Superior ballistic performance of high-nitrogen steels against deformable and non-deformable projectiles

Author’s Accepted Manuscript Superior ballistic performance of high-nitrogen steels against deformable and non-deformable projectiles B.Bhav Singh, G. Sukumar, P.Prakasa Rao, K.Siva Kumar, V. Madhu, R. Arockia Kumar www.elsevier.com/locate/msea

PII: DOI: Reference:

S0921-5093(19)30215-1 https://doi.org/10.1016/j.msea.2019.02.044 MSA37563

To appear in: Materials Science & Engineering A Received date: 4 October 2018 Revised date: 13 February 2019 Accepted date: 14 February 2019 Cite this article as: B.Bhav Singh, G. Sukumar, P.Prakasa Rao, K.Siva Kumar, V. Madhu and R. Arockia Kumar, Superior ballistic performance of highnitrogen steels against deformable and non-deformable projectiles, Materials Science & Engineering A, https://doi.org/10.1016/j.msea.2019.02.044 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.

Superior ballistic performance of high-nitrogen steels against deformable and non-deformable projectiles B.Bhav singha, G. Sukumara*, P.Prakasa Raoa, K.Siva Kumara, V.Madhua, R. Arockia Kumarb

a

Defence Metallurgical Research Laboratory, Kanchanbagh, Hyderabad, India-500058

b

National Institute of Technology, Warangal, India

*

Corresponding author. Armour Design and Development Division. Defence Metallurgical Research Laboratory,Kanchanbagh, Hyderabad- 500058, India Tel.: (O): +91 40 2458 8019; fax: +91 40 2434 2252/ 0683. [email protected] (G.Sukumar)

Abstract: Hot-rolled high-nitrogen steel (HNS) plates were evaluated against AK-47 mild steel core (7.62mm x 39mm, deformable) and 14.5mm armour piercing (AP) (14.5 mm×114mm, hard steel core)

projectiles. The ballistic results of HNS plates were compared with

conventionally used Rolled Homogeneous Armour (RHA) steel plates. The ballistic tests on HNS and RHA steel plates with different thicknesses were conducted against AK-47 mild steel core and 14.5mm AP projectiles at velocities of 740±10 m/s and 1000±10 m/s respectively, in order to find out the minimum thickness required to stop the projectile without perforation. In comparison to RHA steel, HNS exhibited higher ballistic performance (in terms of minimum areal density required) against both AK-47 mild steel core and 14.5mm armour piercing projectiles. The ballistic results against 14.5mm AP was studied using an analytical model reported in literature. Detailed post ballistic (microstructural and hardness measurements) investigations were carried out to understand the ballistic performance of the two steels. Further, an attempt has been made to correlate the initial microstructure and mechanical properties with the failure mechanisms activated during ballistic impact and the resulting ballistic performance.

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Keywords: Rolled Homogeneous Armour (RHA) steel and high-nitrogen steel, ballistic performance, high strain rate, adiabatic shear band

1. Introduction Material selection is an important aspect in design of light weight armoured vehicles used in various applications. Generally, weight reduction is achieved by either changing the design of vehicle and armour system or by using newer materials with improved ballistic performance or by utilizing both routes. While there has been tremendous increase in usage of lightweight composites in various armour applications, still high strength low alloy steel with tempered martensitic microstructure is the most widely used material in various structural armour applications. This is mainly due to its good strength-ductility combination, excellent formability, welding and corrosion characteristics and its low cost of manufacturing. In addition to high strength low alloy steels, many studies have also been carried out on other class of steels such as bainitic steels, secondary hardening steels, auto tempered steels, and nickel free high-nitrogen steels (HNS) etc for potential ballistic applications [1-6]. The aim of the present study is to investigate Nickel free high-nitrogen steels (HNS) for armour applications. The present work on HNS is motivated by earlier studies [6-15]. Nickel free high-nitrogen steels have been reported to possess high strength, toughness, good corrosion resistance, non-magnetic susceptibility and high strain hardening and strain rate hardening ability [6-9].E.Lach et.al. [6] compared the ballistic performance of high-nitrogen steel and martensitic armour steel against the high velocity (2500m/s) impact of tungsten sintered alloy with an L/D ratio of about 14. It was shown that both pre-deformed HNS and conventionally used low alloy steel performed similarly against long-rod projectiles [6]. In 2

