numerical simulation of hyper velocity impact on cfrp-honeycomb

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International Journal of Impact Engineering 33 (2006) 703–712

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Numerical simulation of hypervelocity impact onCFRP/Al HC SP spacecraft structures causing penetration

and fragment ejection

S. Ryana,b,�, F. Schaefera, W. Riedela

aFhG-Ernst-Mach-Institut, Eckerstr. 4, D-79104, Freiburg, GermanybSchool of Aerospace, Mechanical & Manufacturing Engineering, RMIT University, GPO Box 2476V, Melbourne, Australia

Available online 1 November 2006

Abstract

A representative carbon fiber reinforced plastic/aluminum honeycomb sandwich panel (CFRP/Al HC SP) spacecraft

structure has been modeled in the hydrocode AUTODYN using the state-of-the-art ADAMMOmaterial model [Riedel W,

Harwick W, White D, Clegg R. Advanced material damage models for numerical simulation codes. ESA CR(P) 4397,

2003] to study the performance of the structure during impact events that cause perforation and fragment ejection. A new

procedure combining a series of existing theoretical methods has been developed and applied to derive a full set of coarse

material data. The data set has been implemented in AUTODYN, and the results of the numerical simulation have been

compared to experimental impact test data. For impact tests performed near the structural ballistic limit, quantitatively

accurate results were obtained over a range of impact velocities and angles. A further increase in the projectile size resulted

in significant destruction of the sandwich panel front face-sheet and diversion from the experimental damage

measurements. Inspection of the numerical model has shown non-localized propagation of inter-laminar delaminations,

possibly caused by an under-prediction of the laminate dynamic inter-laminar tensile strength. The effects of the

delamination propagation occur over an extended time scale and were not found to affect the state and trends of the

fragment cloud ejected into the satellite interior. Accordingly, experimental trends of fragment cloud dispersion have been

qualitatively reproduced.

r 2006 Elsevier Ltd. All rights reserved.

Keywords: Space debris; Hypervelocity; Hydrocode; CFRP; Simulation

1. Introduction

A recent experimental hypervelocity impact (HVI) test campaign performed at the Ernst-Mach-Institute(EMI) has investigated the behavior of carbon fiber reinforced plastic/aluminum honeycomb sandwich panel(CFRP/Al HC SP) spacecraft structures and the subsequent damages to representative component set-upslocated behind these structural walls [2]. This type of structure is common in satellites (e.g. GOCE, Radarsat2,

ee front matter r 2006 Elsevier Ltd. All rights reserved.

mpeng.2006.09.072

ing author. FhG-Ernst-Mach-Institut, Eckerstr. 4, D-79104, Freiburg, Germany. Tel.: +49 761 2714 402;

14 316.

ess: [email protected] (S. Ryan).

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dx.doi.org/10.1016/j.ijimpeng.2006.09.072

mailto:[email protected]

ARTICLE IN PRESSS. Ryan et al. / International Journal of Impact Engineering 33 (2006) 703–712704

Herschel/Planck, Integral, BeppoSax, etc.), and given the increasing population of orbital debris in keyoperational orbits and the subsequent increased risk of micrometeoroid/orbital debris (M/OD) impacts, it isimportant that the vulnerability of these structures to HVI be thoroughly investigated. Prior to theexperimental study in [2], the amount of HVI test data on CFRP/Al HC SP structures was very limited [3–6].

