

Remennikov et al. [9,10] investigated experimentally the static

and impact performance of non-composite SCS panels in which no

shear connectors were utilised to connect the steel faceplates and

the concrete core. Owing to the specially designed connection de-

tails, the tested panels exhibited tensile membrane resistance at

large deformation. The tensile membrane resistance achieved in

the tests was significantly higher than the flexural resistance. The

panels were capable of resisting large support rotations, more than

15, without collapse. Kong et al. [11] presented FE simulation

techniques for the SCS panels under static and impact loading

and validated the simulation results against the experimental re-sults presented in Remennikov et al. [9].

This paper presents three dimensional FE modelling tech-

niques for the axially restrained SCS panels under impact loading

conditions using the non-linear transient dynamic finite element

program LS-DYNA [12]. These techniques have been validated

against the experimental results of SCS panels under impact load-

ing conditions, utilising the drop hammer facility at the Univer-

sity of Wollongong. Parametric studies include investigation on

the effects of the contact surface algorithms, concrete density,

strain rate effects and erosion of concrete elements were carried

out.

0263-8223/$ - see front matter 2012 Published by Elsevier Ltd.http://dx.doi.org/10.1016/j.compstruct.2012.05.011

Corresponding author. Tel.: +61 2 4221 5574; fax: +61 2 4221 3238.

E-mail address: [email protected] (A.M. Remennikov).

Q1

Composite Structures xxx (2012) xxxxxx

Contents lists available at SciVerse ScienceDirect

Composite Structures

j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / c o m p s t r u c t

COST 4643 No. of Pages 11, Model 5G

26 May 2012

Please cite this article in press as: Remennikov AM, Kong SY. Numerical simulation and validation of impact response of axially-restrained steelconcrete

steel sandwich panels. Compos Struct (2012), http://dx.doi.org/10.1016/j.compstruct.2012.05.011
http://dx.doi.org/10.1016/j.compstruct.2012.05.011mailto:[email protected]://dx.doi.org/10.1016/j.compstruct.2012.05.011http://www.sciencedirect.com/science/journal/02638223http://www.elsevier.com/locate/compstructhttp://dx.doi.org/10.1016/j.compstruct.2012.05.011http://dx.doi.org/10.1016/j.compstruct.2012.05.011http://www.elsevier.com/locate/compstructhttp://www.sciencedirect.com/science/journal/02638223http://dx.doi.org/10.1016/j.compstruct.2012.05.011mailto:[email protected]://dx.doi.org/10.1016/j.compstruct.2012.05.011

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2. Impact tests description

High-velocity impact tests were performed on four axially re-

strained steelconcretesteel panels. The configurations of the

tested panels are described in Table 1. The Control panel had a nor-

mal weight concrete core and mild steel faceplates. For the other

panels, one design parameter was varied to obtain a more detailed

picture of the response of the non-composite SCS sandwich panelsunder impact loading as shown in Table 1.

Details of the experimental setup for the impact tests are illus-

trated in Fig. 1. The panels were axially restrained with specially

designed keyed connections. It consists of keyed inserts, UC sec-

tions, I-beam and bolt connections. The panel was connected to

the UC sections through the keyed connections, while the bottom

flanges of the UC sections were bolted to the I beam using M25

high strength bolts. The bolt connections between the UC sections

and the I-beam, together with the angle bracing restrained the in-

plane movement of the UC sections, providing axial restraint on

the panel. The mass of the drop hammer was approximately

600 kg. It was released from a height of three metres to impact

the panels at the mid-span. During the experiments, a 1600 kN

capacity load cell, mounted on the drop hammer was used to mea-

sure the load time history. The displacement of the panel at mid-

span was measured using a high speed draw wire potentiometer.

The experimental data was acquired by the National Instrument

PXI data acquisition system using a sampling rate of 100,000 sam-

ples per second.

3. LS-DYNA modelling

Finite element analysis of the axially restrained steelconcrete

steel sandwich panels was carried out using the three-dimensional

non-linear transient dynamic finite element code LS-DYNA [12].

This program uses explicit time integration for the transient dy-

namic analysis which is suitable for the application of short dura-

tion events such as vehicle crash and explosion. The explicit

analysis considers dynamic equilibrium of each node at every time

step as follow

Man Fexn Fintn 1

where the Mis the mass on the node, a is the acceleration, Fex is the

external force, Fint is the internal force and n is the time step.