another work, ballistic studies were carried out on cold worked HNS by impacting pure tungsten and tungsten heavy alloy (L/D=10) at velocities ranging from 2000 to 4500m/s [10]. It was found that pure tungsten has higher penetration capability in comparison to tungsten heavy alloy. Apart from studies against sub-scale long-rod penetrators[6,10], it was also shown that HNS has equal ballistic limit in comparison to Rolled Homogeneous Armour (RHA) steel against 5.56mm armour-piercing (AP) projectiles [11,12]. Chen Rong et al [13] evaluated the ballistic performance of high-nitrogen steels against various small arms projectiles and found that HNS had much higher mass efficiency in comparison to similar strength martensitic steel. In all these ballistic studies, the better ballistic performance of HNS was attributed to its higher dynamic strength and good strain hardening and strain rate hardening capacity [6, 10-13]. Bhav singh et.al. [14, 15] evaluated the effect of cold rolling and hot rolling on ballistic performance of high-nitrogen steel. It was observed that the ballistic performance of HNS decreased against deformable projectiles and increased against 7.62 AP projectiles with increase in cold rolling reduction [14].In contrast, the ballistic performance against both deformable projectile and 7.62 AP projectiles improved with increase in hot rolling reduction [15]. Though HNS was found to show good ballistic resistance, the ballistic studies on HNS against different projectiles are few in literature in comparison to RHA steels. In particular, studies against deformable projectiles such as AK47 (mild steel core) are not readily available in literature and this result will be useful in light structural armour applications such as mine proof vehicles. While some studies have reported ballistic results against 5.56 AP [11-12] and 7.62 AP [14.15], the ballistic results against 14.5mm AP is very limited. The ballistic result against 14.5mm AP is very important for heavy structural armour applications such as main battle tanks and infantry combat vehicles. Hence, the present work is undertaken to comprehensively study the ballistic performance of nickel free high-nitrogen steel against these two small arms (AK-47 mild steel core and

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14.5mm AP) projectiles. Further, the ballistic performance of HNS is compared and contrasted with conventionally used RHA steels from the perspective of how the differences in plastic flow behaviour (of these two classes of steels) resulting from different microstructures affect the energy absorption and failure mechanisms during ballistic impact.

2. Materials and Experiments Hot-rolled high-nitrogen steel plates with thicknesses of 50mm and 8mm were used in this study and these were produced at M/S. Jindal Stainless Steel Limited, Hisar, India. These plates were produced by a route consisting electric arc melting, argon oxygen decarburization, continuous casting followed by hot rolling. The initial 190mm thick continuous cast slab (slab soaking temperature 1100˚C-1150˚C) was hot rolled into 50mm and 8mm thick plates. RHA steel plates with thicknesses of 50mm and 11mm were used in this study and these were procured from Rourkela Steel Plant, Steel Authority of India Limited, India. The manufacturing process of RHA steel includes steel making through electric arc furnace, ladle refining and vacuum argon degassing, continuous casting followed by hot rolling and heat treatment. The heat treatment process consists of austenitization at 910°C followed by oil quenching and tempering at 650°C. The chemical composition of the HNS and RHA steel plates used in this study are given in Table 1. Tensile tests of HNS plates were carried out on cylindrical samples taken from longitudinal direction of 50mm thick plates and flat samples taken from longitudinal direction of 8mm thick plates. Tensile tests of RHA steel plates were carried out on cylindrical samples taken from longitudinal direction of 50mm and 11mm thick plates. All tensile tests were carried out using screw driven Instron machine (Make: Instron5500R) at an initial strain rate of 8×10-3/s as per ASTM E-8M standard. The charpy V-notch impact tests were carried out on HNS and RHA steel samples taken from longitudinal direction of 50 mm thick plates and these samples had dimensions of 10 mm×10 mm×55 mm (with 2 mm deep notch). These 4

tests were carried out as per ASTM (E23-02a) standard using Tinius-Olsen machine at room temperature and -40˚C, Quasi-static strain rate compression tests were carried out on cylindrical samples (6mm diameter and 9mm length) taken from normal direction of the HNS and RHA steel plates. These tests were carried out at room temperature and high temperatures (200˚C, 400˚C & 600˚C) using a hydraulic machine (Make:Walter+Bai) at a initial strain rates of

2×10-3/s.