Given the high costs associated with experimental impact testing of composite sandwich panel structuresand the large variety of configurations in use, the utilization of hydrocode software packages in support ofexperimental campaigns is of interest. Previous hydrocode numerical studies (e.g. [7–9]) have investigated theperformance of spacecraft structures subject to HVI, and have shown good correspondence with experimentalresults, allowing extrapolation of experimental data to conditions beyond the reachable limits of experimentalfacilities (i.e. impact velocities). Recent developments [1] in the modeling of fiber-reinforced compositematerials for implementation in the hydrocode AUTODYN allow orthotropic constitutive behavior, non-linear equation of state, orthotropic non-linear hardening, and individual material plane interactive failureinitiation criteria to be described, improving the existing hydrocode capabilities in reproducing the behavior ofcomposite materials during HVI. It was shown in [1] that the implementation of these composite-specificproperties can allow damage caused by HVI of aluminum fragments on composite materials to bequantitatively reproduced. Development and validation of this mesomechanic material model was performedusing an Aramid weave/epoxy composite; however, the possible applicability of this technique to the modelingof high-strength composite materials (e.g. CFRP) has been demonstrated through preliminary studies [1,10]using limited experimentally derived material data.

A composite material is described in the ‘‘Advanced material DAMage Models (ADAMMO) for Numericalsimulation codes’’ model by over 60 material coefficients that define: elastic performance, initial yield surface,reference hardening curve, and failure stresses and energies. If full material data sets are not available, anumber of theoretical techniques can be applied in series to derive a full set of coarse material data forcomposite laminates consisting of multiple uni-directional (u.d.) plies if the u.d. ply elastic and failureproperties (stiffness, Poisson’s ratio, strength) are known, along with the laminate stacking sequence. In thisstudy, CFRP face-sheets (fs) of the Radarsat-2 [11]7Z platform CFRP Al H/C SP have been considered.

2. Derivation of coarse CFRP material data

The specifications of the Radarsat-2 7Z CFRP laminate are given in Table 1.Without the outer fabric layer (FL01), the laminate is symmetrical and quasi-isotropic. The contribution of

the fabric layer is primarily for protection and resistance in the construction phase, i.e. the laminatemechanical properties are primarily dependant on the eight u.d. plies (FL02). As such, it has been assumedthat the contribution of the fabric layer to laminate performance can be conservatively represented byremoving the fabric from the lay-up and subsequently increasing the thickness of each FL02 u.d. ply toconserve the nominal laminate thickness. Therefore, derivation of the laminate mechanical properties isperformed for an idealized quasi-isotropic structure consisting of 8 FL02 u.d. plies, each 0.14125mm thick.The properties of the FL02 u.d. ply are given in Table 2.

A procedure which combines a series of theoretical methods was developed and applied to derive thelaminate properties and material data required for implementation into the ADAMMO material model fromthe limited manufacturer-supplied data (Table 2). The procedure is intended for application whenexperimental material data is not available. Material constants not derived from the theoretical methods(e.g. delamination energy) were estimated based on existing ADAMMO composite material data [1,10]. Thesteps of the derivation procedure are shown in Fig. 1.

Table 1

Specifications of the Radarsat-2 7Z panel CFRP laminate

Type FL01 ¼ HMF196/34 T300-1k fabric

FL02 ¼ HYE 4934C K139 u.d.

Stacking sequence FL01 01, FL02 01, +451, 901, �451, �451, 901, +451, 01

Thickness 1.13mm

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Table 2

Manufacturer-supplied properties of the HYE 4934C K139 u.d. lamina

Parameter Nomenclature Value

Longitudinal tensile modulus E_tx (MPa) 399700

Longitudinal compression modulus E_ty (MPa) 8293

Shear modulus G_xy (MPa) 3550

Longitudinal tensile strength Uts_x (MPa) 1457.1

Transverse tensile strength Uts_y (MPa) 83

Longitudinal compression strength Ucs_x (MPa) 293.8

Transverse compression strength Ucs_y (MPa) 83

Shear strength Us_xy (MPa) 40.4

Inter-laminar shear strength ILSS (MPa) 35.1

Ply thickness t (mm) 0.125

Ply density r (kg/m3) 1800

Classical LaminateTheory (Tsai-Wufailure criterion)[13]

σ-ε (11 & 22)Strength1,2

Hooke’s Law foran OrthotropicComposite [12]

E1, E2, E3G12,G23,G21ν12, ν23, ν31

Rankine-Huginotequations for theshock-jumpcondition [1]

Mie-Gruneisencoefficients(A1,A2,A3,Γ,T1,T2)

Chen’s QuadraticYield Function[14]

Plasticity parameters(a11..a66)Eff. σ-ε master curve

Fig. 1. Flow-chart describing the coarse composite material data derivation procedure [1, 12–14].