12The accuracy of the simulation is determined by monitoring the12total energy in the structures during the analysis. The total energy12in the structures consists of kinetic energy (Ekin), internal energy12(Eint), friction energy (Efr), damping energy (Edamp) and hourglass12energy (Ehg). The internal energy includes elastic strain energy12and work done in plastic deformation. The hourglass energy is12attributed to nonphysical modes of deformation occured in un-12

der-integrated element formulation. It should be less than 10% of 13the peak internal energy of each part.13

Ekin Eint Efr Edamp Ehg Etotal constant 2 1313

134. Numerical model description

13In the finite element models developed for this study, only a13quarter of the experimental setup was considered due to the sym-13metry of the specimen, loading and support conditions, to save the13computational time. The axial restraints, including the keyed in-13serts, bolted connections, steel UC section and steel I-beam were14modelled in detail, as shown in Fig. 2. From the convergence study,14a mesh size of 10 mm was found to be appropriate for the concrete14core and the steel faceplates. Fully integrated selectively reduced14(S/R) solid element formulation was applied to the steel UC section,14I-beam, and the bolts, to avoid hourglass effects in these elements.14Under large deformation, the fully integrated S/R solid elements14become unstable and may cause error termination (negative vol-14ume) in the simulation. To avoid the negative volume effect, the14concrete core of the panel was modelled using constant stress solid14elements. The steel faceplates were modelled using Belytschko1Tsay shell elements. The HughesLiu with cross section integration1beam elements was used to model the reinforcing steel elements.1The bolts of the keyed inserts were simplified as square bars with1the cross sectional area equivalent to M16 bolts. This simplification1has no significant effect on the accuracy of the model as shown by1Lee et al. [13].

14.1. Boundary conditions

1The models were defined so that they were symmetrical in the1x- and y-directions. For symmetry in the x-direction, the x transla-1tional degree of freedom of the solid elements was restrained. For16the symmetry in the y-direction, the y translational degree of free16of solid elements were restrained. For the shell elements, there are

Table 1

The parameters evaluated in the impact tests for steelconcretesteel panels.

No. Panel Parameter

11 Control pa ne l (CP) N orma l we ig ht conc re te c ore a nd mid ste el f ace plat es

22 S ta inless stee l p anel ( SP) S tainless ste el f ace plat es

43 Lightweight concrete core panel (LP) Lightweight concrete core, with a density and concrete compressive strength of 1400 kg/m3

and 10.5 MPa respectively64 Reinforced concrete core panel (RP) Normal weight concrete core was reinforced with two layers of 4@50 mm wire meshes.

Fig. 1. Experimental setup for an axially restrained steelconcretesteel panel subjected to mid-span impact.

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2 6 of freedom at each node, namely translational and rotational at3 the x-, y- and z-direction. For the x symmetry of the shell elements,4 the x translational, and y and zrotational degrees of freedom were5 fixed. While the y translational degree of freedom, and x and zrota-6 tional degrees of freedom of shell elements were restrained for the7 y symmetry.8 The I-beam was bolted down to the strong floor in the experi-

9 mental setup. In the model, all the nodes at the bottom surface0 of the I-beam were restrained in the three translational degrees1 of freedom, x, y and z, so no movement was allowed at these nodes.2 The impactor was positioned 2 mm from the top faceplate, and it3 was assigned with an initial velocity in the z direction that corre-4 sponded to the drop height used in the tests. By using the energy5 conservation approach, an initial velocity of the drop hammer6 immediately before it strikes the panel can be calculated. The ini-7 tial velocity was determined as 7.67 m/s for the 3 m drop height.

8 4.2. Contact surfaces

9 The interaction between different parts in the model was0 important in order to predict the behaviour of the SCS panels cor-1 rectly. In this study, the Automatic-Surface-to-Surface contact2 algorithm in LS-DYNA was used to model the interaction between3 the following parts in the model:

4 1. Impactor and steel faceplates.5 2. Steel faceplates and the concrete core.6 3. Steel faceplates with the UC section and keyed inserts.7 4. Steel UC section and I beam.8 5. Bolts and the UC section and I beam.9

0 In this penalty-based contact algorithm, when a penetration is1 found for the parts in contact, a force proportional to the penetra-2 tion depth is applied to these interfaces to eliminate the penetra-3 tion [14]. Thus, the impact load time histories for the models can4 be obtained by using this contact algorithm between the impactor5 and the steel faceplates. The definition of the master surface and

slave surface is arbitrary, but normally the surface with a finer

mesh will be defined as the slave surface.

