EXPERIMENTAL AND NUMERICAL STUDY OF THE ...

EXPERIMENTAL AND NUMERICAL STUDY OF THE SHAPED CHARGE JET PERFORATION AGAINST CONCRETE TARGET A.D. Resnyansky1, G. Katselis1 and A.E. Wildegger-Gaissmaier2 1 Weapons Systems Division, Defence Science and Technology Organisation, Edinburgh SA 5111, Australia 2 Science Policy Division, Defence Science and Technology Organisation, Canberra ACT 2600, Australia This paper presents experiments and LS-DYNA hydrocode analysis of the Shaped Charge jet penetration of concrete targets. The effect of liner angle, charge height and charge stand-off distance are considered. Afterimpact damage characteristics from experiments on concrete targets are compared with LS-DYNA simulations. The simulation demonstrates the importance of the liner angle and the stand-off charge-target distance to the shape and velocity parameters of the jet, and in turn, to the damage parameters of the target. INTRODUCTION The majority of man-portable direct fire guided hard target defeat weapons are optimised for armour penetration. Most of the weapons utilise shaped charge (SC) technology employing tandem shaped charge warheads. It is expected that the use of these weapons against other targets like concrete structures (bunkers, buildings, and field fortification) is less effective. To defeat armour the hypervelocity SC jet penetrates even reactive armour and creates behind armour debris, which can ignite and penetrate materials/components inside the vehicles and contribute to crew incapacitation and injuries. SC jet penetration of brittle building materials such as concrete is governed by different mechanisms. SC weapons optimised for building targets often use a SC jet as a precursor projectile with a follow-up grenade. Therefore, an important factor for this case, in contrast to the ductile armour targets, is the borehole diameter in a concrete target in order to allow the follow-up grenade to pass through the concrete wall. One potential way to increase the diameter is to increase the charge mass as reported in [1]. However, the weight penalty associated with this method may not be an option for man-portable weapons.

This paper presents experiments and numerical analysis of the SC jet penetration in concrete targets. Flash X-ray photography is used for observation of the SC liner collapse. In the previous paper [2] we reported calculations of the liner collapse using a hydrodynamic elasto-plastic model that resulted in a significant deviation of the liner shape from that observed in the experiment. The present paper employs a rate sensitive model providing better agreement with the experiment. In the present study the shaped charges utilised aluminium liners. Besides the liner angle, the influence of the charge height and charge-target stand-off distance on the jet parameters is also considered. Hydrocode analysis of the target response employs a model, which has been published elsewhere [3]. This model has previously been used for simulation of damage to glass [3] and concrete [2] targets. The model is implemented in a commercial version of LS-DYNA3D and is being employed with the Eulerian option. Modelling results have been compared with the experimental data, when varying parameters of the problem. The analysis demonstrates that the important factors affecting the resulting borehole diameter achieved in the target are the liner angle and the stand-off distance between the SC charge and target. EXPERIMENT The experiments using concrete targets were aimed at a study of the influence of the SC parameters and the stand-off distance between the charge and target on target damage.

FIGURE 1. Schematic of the experimental set-up (axial cross-section).

The experimental set-up includes: 1) a SC conical liner of thickness τ, made of aluminium with the internal liner angle θ (liner L in the schematic in Fig. 1); 2) a cylindrical profiled charge of high explosive HE (Composition B) with the maximum height h (see Fig. 1); 3) a booster with detonator B; 4) an aluminium casing C of 3mm-thickness, containing the charge and the liner; and 5) a target T of thickness ∆ – (300mm in the experiments), located

at the stand-off distance D from the shaped charge. The charge diameter is 76mm; the target is approximately 70cm in diameter and it is encased in a steel shell S. Several tests have been conducted without targets in order to obtain the Flash X-Ray images of the SC jet for observation of the liner collapse and validation of the modelling results. The experiments using concrete targets varied the following parameters: i) charge height (h); ii) liner angle (θ); iii) liner thickness (τ); and iv) stand-off distance (D). Results of the tests are summarised below in Table 1. TABLE 1. SUMMARY OF THE SC PENETRATION TESTS AGAINST CONCRETE TARGETS Ntest h(mm) θ (˚) τ(mm) D(mm) Fd(mm) Fs(mm) 1 46 90 3 127 44 280 2 46 90 4 127 45 270 3 46 90 4 80 51 420 4 36 100 4 128 47 270 5 36 100 3 153 49 270 6 36 110 3 93 22 455 7 36 110 3 153 26 450 8 36 110 4 153 24 350