S. Ryan et al. / International Journal of Impact Engineering 33 (2006) 703–712 705

In previous studies (e.g. [7]) it has been noted that classical laminate theory (CLT) cannot be used to analyzethe 3D characteristics of a moderately thick laminate. This statement is based on the assumption in CLT thatthe transverse shear is negligible, i.e. under deformation; the transverse shear strains are zero. The thickness ofthe Radarsat CFRP laminate under consideration in this study is 1.13mm, thus the assumption of plane stressand negligible transverse shear is valid. However, care should be taken when applying this technique to thickerlaminates.

An overview of the key parameters of the derived material model is given in Table 3. The remainder ofthe material data can be determined by following the procedure outlined in Fig. 1. In this material data set, the33-direction is defined as the through-thickness.

3. Set-up

3.1. Experimental set-up

Experiments described in this paper were performed at EMI’s LGG facilities. In the experiments, thestructure panel was mounted on a frame holder that is connected to the witness plate (WP) spaced at 100mmvia four threaded rods. Further details of the experimental set-up can be found in [2].

3.2. Numerical set-up

Considering the large variation in projectile diameters, and the subsequent induced damages, investigatedin this study, two numerical models were constructed: the ‘‘fragmentation model’’ for examining thefragmentation cloud produced as a result of penetration; and the ‘‘ballistic limit (BL) model’’, used forvalidation of the numerical model through comparison with experimental structure ballistic limit tests.

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Table 3

Overview of derived ADAMMO material data set for HYE 4934C K139

Description Parameter Data

Equation of state: orthotropic

Reference density rREF (g/cm3) 1.800E+00

Longitudinal tensile modulus E11 (kPa) 1.389E+08

Transverse tensile modulus E22 (kPa) 1.389E+08

Through-thickness tensile modulus E33 (kPa) 1.085E+07

In-plane Poisson’s ratio n12 (—) 3.291E–01

Out-of-plane Poisson’s ratio n23 (—) 3.250E–01

Out-of-plane Poisson’s ratio n31 (—) 2.540E–02

In-plane shear modulus G12 (kPa) 5.223E+07

Out-of-plane shear modulus G23 (kPa) 3.645E+06

Out-of-plane shear modulus G31 (kPa) 3.645E+06

Strength model: orthotropic yield

Plasticity parameter (normal stresses 11) A11 (—) 1.000E+00



Plasticity parameter (interaction between normal stresses 11 and 22) A12 (—) 0.000E+00



Plasticity parameter (shear stresses 23) A44 (—) 5.633E+01



Failure model: orthotropic softening

Longitudinal tensile failure stress 11 (kPa) 3.726E+05

Transverse tensile failure stress 22 (kPa) 3.726E+05

Through-thickness tensile failure stress 33 (kPa) 8.300E+04

Maximum in-plane shear stress 12 (kPa) 2.634E+05

Maximum out-of-plane shear stress 23 (kPa) 3.510E+04

Maximum out-of-plane shear stress 31 (kPa) 3.510E+04

Table 4

Results of the ballistic limit experiments (EB) and numerical simulations (SB)

Test EMI no. V (km/s) dp (mm) a (deg.) Exp Result, P/NP Front FS Rear FS

dh (mm) DSP (mm) dh (mm) DSP (mm)