This sliding contact algorithm only considers the friction inter-

action between the contact interfaces. The dynamic coefficient of

friction of 0.2 was applied to the mild steel and concrete core inter-

faces. For the contact interfaces between the stainless steel and the

concrete core, the dynamic coefficient of friction of 0.1 was used

considering the surface of the stainless steel was smoother thanthe mild steel. In these models, the chemical bonding of the con-

crete was ignored. This was a realistic assumption because the

chemical bonding failed during the panel handling and experimen-

tal set up before the tests commenced. For the Reinforced core pa-

nel, the nodes of the beam elements were merged with the nodes

of the concrete elements and the slippage between these interfaces

in reality was ignored.

4.3. Material models

4.3.1. Mild steel and stainless steel

The complete stressstrain relationships for both the mild steel

and the stainless steel faceplates were modelled using the LS-

DYNA Piecewise Linear Plasticity material model. For the mild

steel, the yield stress was 271 MPa from the tensile coupon tests,

whilst for the stainless steel, the yield stress was 291 MPa. The

non-linear behaviour after yielding was considered by defining

plastic stressstrain relationships for both steels according to the

tensile coupon tests, as shown in Fig. 3. The strain rate effects of

the mild steel and stainless steel was considered in the model by

specifying the CowperSymonds coefficients. The CowperSy-

monds coefficients for mild steel are 40.4 (D) and 5 (q), while for

the stainless steel, they are 100 (D) and 10 (q) [15]. The impactor

was assumed to be absolutely rigid since there was no deformation

observed on the drop hammer during the tests. The steel UC sec-

tion, I-beam, bolts and wire meshes were assumed to behave as

elastic perfectly plastic material and modelled using the LS-DYNA

Plastic Kinematic material model. The yield stress for the UC sec-

tion and I-beam was assumed as 300 MPa, while the yield stressfor the bolts was assumed as 600 MPa. The yield stress of the wire

mesh was assumed as 450 MPa.

4.3.2. Concrete infill

The CSCM (Continuous Surface Cap Model) Concrete (Mat 159)

material model was developed for the US Federal Highway Admin-

istration (FHWA) to simulate the reinforced concrete structures

subjected to impact loading conditions [16]. This model is simple

to use as it can generate default parameters for the concrete by

only requiring some basic inputs. The details of theoretical descrip-

tion and validation of the LS-DYNA CSCM_Concrete model are pro-

vided in the Federal Highway Administration reports [16,17]. For

this model, the three inputs required to generate the default

parameters are the unconfined compressive strength, aggregatesize and the units used in the finite element model [12].

According to the FHWA report [17], the model is applicable for

concrete grade between 20 MPa and 58 MPa with the aggregate

size between 8 mm and 32 mm. The unconfined compressive

strength will affect all the generated parameters such as stiffness,

three dimensional yield strength, hardening and softening, while

the aggregate size only affect the softening behaviour. The param-

eters are generated based on the Comite Euro-International Du Be-

ton-Fdration Internationale de la Prcontrainte Model (CEB-FIP)

Code.

The yield surface of this concrete model is defined by three

invariants and the cap hardening parameter j, as follows:

fJ1;

J02;

J03;j J

02 K

2

F2fFc 3

Fig. 2. A quarter model of the experimental set up for axially restrained SCS panels

under the drop mass impact test.

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4 for solid elements (type 3) and FlanaganBelytschko stiffness form5 with exact volume integration for solid elements (type 5) were not6 effective to control the hourglass energy. This is due to localised7 impact condition and the panel underwent very large deformation.8 The hourglass energy in the concrete core can be reduced to about9 15% of the peak internal energy when the strain rate effect was ig-0 nored. Therefore, the strain rate effect of the concrete was ignored1