Room temperature high strain rate compression tests were carried out on samples taken from normal direction of the plates (6mm diameter and 3mm thickness) using a Split-Hopkinson Pressure Bar (Make:Thiot-France).The working principle of the Split-Hopkinson Pressure Bar can be found elsewhere [16]. The average strain rate during SHPB tests was around 2000-3300 /s. Microstructures of HNS and RHA steel plates in long transverse direction were studied under optical microscope and some RHA steel samples were studied under transmission electron microscopy. For microstructural observations under optical microscope, Aqua-regia (75%HCl, 25%HNO3) and Nital solution were used to etch HNS and RHA steel plates respectively. XRD experiments were carried on initial samples of HNS and RHA steels using Co-Kα radiation. The ballistic tests against small arms projectiles were carried out in a small arms range at DMRL, Hyderabad, India. Details of the different small arms projectiles used in present study are given in Table.2. Both small arms projectiles were fired from a distance of 10 meters with full velocity and the angle of attack was zero degree. The actual velocities will be discussed later. The ballistic test arrangement used for small arms testing is given elsewhere [17]. For ballistic testing against AK-47 projectiles, 8mm thickness HNS plates were machined to 7mm, 6mm, 5mm and 4mm thickness, while the RHA steel plates with thickness of 11mm were machined to 7.5mm, 7mm and 6mm. For ballistic testing against 14.5mm AP projectiles, HNS and RHA steel plates with different thicknesses such as 50mm, 5

48mm, 45mm, 43mm and 40mm were used. The lower thickness plates (<50mm) were machined from 50 mm thick plates. The plates were machined using coated carbide tool with a speed of 40m /min and depth of cut of 0.2 mm under wet (under coolant) condition. The ballistic performance against both the small arms projectiles was evaluated by observing whether a plate at a given thickness is able to stop the projectile or not. In addition, minimum areal density required to stop the projectile (critical thickness required to stop the projectile× density) was also calculated to compare the ballistic performance. After ballistic trials, the damage patterns on the front and rear face of target plates were photographed and studied. Microstructures of crater cross section were studied using optical microscopy and scanning electron microscopy. The micro hardness of initial steel plates and the micro-hardness variations across the crater cross sections were measured by Vickers hardness tester under a load of 300g.

3. Results and Discussion 3.1 Microstructure and Mechanical properties Fig.1a and b show the optical micrographs of the HNS plates at different thicknesses. While 50 mm thick hot-rolled HNS plate (Fig.1a) had equiaxed austenitic grains, the lower thickness 8mm plate (Fig.1b) showed elongated austenitic grains. The presence of single phase fcc austenitic structure in both 50mm and 8mm thick plates was confirmed by X-ray diffraction patterns (Fig.1c). The scanning electron micrographs of RHA steel plates (50mm and 11mm thickness) show needle like precipitates in lath type tempered martensite matrix. (Fig.2a and b).TEM image of 50mm thick RHA steel plate also clearly reveals the presence of needle like precipitates in lath type matrix (Fig.2c). Fig.2d shows the XRD patterns of RHA steels which correspond to tempered martensite (or ferrite) phase. The diffraction peaks from second phase precipitates were not observed in the XRD patterns, which may be due to

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their lower volume in the sample. The composition of RHA steel used in the present study is close to the AISI 4340-low alloy steels. The amount of different alloying elements in RHA steel and AISI 4340-steel are in the same range, except chromium which is higher in the case of RHA steel [18,19]. Hence, based on the composition and heat treatment used in present study and based on earlier studies on AISI-4340 steels, the RHA steel is likely to contain carbides (white particles in Fig 2a and b and dark phase in Fig.2c) in tempered martensite matrix [18,19]. However, further studies are needed to identify the exact type of the carbides. Table.3 shows the mechanical properties of HNS and RHA steel plates used in the present study. It can be seen that 50mm thickness HNS plates have lower yield strength and ultimate tensile strength and higher ductility in comparison to 8mm thickness HNS plates. Higher yield and ultimate strength of 8mm HNS plates could be attributed to their fine partially recrystallized microstructure. When mechanical properties of 50mm thick HNS plate is compared with 50mm thick RHA steel plate, the HNS plate shows lower yield and ultimate tensile strength but with a much higher % elongation and charpy-V notch impact energy (CVN) values (at both room temperature and -40°C). The superior Charpy impact values of HNS at both room temperature and -40°C can be attributed to the single phase austenitic microstructure with FCC crystal structure. The 8mm thick HNS plates had slightly lower yield strength values in comparison to 11mm thick RHA steel plates, but its ultimate tensile strength is slightly higher in comparison to 11mm thick RHA steel indicating higher strain hardening ability. Fig.3 shows the strain hardening behaviour of HNS and RHA steel plates analyzed using log-log plot of plastic strain vs.true stress. The understanding of strain hardening behaviour will be useful in carrying out the room temperature forming and shaping operations. Both HNS and RHA steel plates exhibited dual slope behaviour and the slope change is significant in case of HNS plates (Fig.3). The strain hardening coefficient ‘n’ was