EB-1 119 6.24 1.0 0 P 4.3� 4.2 8.3� 8.9 2.2� 3.9 5.9� 7.7

SB-1 — 6.24 0.99 0 P 3.7� 3.7 6.4� 6.9 3.7� 4.7 4.1� 8.2

EB-2 4580 6.12 2.0 45 P 7.5� 8.0 10.6� 10.1 o1mm 4.0� 3.2

SB-2 — 6.12 1.98 45 P 8.5� 10.3 9.5� 12.2 5.2� 4.8 5.1� 5.9

EB-3 107 1.40 1.25 0 NP 2.4� 2.2 5.0� 4.7 No damage

SB-3 — 1.40 1.27 0 NP 2.4� 2.2 3.0� 2.5 No damage

S. Ryan et al. / International Journal of Impact Engineering 33 (2006) 703–712706

The fragmentation configuration was modeled in single-axis symmetry with an effective lateral extension of100� 100mm to completely contain the deformed areas (even in the case of oblique impact). The front face-sheet was meshed entirely using SPH particles. For sufficient resolution in the projectile diameter and throughthe thickness of the face-sheets, 0.2825mm diameter SPH nodes were used. This corresponds to four SPHnodes through the thickness of the face-sheets. It was not possible in all cases to mesh the SPH objects with aconsistent-SPH particle size while maintaining the nominal diameter of the projectile. Thus, slight variationsbetween the experimental and simulated projectile diameters can be seen in Tables 4 and 5. The rear face-sheetwas discretized using Lagrangian volume elements with a central grading bias. The central impact region on

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Table 5

Comparison of the fragment ejection experiments (EF) and simulations (SF)

Test EMI no. V (km/s) dp (mm) a (deg.) Front FS Rear FS b Witness plate

dh (mm) DSP (mm) dh (mm) DSP (mm) Int (deg.) Ext (deg.) DWP99 (mm) DDUST (mm)

EF-1 4635 6.64 3.00 0 8.4� 8.8 11� 10 5.8� 6.4 15� 10 �2.2 55.6 49� 49 113� 110

SF-1 — 6.70 3.108 0 8.1� 8.4 39� 48 8.4� 4.2 11� 12 �2.2 43.2 0� 0 96� 75

EF-2 4671 2.60 4.00 0 7.8� 8.2 9� 12 4.5� 5.4 15� 19 �3.4 26.5 33� 13 50� 54

SF-2 — 2.50 3.955 0 17.3� 18.1 20� 25 12.2� 11.4 18� 13 �6.5 49.0 44� 28 104� 102

EF-3 4633 6.71 5.00 0 11.3� 11.7 14� 14 53.6� 38.8 73� 52 37.0 52.6 72� 73 137� 153

SF-3 — 6.70 5.085 0 27.5� 33.0 102� 108 11.7� 12.4 22� 27 �19.9 59.8 79� 67 146� 108

Fig. 2. Propagation of the fragment cloud. Top: high speed framing camera images from exp 4671, 4.0mm, 2.60 km/s, 01; Bottom: initial

images from numerical simulation of Exp 4671 and further extrapolation of the fragment cloud.


the rear face-sheet was described using SPH particles confined to 50� 50mm and coupled to the surroundingLagrangian elements.

The BL model was confined to a lateral extension of 25� 25mm, using coupled SPH/Lagrangiandiscretisation for both face-sheets. SPH particles of 0.14125mm diameter provided sufficiently high resolutionvia eight nodes through the face-sheet thickness for the sensitive BL validation cases.

The HC consisted of Al 5056 foils (thickness 0.0254mm) regularly folded at 7601 and pin-joined togetherto form hexagonal cells. The HC material was described by a linear equation of state and the Johnson-Cookstrength model. The impactor in all simulations was an Al 2024-T351 sphere. The material was described usinga shock equation of state and the Johnson-Cook strength model.

Simulation of the fragment cloud impacting on the WP was not performed in order to minimize runningtimes. Instead, a user subroutine which recorded the properties of fragments passing through a virtualmembrane located 10mm behind the structural wall (SP) rear face-sheet was implemented. The fragment datawas extrapolated to a point in space representing the experimental WP (extrapolation shown in Fig. 2). Thismethod is unable to identically reproduce the composition of the fragment cloud due to the resolution of theSPH nodes. Additionally, reproduction of CFRP dust deposits on the WP is difficult as the CFRP dust doesnot crater the WP surface in an impact experiment [2]. As a result, it is not possible to adjust the SPHresolution to the expected individual crater size, nor is it possible to account for ejecta particles, which do notadhere to the WP surface. To account for this, the low-density extremities of CFRP deposits on the WP wereexcluded from the numerical dust extension measurements.