in this study to minimise the hourglass energy in the concrete core.2 The concrete compressive strength for panels was different be-3 cause the panels were casted using different batches of concrete.4 For the Control panel and the Reinforced core panel, the concrete5 compressive strength was 23 MPa. For the Stainless steel panel,6 the concrete compressive strength was 37 MPa while the concrete7 compressive strength for the lightweight panel was 10.5 MPa.8 The density of the lightweight concrete was 1400 kg/m3 and no9 aggregates were used in the mix. Single element simulation [16]0 was carried out to evaluate the ability of the concrete model CSCM1 (Mat 159) to generate parameters for the lightweight concrete. It2 was found that using the density of 1400 kg/m3, concrete compres-3 sive strength of 16 MPa and ignoring the aggregate size, the4 concrete model can generate a stressstrain curve with the com-5 pressive strength of 10.8 MPa and tensile strength of 0.9 MPa, as6 shown in Fig. 3b. It was assumed that this stressstrain relation-7 ship was appropriate for the lightweight concrete in this study.8 The stressstrain relationships for different grades of concrete used9 in this study were generated by a single element simulation and0 the results were shown in Fig. 3b. The stressstrain relationships1 included the compressive strength, tensile strength, softening2 curves after the concrete reached its maximum strengths.

3 4.4. Element erosion

4 An erosion algorithm is available in the LS-DYNA code which al-5 lows computation to be carried out without the need for re-zoning6 distorted regions of the mesh during a large deformation loading.7 The erosion algorithm is based on the concept that the highly

8 strained elements of the deformed mesh have failed totally and9 may no longer contribute to the physics of structural response.0 Upon erosion, the sliding interface between the steel faceplates1 and the concrete core needs to be re-defined dynamically due to2 the total element failure.3 The erosion of the elements is based on somewhat ad hoc crite-4 ria related to a deformation or stress measure in the element. In5 the LS-DYNA concrete material model CSCM_Concrete, the element6 removal can be activated by specifying an erosion coefficient.7 When the erosion coefficient is less than unity, no erosion occurs,8 while erosion coefficient of 1 means the erosion is independent of

the strain. For erosion depending on the maximum principal strain,

erosion coefficient is set to values greater than 1. For example,

when erosion coefficient is set to 1.05, the element is deleted once

it reaches maximum principal strain of 5%.

It should also be realised that the erosion strain has no correla-

tion with the fracture strain and is solely a measure of how much

plastic deformation an element can undergo before it is removed

from the numerical computation. In this study, based on a prioriknowledge of the experimental outcomes, it was found that it

would not be beneficial to include the eroding-element technique

for the concrete core of the panels. This was because it may result

in significant underestimation of the overall ultimate load-carrying

capacity of the non-composite SCS sandwich panels under impact

loading.

5. Numerical results and discussion

5.1. Control panel

The predicted load and displacement time histories of the Con-

trol panel are compared to the experimental results in Fig. 5. It

shows that the model has the capacity to predict initial inertial ef-

fects during the instrumented impact testing and flexural response

of the Control panel quite closely. After the initial flexural capacity

was reached, the flexural resistance of the Control panel reduced

significantly to about 12 kN due to damage of the concrete core.

Then, the tensile membrane action started to develop in the steel

faceplates resulting in the significantly increased load-carrying

capacity of 356 kN. The model predicted the residual flexural

strength of 48 kN, which is higher than the experimental residual

strength of 36 kN. The model was able to predict the development

of the tensile membrane resistance and showed the peak tensile

membrane resistance of 384 kN. The maximum experimental dis-

placement of the panel was 200 mm, compared to the maximum

predicted displacement of 182 mm. The model underestimated

the maximum displacement by 9% compared to the experimental

result.Fig. 6 illustrates the damage of the concrete core observed after

the test. It shows that the concrete core at the mid-span was se-

verely damaged and large fragments of concrete fell out from the

panel. The model predicted extensive damage of the concrete core

at the impact zone and near the support similar to the experimen-

tal observation, as shown in Fig. 7. Fig. 8 shows the indentation of

the top steel faceplate observed after the test. Contour plot of the

Von Mises stresses on the top steel faceplate is shown in Fig. 9.

The indentation of the top faceplate is visualised by the elements

with a very high stress concentration forming a circular shape at

Fig. 5. A comparison between experimental and predicted results for the Control panel (a) load time histories, (b) displacement time histories.



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the mid-span. The maximum stress on the top faceplate was

623 MPa, exceeding the static yield stress of the mild steel due to

the strain rate effects included in the model. From the comparison

of the numerical and test data for the impact load, maximum dis-placements, physical damage of the concrete and steel plates, it can

be concluded that the finite element model is capable of capturing

the most important structural response characteristics of the Con-

trol panel.