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calculated by fitting the Holloman equation [20] to the later stages of plastic strain vs. true stress curve (Fig.3). It can be observed that 50mm thickness HNS plate show higher strain hardening coefficient in comparison to 8mm thick HNS plates. The strain hardening coefficient of RHA steel and HNS plates are similar up to the plastic strain value of 0.1. However, at larger strains, the strain-hardening coefficient of HNS is higher than that of RHA steel plates. Strain hardening behaviour of high nitrogen steels during static tensile testing has been widely reported literature [21,22,23]. The HNS has FCC crystal structure which contains many close packed slip planes and directions. Hence dislocation interaction and intersection is expected to be higher. Due to the presence of nitrogen, the stacking fault energy of HNS is drastically reduced and there will also be the formation of short range ordering [21,24]. These factors promote planar dislocation movement and consequently results in pileup of dislocation at grain boundaries. The pile up of dislocations produces large back stresses, which results in increased work hardening. It has been shown in earlier studies that as the amount of deformation increased the deformation mechanism changes from planar glide at lower strains to twinning at higher strains [21,22,23,24]. This change in deformation mechanism is the basis for the dual slope behaviour observed during tensile testing of HNS steel (Fig.3). The deformation twins act as effective barriers for dislocations movement and increases the strain hardening rate of HNS especially at larger strains [21,22,23,24]. The lower rate of work hardening in RHA steels can be understood from interaction of dislocations with various barriers in the microstructure. Due to the high temperature tempering (650°C), the RHA steel contains lower numbers of dislocations and coarse carbides (Fig.2c). In this microstructural condition, the dislocation–dislocation interaction will be less and dislocations can easily by-pass the coarse carbides, which results in lower work hardening rate[25,26].

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During ballistic impact of small arms projectiles, material near the impact zone undergoes large strain deformation (up to strain of ~2) at wide range of strain rates (1/sec to 105/sec) [27]. It was shown in a numerical study that ballistic impact of 7.62 AP (with a velocity of 800 m/s) on 6mm thick RHA steel plate generates an average strain rate of 104/sec and average plastic strain of 0.27 [27]. Hence, it is important to study the dynamic flow behaviour of target materials in order to understand the ballistic performance. Compression stress-strain curves of HNS and RHA steel plates tested at different strain rates are shown in Fig.4 a-d. Table.4 compares the quasi-static and dynamic flow stress values (measured at strain rate of 3300/s or 3000/s and plastic strain value of 0.1) of HNS and RHA plates. It can be observed that both HNS and RHA steels show substantially higher dynamic flow stress values in comparison to their respective quasi-static flow stress values. At higher strain rates, there is less time available for thermal energy to assist the dislocations in overcoming the short range barriers. Hence, higher stress is required to move the dislocation. The quasi-static flow stress values of 50mm and 8mm thick HNS plates are lower in comparison to the 50mm thick and 11mm thick RHA plates respectively. But, the dynamic flow stress of 50mm and 8mm thick HNS plates are higher in comparison to 50mm and 11mm thick RHA steel plates respectively. The dynamic flow stress of a material at a given strain is decided by the combined effects of strain rate hardening, strain hardening and thermal softening effects. The higher dynamic flow stress of HNS is mainly due to its higher strain rate sensitivity and higher dynamic strain hardening rate. These aspects are discussed in subsequent sections. The strain rate hardening capacity of the two steels were measured by calculating strain rate sensitivity parameter “m”. The strain rate sensitivity index ‘m’ was calculated using the formula, ̇

(1) and these values are listed in Table.4. The flow

stress values measured at a plastic strain of 10% (during quasi-static and high strain