As seen in Fig. 2, the extremities of the fragment cloud cannot be precisely captured in the high-speedshadowgraphs as a result of the low density and size of the individual CFRP fragments. However, it can beseen that the visible high-density ejecta cone is reproduced in the simulation extrapolation.


4. Uncertainties and scatter

In the numerical reproduction of HVI, there are a number of uncertainties which affect the ability to exactlyreplicate the impact experiment, thus complicating direct comparison of the experimental and numericalresults. Amongst the structures tested in [2], significant variation of the structural components from thenominal specifications was found (up to +33% in the CFRP face-sheet thicknesses). Furthermore, distortionsand locally inhomogeneous sections may exist in the honeycomb core (as shown in Fig. 3). In all simulationsthe numerical model was constructed to nominal specifications.

The global damage produced by HVI on a test sample is subject to significant ‘scatter’. The scatter may bedue to the afore-mentioned variation in the target parameters. As such, two experiments can be performed atnear-identical impact conditions on nominally identical target configurations and the results can varysignificantly. Evidence of experimental scatter is given in Fig. 4 for two tests performed at EMI on theRadarsat-2 7Z SP structure (a ¼ 01). The effect of these uncertainties may account for �725% variation incomparison of numerical and experimental results.

5. Preliminary validation of numerical model

Validation of the numerical model was performed using both ballistic limit and fragment ejection-inducingimpact tests over a range of projectile sizes, impact velocity regimes and angles. In total, three ballistic limitand three fragmentation experiments were used to validate the numerical model. Due to the extremely thin,brittle nature of the CFRP face-sheets, differentiating between a detached spall and perforation case isdifficult. As such, specification between detached spallation and perforation was not made in this study.Failure is defined as the ejection of any material in the half space behind the SP.

Fig. 3. Left: distorted SP HC core (Rosetta CFRP Al HC SP, sample DLR-A/1), right: ideal numerical model HC core.

Fig. 4. Face-sheet damage experimental scatter for EMI exp 4633 (dp ¼ 5.0mm, V ¼ 6.75 km/s, a ¼ 01) and 4653 (5.0mm, 6.69 km/s, 01).

Left to right: EMI No. 4633 front fs, 4653 front, 4633 rear, 4653 rear.

ARTICLE IN PRESSS. Ryan et al. / International Journal of Impact Engineering 33 (2006) 703–712 709

5.1. BL validation cases

Validity of the BL simulations was assessed via comparison of the overall experimental result (i.e.perforation/no perforation) and measurement of the damage zones in both SP face-sheets: ‘clear perforationhole’ (dh) and ‘detached spall’ (DSP), shown in Fig. 5.

The results of the BL validation simulations and experiments [2] are given in Table 4.It can be seen in Table 4 that the overall result (perforation/no perforation) of all three BL validation cases

has been reproduced in the simulations. The two perforating experiments (EB-1 and EB-2) displayed resultsminimally above the BL (as can be seen in the small perforation hole diameters of the rear face-sheet). That thenumerical model was able to reproduce these sensitive perforation results, for both normal and oblique impactangles, while also reproducing the no-perforation result of experiment EB-3 confirms the applicability of thematerial data. The damage measurements in the vast majority of cases are reproduced to a reasonable degreeof accuracy, with the exception of the rear face-sheet in SB-2. In this case the perforation hole is significantlylarger than in the experimental case (5.2� 4.8 vs. o1mm); however, the spall area is reproduced morereasonably (5.1� 5.9 vs. 4.0� 3.2).