An additional finite element model was generated to investigate

the effect of erosion of the concrete elements on the behaviour of

the Control panel under impact loading. Using the LS-DYNA con-

crete model 159 (CSCM), the erosion coefficient was assigned the

value of 1.0. This means that the concrete elements will be elimi-

nated from the model once they reach the failure strain. The sim-

ulation results were compared to the experimental results in

Fig. 10. It shows that the predicted residual flexural strength was

reduced and correlated better with the experimental residual flex-

ural strength of the panel. However, the predicted tensile mem-

brane resistance reduced to only 200 kN, and the maximum

predicted displacement was significantly lower than the experi-

mental displacement. This can be attributed to the elimination of

a large number of concrete elements from the model, causing a

shattering effect on the concrete core of the panel as shown in

Fig. 11. From these results, it can be concluded that modelling

the concrete core of SCS sandwich panels with the element erosion

option activated may result in significant underestimation of the

panels ultimate capacity and peak deformation. Hence, the erosion

algorithm is not recommended in the simulation of non-composite

steelconcretesteel panels under impact loads.

445.2. Stainless steel panel

44This panel used stainless steel faceplates. The stress strain rela-44tionship of the stainless steel shown in Fig. 3 was defined in the44model. The CowperSymonds coefficients of 100 and 10 were de-44fined to include the strain rate effects in the material model of44the stainless steel. As shown in Fig. 12, the model produced reason-

44able prediction of the inertial effects and flexural capacity of the 44Stainless steel panel. Between 0.01 s and 0.02 s of the response,4the model predicted a significantly higher flexural strength than4the experimental residual flexural strength of the panel for the4same reasons discussed above for the Control panel. However,4the peak tensile membrane resistance predicted by this model4was 380 kN, and it was only slightly higher than the experimental4peak resistance of 377 kN. The maximum displacement of the bot-4tom faceplate for this model was 162 mm, which is 10% lower than4the experimental displacement of 181 mm. In the test, the concrete4core of the Stainless steel panel was extensively damaged at the4mid-span and near the support, while the top faceplate was signif-46icantly indented by the impactor. These failure modes were similar46to the experimental observations of the Control panel shown in46Figs. 6 and 8. The contour plot of concrete damage distribution46and the Von Mises stresses for the model of the Stainless steel46panel were similar to the Control panel as shown in Figs. 7 and 9.

465.3. Panel with lightweight infill

46Lightweight concrete with a density of 1400 kg/m3 and the con-46crete compressive strength of 11 MPa was used as the infill for this46panel. The load and displacement time histories predicted by the46numerical model wre compared to the experimental results in4Fig. 13. The model predicted higher values for the initial flexural4capacity and the residual flexural strength than the experimental

Fig. 6. Damage of the concrete core of the Control panel after the impact test.

Fig. 7. Damage contour plot for the concrete core of the FE model for the Control panel.

Fig. 8. Indentation of the top faceplate of the Control panel observed in the test.



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2 results. As previously discussed, the erosion of the concrete ele-3 ments can reduce the predicted residual flexural strength, but it4 will also affect the tensile membrane resistance of the steel face-5 plate due to the shattering of the concrete elements. Therefore,6 the erosion formulation was not considered in the model of the7 lightweight concrete panel. The peak tensile membrane resistance8 predicted by the model was 358 kN compared to the experimental9 values of 333 kN. The maximum predicted displacement of the0 bottom faceplate was 174 mm. The model underestimated the

1 experimental maximum displacement of 196 mm by 11%. In the2 test, the physical damage of the lightweight concrete panel was

similar to the Control panel, except the concrete core of the light-

weight concrete panel experienced more severe damage. The

numerical model was capable to represent the main stages of the

response of the panel with lightweight concrete infill with reason-

able accuracy.

5.4. Panel with reinforced infill

In this panel, two layers of 4@50 mm wire mesh were used to

reinforce the concrete core. The wire meshes were modelled usingthe Hughes-Liu beam elements and their nodes were merged with

Fig. 9. Von Mises stress contour plot on the top faceplate of the FE of the Control panel.

Fig. 10. A comparison between experimental and predicted results using erosion formulation for the Control panel (a) load time histories, (b) displacement time histories.

Fig. 11. Shattering of the concrete core due to erosion of concrete elements.