Damage of the targets appeared as an extensive front scabbing, a localised borehole, and a rear scabbing, which can be extensive as well. The most relevant characteristic to the current study is the minimal borehole diameter, denoted by Fd, and shown in Table 1 for each of the tests. General damage was assessed by measuring the maximum visible area of front scabbing. This was approximated by an average diameter of the affected zone at the front surface of the target; denoted by Fs and shown in Table 1.

FIGURE 2. After-impact photographs of residual damage of the concrete targets in tests 1 ((a) and (b)) and 3 ((c) and (d)). Views of the front surfaces ((a) and (c)) and the rear surfaces ((b) and (d)).

Photographs in Fig. 2 demonstrate typical patterns of the damage for tests 1 and 3 (Ntest is the test number in Table 1). Traces of burning are noticeable on the front surfaces

of the targets (Figs. 2(a) and (c)). As it is seen in Figs. 2(b) and (d) the scabbing at the rear of the targets is significant. A narrow borehole is the principal damage in the middle of the target. This damage classification is typical for the majority of the tests listed in Table 1. Exemptions are tests 6 and 7 where the localised borehole diameter changed stepwise with penetration depth. In those cases Table 1 states the minimum borehole diameter. Results in Table 1 indicate that in general lower liner angles (90° and 100°), as seen in tests 3 and 5 respectively, achieve larger borehole diameters. For the 90° cases (tests 1-3) a decreasing stand-off distance and increasing liner thickness result in an increase in the borehole diameter. The opposite trend is observed in the 100° liner angle tests (4, 5) where decreasing liner thickness and increasing stand-off distance result in an increase in the borehole diameter. The differences in results are small however for cases 4 and 5. It is thought that the damage to the target is dependent on the shape and velocity of the jet at impact, which are dependent on the liner angle, liner thickness and stand-off distance. It is expected that there is an optimum jet diameter and jet velocity, which results in the maximum achievable borehole diameter. The results would indicate that for higher liner angles a longer stand-off distance is required to develop the optimum jet diameter/velocity. The results also show that the stand-off distance influences the general damage area (Fs) of the target. To analyse the trends in detail we conducted numerical simulation for a number of cases in following sections. PROBLEM DEFINITION AND MODEL IMPLEMENTATION For the numerical modelling we considered a set-up, which is similar to that in Fig. 1. Detonation starts at point I (see in Fig. 1). We ignore the steel shell S around the target imposing the boundary condition of ambient material of the same type outside the target area (a soft boundary condition in the Eulerian calculation). In a previous publication [2] we split the problem into three steps, considering separately the SC jet formation, the loading of a concrete target by the SC jet projectile, and damage growth due to the impact load. In that study we introduced inaccuracies by 1) approximating the SC projectile with cylindrical blocks; 2) interpolating the pre-impact velocity of the blocks from the first stage of the calculation (the SC liner collapse); 3) extrapolating the load obtained from the impact calculation throughout the target axis; and finally 4) the borehole diameter was approximated by a maximum damage zone, when calculating the damage due to the extrapolated load. Therefore, there were a significant number of intermediate approximations introducing inaccuracies into the final result. In the present paper we

consider the process within a single Eulerian calculation. The problem employs 4 materials within a multi-material set-up: 1) void material (the Null-material from the DYNA material database). This is used to represent the air adjacent to the liner and casing at the freesurface sides and surrounds the target at the front and rear sides; 2) HE for the explosive charge simulated by the High-Explosive-Burn material from the DYNA material database with the JWL equation of state. The parameters of HE correspond to the Composition B;. 3) a material for simulation of the aluminium liner and the casing. A user-defined material model is used for the simulation represented by a viscoelastic rate-sensitive model [4]. Its implementation in DYNA was described previously in [5]; and 4) a material representing concrete. A user-defined material model of this target material [3] was implemented in LSDYNA3D as described in [6]. The space resolution chosen for the simulation was of the order of millimetres, similar to the liner thickness dimensions. This was necessary to obtain a reasonable resolution of the liner as well as the space of the stand-off distance and target thickness and to minimise the computational resources required. This however results in significant numerical erosion of the millimetre-range projectile when travelling through the stand-off region and target. Because of this we had to limit ourself to a more modest set-up than the experimental test: we considered usually thicker liners (5mm) and thinner targets (150 mm) to be able to complete the calculation. The problems considered are shown in Table 2. This table also lists the closest corresponding experimental set-ups from those listed in Table 1 (Ntest) and corresponding borehole/damage characteristics Fd and Fs in Table 1. N 1 2 3 4 5 6 7 8 9