5.2. Fragment-ejection validation cases

Validation of the fragmentation model was assessed via: dispersion of the fragment cloud both internally,i.e. inside the honeycomb (bint), and externally (bext); SP face-sheet damages (dh and DSP), and ejecta depositson the witness plate (DDUST, DWP99). The witness plate deposits are characterized into two groups: dustdeposits from carbonized CFRP fragments (DDUST), and micro-craters from Al fragments (DWP99). Allfragmentation validation measurement techniques are shown in Figs. 5 and 6.

For the fragmentation validation case 1 (EF-1/SF-1) the sandwich panel face-sheet clear holemeasurements, dispersion angles and witness plate dust deposit measurements all agree to within 30% ofthe experimental results. The front fs spall is significantly over-predicted in the numerical simulation, and noaluminum deposits are predicted on the WP. In the experiment, the aluminum deposits are spread very lightly(i.e. minimal volume) over the 49� 49mm area and may possibly be reproduced in the numerical result if thesimulation was executed over a longer time-frame. Another possible cause of lower aluminum ejected mass inthe simulation is the application of an erosion algorithm required in modeling the honeycomb core shellelements with a Lagrangian discretization scheme. Following severe distortion of the shell elements, the cell iseroded and the cell mass is withdrawn from the numerical model. Thus all HC cells significantly distorted (i.e.in the impact locality) do not contribute to the Al ejecta, which may not be a true representation of theexperiments.

Fig. 5. Experimental and numerical damage measurements on CFRP front (left) and rear (right) face-sheets from BL validation case 1

(EB-1 and SB-1 respectively) (1.0mm, 6.24 km/s, 01).

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Fig. 7. Comparison of front CFRP face-sheet impact damage locality of EF-3 and SF-3 (5mm, 6.7 km/s, 01), dh,front marked.

βint/2

dh,front

dh,rear

tHC

dhc

βext/2 βext/2

Ddust

WP

DW

P99

,2

DDUST,1

DWP99,1

DD

US

T,2

WP

Fig. 6. Validation measures in fragmentation configuration. Left: internal dispersion angle bint; center: external dispersion angle bext; right:witness plate deposit measurements.

S. Ryan et al. / International Journal of Impact Engineering 33 (2006) 703–712710

In fragmentation simulations 2 (SF-2) and 3 (SF-3), damage measurements deviate considerably fromexperimental results, well beyond the previously defined acceptable scatter limits. In both cases the numericalmodel’s front surface damage values are much larger than seen in the experimental samples. For the rearCFRP face-sheets, SF-2 under-predicts the damage zone, while SF-3 shows an over-prediction of the clearhole diameter.

In general it can be noted that an increase in the projectile diameter corresponds to an increase in theinaccuracy of numerical model damage values. For SF-1, simulated normal impact of a 3.0mm projectile athigh velocity, the damage measurements agree reasonably well with the experiment. An increase in theprojectile size to 4.0mm (SF-2) results in significant (�200%) over-prediction of the damage measurements forboth SP facesheets and WP deposits. Following a further increase in the projectile diameter to 5.0mm (SF-3),the deviation of the simulation damage values from the experimental measurements was even morepronounced: front surface measurements were over-predicted by up to 600% while rear surface measurementswere under-predicted by up to 230%.

From inspection of the simulation model, some phenomenological discrepancies are noted. In Fig. 7, it isapparent that the damage in the front face-sheet of the simulation model is not limited in the vicinity of theperforation hole, instead the damage propagates through an extended area of the face-sheet, leading tothe excessively large front face-sheet damage measurements. In simulations performed for smaller projectilediameters, i.e. BL validation simulations, this response was not seen (Fig. 8 bottom). For the BL andfragmentation cases, the only differences in the two numerical models were the size of the SPH particles andthe lateral extension of the SPH-discretized zones. It is not expected that the lateral extension would affect themodel, given that the damage was contained in the SPH zone in both models.