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the concrete elements. Fig. 14 shows the comparison between the

experimental results and the simulation results.

From Fig. 14a, it can be observed that the numerical model pre-

dicted the initial response of the panel quite accurately. The flex-

ural strength predicted by the model was significantly higher

than that in the test. This could be attributed to the modelling of

the bond between the wire meshes and the concrete core by merg-

ing the nodes of the beam elements to the nodes of the concreteelements in the model. This full interaction between the concrete

50and the wire meshes ignored the slippage at the steel wire-con-50crete interface thus causing a higher flexural capacity. The peak50tensile membrane resistance predicted by the model was 294 kN,50which was slightly higher than the experimental result of50284 kN. The maximum displacement of the bottom faceplate pre-50dicted by the model was 168 mm. It was 8% lower than the exper-50imental displacement of 183 mm.

50In the test, it was observed that the concrete core at the impact50zone was crushed at the top and cracked at the bottom, as shown

Fig. 12. A comparison between experimental and predicted results for the stainless steel panel (a) load time histories, (b) displacement time histories.

Fig. 13. A comparison between experimental and predicted results for the lightweight core panel (a) load time histories, (b) displacement time histories.

Fig. 14. A comparison between experimental and predicted results for the Reinforced core panel (a) load time histories, (b) displacement time histories.



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0 in Fig. 15a. The top layer of the wire mesh at the mid-span buckled1 and the top steel faceplate was indented. There was some cracking2 at the flared zone, as shown in Fig. 15b. The damage contour plot of3 the concrete core was illustrated in Fig. 16. It shows that the model4 predicted similar damage of concrete core at the mid-span and5 flared zone with the experimental observations. The axial force6 contour plot of the wire meshes was illustrated in Fig. 17. It shows

7 that the top layer of wire mesh suffered local buckling at the im-8 pact zone and fractured near the support.

6. Conclusions

An extensive study on the dynamic response of non-composite

steelconcretesteel sandwich panels under condition of normal

impact by a falling mass has been undertaken. The numerical

simulations presented a full three-dimensional modelling of the

impact by a falling hammer from a height of 3 m. The numerical

results demonstrated reasonable correlation with the test results.They also revealed some differences in modelling the failure

Fig. 15. Damage of the concrete core of the Reinforced core panel (a) cracking, crushing of the concrete and buckling of top layer of the wire mesh, (b) cracking at the flared

zone.

Fig. 16. Damage contour plot for the concrete core for the FE model of the Reinforced core panel.

Fig. 17. Deformation and axial force distribution in the wire meshes for FE model of the Reinforced core panel.



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mechanisms between the steel plates and the concrete infill under

impact conditions. Several important modelling observations were

summarised below:

1. The contact surface algorithm: The LS-DYNA Automatic-

Surface-to-Surface contact algorithm which only considers the

friction resistance between interfaces was appropriate for the

non-composite steelconcretesteel panels since the chemicalbonding of the concrete has failed before the testing. However,

the assumption of a full bond between the wire meshes and the

concrete core in the model with reinforced concrete infill may

overestimate the flexural capacity of the panel.

2. The strain rate effect: The strain rate effect of the steel was con-

sidered by defining the CowperSymonds coefficients. These

coefficients increased the static yield stress of the steels. The

maximum stress achieved by the steel plates in the simulation

was higher than 600 MPa. It was found that the hourglass

energy exceeded 50% of the internal energy of the concrete core

when the strain rate effects of the concrete were considered.

Therefore the strain rate effects of concrete were ignored to

minimise the hourglass energy in this study.

3. Erosion formulation of concrete elements: The concrete ele-

ments can be eliminated once they exceeded the strain thresh-

old defined in the LS-DYNA concrete material model CSCM

Concrete (Mat 159). The erosion of the concrete elements could

improve the prediction of the residual flexural strength of the

panel. On the other hand, the erosion formulation was not rec-

ommended in this study because it may result in significant

underestimation of the tensile membrane resistance of the

non-composite axially-restrained sandwich panels under large

deformation.

Acknowledgements

This research was supported under Australian Research Coun-

cils Discovery Projects funding scheme (Project No. DP0879733),this support is gratefully acknowledged. The authors would like

56to thank senior technical staffs, Mr. Alan Grant and Mr. Ian Bridge56for their assistance in conducting the experiments.

56References

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60

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