TABLE 2. SUMMARY OF NUMERCAL SET-UPS Ntest h(mm) θ (˚) τ(mm) D(mm) Fd(mm) Fs(mm) 36 90 3 N/A FXR 46 90 5 80 51 420 3 46 90 5 127 45 270 2 36 90 5 127 45 270 2 36 90 3 127 44 280 1 36 100 5 127 47 270 4 36 100 5 153 49 270 5 36 110 3 93 22 455 6 36 110 5 127 24/26 350/450 7 & 8

Due to the restrictions on computational resources, the present calculations cannot provide predictive capacity but they can be used for qualitative comparative analysis to study influence of the varying parameters on the target damage.

MODELLING AND COMPARISON WITH EXPERIMENT Development of the SC projectile for the first simulation (N=1 in Table 2) is shown in Fig. 3(c). This figure shows the modelling results at the different stages (36, 40, 44, 48 and 52µsec) of SC jet development. Fig. 3(a) shows the Flash-X-Ray (FXR) photograph taken at time 39.3µsec after the detonation initiation. The modelling result at 40µsec is overlayed on the FXR photograph in Fig. 3(b). As can be seen good agreement between the modelled and experimental collapsed liners was obtained. As mentioned above a very coarse computational grid was used for the simulation. This resulted in a numerical erosion of the jet tip and slug.

FIGURE 3. Flash X-Ray photograph (a) and comparison with the simulation 1 (b) at t=40µsec. (c) – Simulation 1: jet development at five moments of time from 36 up to 52µsec (5 frames).

The following discussions refer to the simulation number N in Table 2 to specify the numerical set-up. In simulations 2-9, illustrated below in Figs. 4-11, we usually draw the first frame at t=40µsec in order to give an image of the liner shape for each simulation before impact. Then we use the moment of the SC jet impact on target as a starting point for the follow-up frames. It is given in every figure as ti. The majority of the figures are calibrated to this instant and the interframe time is set to be 40µsec. This allows the frames to be compared with each other for most simulations. The numerical set-up number is referred to in Table 2 and shown on every figure followed by the letter ‘N’. Every frame in the figures also indicates the physical time t from the time of detonation initiation. The initial SC location is shown in each figure along with the Eulerian frame surrounding the SC and target, that contains the air (void) material. At initial SC jet penetration, the figures

show one or two frames with damage concentration contours on the target (e.g., frames 1 and 2 in Fig. 4 show the damage contours). These are drawn at three levels of the damage concentration (see [3]).

FIGURE 4. Simulation 2: h=46mm, θ=90˚, τ =5mm, and D=80mm.

FIGURE 5. Simulation 4: h=36mm, θ=90˚, τ =5mm, and D=127mm.

Figs. 4 and 5 show results of simulations 2 and 4 for a 90° SC with a 5mm liner thickness at two different stand-off distances. As one would expect, the shorter stand-off distance (Fig. 4) results in a larger diameter jet impacting on the target. This results in a larger borehole diameter and qualitatively confirms the results obtained in the experiments. The front damage is larger as well for the shorter stand-off distance (frames 2 in Figs 4 and 5). It is seen that with larger stand-off distance the jet particulation increases which reduces the penetration efficiency. This tendency also develops with decreasing liner thickness, which is noticeable in Fig. 6. Thus, the increase of stand-off distance and decrease of the liner thickness are somewhat in correlation at this liner angle. This is also confirmed by the

comparison of tests 1 and 2. Both the general damage and the tendency to increase borehole diameter are in agreement with tests 2 and 3 for a liner angle of 90 degrees.