For the BL simulations, there were eight SPH nodes through the thickness of the CFRP face-sheets,compared to four nodes in the fragmentation model. However, for SF-1 the low through-thickness resolutionwas used and the results obtained were reasonably accurate. It is likely, therefore, that the excessivedelamination predictions are a result of the material description.

One possible cause of the through-thickness expansion clearly shown in Fig. 8 is superposition of the releasewaves reflected from the upper and lower surfaces of the face-sheet. For larger projectile diameters, the shockduration is longer and therefore the release takes place at a later stage of the perforation process. In definingthe Radarsat CFRP material model, the static laminate through-thickness tensile strength is used (effectively

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Fig. 8. Localized front face-sheet damage of ballistic limit validation case 1 (SB-1) (1mm, 6.24 km/s, 01).

0

50

100

150

200

250

2 6 10

Impact velocity [km/s]

WP

du

st d

isp

ersi

on

[mm

]

Dust (dp=4) Dust (dp=3) Exp dust (dp=4)

4 8

Fig. 9. The effect of impact velocity on the dispersion of WP CFRP dust deposits. All experiments and simulations at 01 impact.


the matrix tensile strength). Further investigation into the dynamic strength of these materials and themechanisms leading to inter-lamina failure or delamination is suggested.

It should be noted that for the primary material model developed and validated in [1] (Kevlar-Epoxy),validation was achieved via impact of fragment clouds (i.e. multiple small projectiles) at velocities below7.0 km/s. As seen in Table 5, the performance of the Radarsat-2 CFRP numerical model for impact offragments (i.e. impact on the rear fs) does not experience the damage extension anomalies seen in the front fs,the problem is noted only in the case of large single-projectile impact on the CFRP laminate. Additionally, thepreliminary CFRP material model developed in [1] and applied in [10] considered impact of projectiles up to amaximum of 3mm in diameter (4.49 km/s, 01), conditions at which no anomalies have been measured in thisnumerical study. Subsequently, the validity of the ADAMMO material model applied to impact cases withlarge projectile diameter:plate thickness (dp:tCFRP) ratios and the subsequent high-amplitude, long durationimpact shock and release effects is not confirmed. The performance of the SPH algorithm in such cases shouldalso be investigated further.

The time scale at which the front fs delamination occurs is significantly large when compared to thepropagation of the fragment cloud through the SP structure and into the satellite interior. In Table 5 it can benoted that the WP deposit trends (i.e. dispersion with respect to impact conditions) are reproduced. A series ofdesigned to investigate the effect of impact velocity on fragment cloud dispersion were performed in parallel withan additional numerical campaign. Details of the experiments and numerical simulations are given in Fig. 9.

It can be seen in Fig. 9 that the extensive front surface delaminations highlighted in SF-2 and SF-3 do not affectthe phenomenological dispersion performance of the fragment cloud following ejection into the satellite interior.As such, it is considered reasonable to apply the numerical model in future fragment cloud predictive studies.

6. Conclusions

A new technique for calculating a full set of coarse CFRP material data for implementation with the state-of-the-art ADAMMO [1] material model in the hydrocode AUTODYN has been developed based on a series of


theoretical methods. This procedure enables predictive simulation campaigns to be performed for highlycomplex composite materials in the case when experimentally derived material data is not available. Applicationof this procedure has produced a coarse set of material data for the CFRP laminate used as face-sheets on theRadarsat-2 7Z structural sandwich panel. In combination with an experimental impact campaign, it has beenshown the subsequent numerical model is able to quantitatively reproduce the ballistic performance for impacttests near the ballistic limit of the structure over a range of impact velocities and angles.

Following validation of the numerical model through BL simulations, the impacting projectile size wasincreased to examine the behavior of the fragment cloud ejected from the SP structure. In these simulations,problems in the numerical behavior of the CFRP front facesheet have been observed. The numericalsimulations significantly over-predicted damage measurements’, suggesting the validity of the numerical modelis not assured for HVI exceeding the structure’s BL. The numerical models show disruption of the face-sheet isnot restricted to the impact-locality; instead inter-laminar delamination propagates through large areas ofthe face-sheet, not reflecting the experimental results. It is proposed that a possible cause of the damage-extension is an under-prediction of the dynamic inter-laminar tensile strength of the material. This should beinvestigated in further studies. Furthermore, it was noted that the ADAMMO material model has not beenvalidated for impact of large projectiles (dp40.2mm).