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compression tests) were used for the calculation of strain rate sensitivity parameter. It is clear that the strain rate sensitivity of HNS is higher than that of RHA steels. The effect of strain rate on flow stress is mainly attributed to the increase in dislocation density, role of thermal activation in dislocation movement and dislocation drag effects observed at higher strain rates [28,29]. It has been pointed out that the thermal part of the flow stress (short range barriers) is mainly responsible for strain rate and temperature effect on flow stress [28,29,30]. In case of FCC metals, the main short range barrier for dislocation motion is forest dislocations [28,29,30]. The forest dislocation cutting mechanism is weakly sensitive to temperature and strain rate. Hence, the yield stress of FCC metals is usually not affected by strain rate and they exhibit lower strain rate sensitivity (m<0.01) [30,31]. The strain rate sensitivity of FCC metals mainly comes from dislocation substructure evolution (increase in dislocation density at higher strain rates) [32]. However, the higher strain rate sensitivity of HNS observed in the present study can be attributed mainly to the interaction of nitrogen with moving dislocations (which increases thermal part of the flow stress significantly) [6,33] and dislocation substructure evolution [32]. The RHA steel also shows substantial strain rate sensitivity (although it is lower than that of HNS), since its thermal part of the flow stress is mainly due to the lattice friction effects (peierls barrier mechanism) coming from the tempered martensite matrix. The lattice friction effects are quite sensitive to strain rate [28,29,30]. Fig.5a and b show work hardening rates of HNS and RHA steels (measured by best fit to the data between the plastic strain values of 0.05 to 0.25) in quasi-static and dynamic compression tests. It should be noted that the dynamic work hardening rate of 50mm HNS was calculated from the plastic strain value of 0.15 to 0.25 only. Variation of work hardening rate with strain is a result of competition between dislocation accumulation and dislocation annihilation processes [30,31]. It can be seen that work hardening rate of HNS is higher in comparison to RHA steels at both quasi-static and dynamic compression tests (Fig.5 and 10

Table.4). Higher work hardening rates of HNS can be mainly attributed to formation of mechanical twins during plastic deformation which divides the grains and results in increased resistance to dislocation movement [34]. The work hardening rates of both HNS and RHA steel during high strain rate test are lower in comparison to their respective quasi-static counterparts. At higher strain rates, accumulation of defects such as dislocations could be faster due to the higher stresses [23,35]. This could result in early onset of dislocation recovery processes (such as cross slip) which causes lower rates of work hardening [23,35]. In case of HNS, twinning shear is also reported to result in reduction of work hardening rate at large strains [34]. It should be noted that high strain rate deformation results in adiabatic temperature rise of the material, since there is less time for the diffusion of heat. This heat generation will also result in lowering of work hardening rate by assisting dislocation recovery processes [35]. It is clear from above discussions that the higher dynamic flow stress of HNS is mainly due to its higher strain rate sensitivity and higher strain hardening rate at dynamic strain rates. The high strain rate sensitivity and high strain hardening rate are also useful in resisting the thermal softening effects resulting from adiabatic temperature rise. It can be seen that RHA steel plates with lower strain rate sensitivity and work hardening rate show highly negative strain hardening rates in comparison to HNS plates with superior work hardening capability (Table.4)

3.2 Ballistic performance against 14.5mm armour piercing projectiles Fig.6a and b show the front face, crater cross-section and rear faces of 50mm thickness HNS and RHA plates tested against 14.5mm AP projectiles. It can be observed that both HNS and RHA plates could completely stop the projectile. The average depth of penetration in 50mm thick HNS and RHA steel plates were 38±1 mm and 46±1 mm 11

respectively. It was observed that a minimum of 43mm HNS and 48mm RHA steel plate is required to stop the 14.5mmAP projectiles.Table.5 summarizes the ballistic results against 14.5mm AP projectiles. It is clear that the HNS shows better ballistic performance in terms of critical thickness (~10% lower areal density) required to stop the projectile. In order to understand the ballistic performance against small arms projectiles, it is important to understand the different energy absorption and failure mechanisms involved during the projectile-target interaction. Various energy absorption mechanisms and failure modes can operate simultaneously in a target based on the projectile characteristics (size, shape, velocity and density etc) , target characteristics (strength, ductily, toughness, thickness etc) and the test conditions [36,37]. Against armour piercing projectiles such as 14.5mm AP, soft ductile targets (with plate thickness to projectile diameter ratio> 1) mainly fail by ductilehole growth process and petalling. During impact on target plate, the hard steel core of the projectile will not deform and maintains its shape (except against high strength targets). The ductile-hole growth process involves pushing aside of the target material, which mainly occurs by plastic deformation [27,36, 37]. Hence material with higher dynamic flow stress will have higher resistance to projectile penetration. Better ballistic performance of HNS (10 % lower areal density required to stop the projectile) in comparison to the RHA steel can be attributed to its higher dynamic flow stress [27]. In order quantify the effect of dynamic flow stress on ballistic performance, an attempt has been made to calculate the critical thickness at which perforation occurs using an analytical model proposed in literature [27,38]. Further, these calculated values are compared with experimentally observed critical thickness at which target is perforated. Various analytical models [38-40] have been proposed in literature for the calculation of resisting stress exerted by target plate on rigid-sharp nosed projectiles. This resisting stress values can be further used for calculating ballistic limit velocity. While some of these models 12