FIGURE 6. Simulation 5: h=36mm, θ=90˚, τ =3mm, and D=127mm.

FIGURE 7. Comparisons for the study of influence of (a) the charge length (simulations 3 with h=46mm and 4 with h=36mm) and (b) the stand-off distance (simulations 6 with D=153mm and 7 with D=127mm).

We did not notice a significant charge length effect as can be seen from comparison of the results from simulations 3 and 4 (the pair (a) in Fig. 7). In the range of lengths studied, the longer charge provided slightly higher jet velocity that resulted in a small increase of damage and borehole diameter which is in agreement with study [1]. Increasing the liner angle to 100° still preserves the cumulation regime. This can be seen from the observed liner shape in available experiments and from simulations 6 and 7 in Figs. 8 and 9. Figs. 8 and 9 show simulations for the 100° liners for 2 different stand-off distances,

127mm and 153mm. The results indicate that there is no significant difference between the borehole diameters in these cases. It appears that for a given liner angle and thickness, the borehole diameter is less sensitive to stand-off distances above a certain height. This is confirmed by comparing the simulations near the final stage of penetration, as shown in Fig. 7(b). Experiments confirm this statement as well, however, it should be kept in mind that the experiments (tests 4 and 5) are for different liner thicknesses. Because the thinner liner induced greater damage effects in the experiments, it appears that the jet particulation fragmentation at this liner angle has less effect on the reduction of the damage than does the jet erosion. At this angle, therefore, the easier jet formation for a thinner liner is more beneficial than the overall jet mass.

FIGURE 8. Simulation 6: h=36mm, θ=100˚, τ =5mm, and D=127mm.

FIGURE 9. Simulation 7: h=36mm, θ=100˚, τ =5mm, and D=153mm.

The last two simulations were conducted at a liner angle of 110°. Simulations with different liner thicknesses and stand-off distances (Figs. 10 and 11) demonstrate that jet erosion is dominant at this angle because this regime of the liner collapse essentially diverges from the cumulation regime. This results in a larger impact area (insufficient compactness) of the SC projectile and low velocity characteristic.

FIGURE 10. Simulation 8: h=36mm, θ=110˚, τ =3mm, and D=93mm.

FIGURE 11. Simulation 9: h=36mm, θ=110˚, τ =5mm, and D=127mm.

The increase in the stand-off distance reduces the velocity, whereas reduction of the distance reduces the liner compactness. It is seen from the calculations that the general damage is extensive but the localised damage (borehole diameter) is stepwise and poor. This is confirmed very well with tests 6-8.

DISCUSSIONS AND CONCLUSIONS An experimental and theoretical study was conducted to investigate the parameters influencing the borehole diameter due to a SC jet penetration of a concrete target. Turning to the calculation results in Fig. 3, it can be seen that the SC jet with a 90° liner angle is within an optimal range from the jet shape/velocity point of view at 3040µsec after the initiation. Shortly thereafter, the jet significantly elongates and starts fragmenting. Therefore, the stand-off distance in 70-100mm is reasonably optimal at the given thickness range; collapse of essentially thicker liners might move away from the regime of cumulation. The liner angle of 100° still preserves the liner in the cumulation regime at the thickness range, however, the jet formation takes longer and, therefore, larger stand-off distances are required for the jet to take nearly optimal shape characteristics. However, the velocity might be not high enough when increasing the stand-off distance. Projectile erosion (lack of mass for penetration) on the other hand is a prevailing factor if the liner thickness is reduced in order to increase the jet velocity. The 110° liner regime show the most deviation from cumulation among those considered; the velocity characteristic does not reach the required value for optimal penetration. Summarising, we can conclude that the mathematical model predicted the trends well and can be used to optimise the SC design and to select the stand-off distance. It seems the present SC configuration for the 90° liner is nearly optimal. To essentially increase the damage characteristics of the targets will require novel methods of initiation, advanced liner materials, or enhanced high explosives. ACKNOWLEDGEMENTS The studies conducted in the paper have been supported by a DSTO task managed by Phil Winter. The authors also acknowledge assistance in conducting the flash X-Ray photography by Dave Fraser, preparation of shaped charges by Max Joyner, and conducting the firings by Dave Harris. REFERENCES 1.

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