The dispersion of CFRP dust deposits on a WP offset 100mm was studied to determine the effect of impactvelocity on ejecta dispersion. In comparison with experimental results, it was determined that an increase inimpact velocity while maintaining projectile diameter and impact angle resulted in an increase in the lateraldispersion of the CFRP fragment cloud. Simulations performed with two different projectile diameters (3and4mm) were able to qualitatively reproduce these trends, suggesting the problems encountered with frontCFRP surface delamination occur over a time scale sufficiently large enough (compared to the perforationprocess and fragment propagation through the SP), that the resulting fragment cloud ejected into the satelliteinterior is not significantly affected.

Acknowledgments

All experimental impact tests have been performed as part of ESA Contract 16721/02/NL/CK ‘‘CompositeMaterials Impact Damage Analysis’’.

Reference

[1] Riedel W, Harwick W, White D, Clegg R. Advanced material damage models for numerical simulation codes. ESA CR(P) 4397,

2003.

[2] Schafer F, Destefanis R, Ryan S, Riedel W, Lambert M. Hypervelocity impact testing of CFRP/Al honeycomb satellite structures. In:

Proceedings of the fourth European conference on space debris, ESA, Darmstadt, 2005.

[3] Lambert M. Hypervelocity impacts and damage laws. Adv Spacecraft Res 1997;19(2):369–78.

[4] Lambert M, Schaefer F, Geyer T. Impact damage on sandwich panels and multi-layer insulation. Int J Impact Eng 2001;26:369–80.

[5] Frost C, Rodriguez P. AXAF Hypervelocity impact test results. Proceedings of the second European conference on space debris, ESA

SP-393, Darmstadt, 1997.

[6] Taylor E, Herbert M, Vaughan B, MacDonnell J. Hypervelocity impact on carbon fiber reinforced plastic/aluminum honeycomb:

comparison with whipple bumper shields. Int J Impact Eng 1999;23:883–93.

[7] White D, Taylor E, Clegg R. Numerical simulation and experimental characterization of direct hypervelocity impact on a spacecraft

hybrid carbon fiber/kevlar composite structure. Int J Impact Eng 2003;29:779–90.

[8] Taylor E, Glanville J, Clegg R, Turner R. Hypervelocity impact on spacecraft honeycomb: hydrocode simulation and damage laws.

Int J Impact Eng 2003;29:691–702.

[9] Hayhurst C, Livingstone I, Clegg R, Destefanis R, Faraud M. Ballistic limit evaluation of advanced shielding using numerical

simulations. Int J Impact Eng 2003;26:309–20.

[10] Ryan S, Riedel W, Schafer F. Numerical study of hypervelocity space debris impacts on CFRP/Al honeycomb spacecraft structures.

In: Proceedings of the 55th IAC, IAF, Vancouver, 2004.

[11] Radarsat International. Radarsat 2 Mission Overview. MacDonald Dettwiler, viewed 2 June 2005, /http://www.radarsat2.info/

rs2_satellite/overview.aspS.

[12] Berthelot J. Composite materials: mechanical behavior and structural analysis. New York: Springer; 1999.

[13] Daniel I, Ishai O. Engineering mechanics of composite materials. New York: Oxford University Press; 1994.

[14] Chen J, Allahdadi F, Sun C. A quadratic yield function for fiber-reinforced composites. J Compos Mater 1997;31(8):788–811.

http://www.radarsat2.info/rs2_satellite/overview.asp

http://www.radarsat2.info/rs2_satellite/overview.asp

numerical simulation of hyper velocity impact on cfrp-honeycomb

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