used cavity expansion analysis in calculating resisting forces [39], Rosenberg and Dekel [38] derived the resisting stress from numerical simulations. According to Rosenberg et al [27] , the ballistic limit velocity is calculated by

(2) Where Vbl is the ballistic limit velocity, H is the target thickness, σr is the resisting stress , ρp is the density of projectile and Leff is the effective length. The calculated from the relation,

can be

.In this relation m is the mass of the projectile

core and r is the radius of the projectile. The mass and radius of the 14.5mm AP core is given Table.2. The resisting stress for the present case (H (target thickness)/ D (projectile diameter) >1) is given as [38] )

(3)

From the equation (2), critical thickness (H) at which perforation occurs can be calculated when Vbl is taken as initial impact velocity (V0). Rosenberg et al [27] recommended that flow stress value Y (which represents the effective strength) should be calculated from compression kolsky bar experiment carried out at a strain rate ranging from 5000/s to 10000/s at a plastic strain value of 0.27. This value represents the average strain rate and strain around the crater [27]. In the present case, high strain rate compression experiments were carried out only to maximum strain rate of 3000-3300/s. Hence flow stress value Y was taken from experiments carried out a stain rate of 3000 or 3300/sec at a plastic strain value of 0.27. This way of flow stress value (Y ) selection is reasonable, since RHA steel and 304L austenitic steel have been shown to have less strain rate sensitivity beyond the strain rates of 3000/sec [27]. In this work, the flow stress value (Y) for 50mm thick HNS and RHA steel is taken as 1.68 GPa and 1.27 GPa respectively. Vbl is taken as 1000m/s, which is the initial impact velocity. From

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equation (2), the critical thickness for perforation is calculated to be 35mm and 44mm for HNS and RHA steel respectively. The value for HNS is slightly lower in comparison to experimentally observed critical thickness values at which perforation occurs (40mm). This discrepancy can be attributed to the fact that the equation-(2) assumes that only core part of projectile is involved in the perforation process. But, the bullet used in the present study had the jacket part along with rigid non-deformable core. It was shown in an earlier study on the perforation of RHA steel by 7.62 AP projectiles that Vbl observed for perforation by actual bullet with jacket is about 6% lower in comparison Vbl observed for perforation by core alone. This kind of data is not available for 14.5mm AP projectiles in literature. Assuming similar amount of Vbl reduction occurs for 14.5mm AP as in the case of 7.62 AP, the Vbl for the present case (perforation by core alone) has to be taken as 1060 m/s. Using this Vbl value, the critical thickness at which perforation occurs is recalculated to be 39mm and 48mm for HNS and RHA steel. These values are reasonably close to the experimentally observed critical thickness values at which perforation occurs (40mm for HNS target and 45mm for RHA target). It can be inferred from the above analysis that dynamic flow stress measured at large strain (effective strength as proposed by Rosenberg et al [27]) is the most important factor in deciding the ballistic performance of soft targets against rigid armour piercing projectiles and the analytical model proposed by Rosenberg et.al [27] can be used for predicting the critical thickness required perforation. The difference in the ballistic performance of HNS and RHA steel could be attributed to the difference in their dynamic flow stress values. Apart from higher dynamic flow stress, HNS has higher strain hardening rate in comparison to RHA steel. The higher strain hardening rate helps in spreading the plastic deformation, since the plastic wave velocity is directly proportional to the slope of the stressstrain curve in plastic region. [41]. Higher plastic wave velocity will result in higher volume

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of material involved projectile energy absorption, which could be useful for ballistic resistance. Fig.7 shows the normalized hardness values measured near the craters of HNS and RHA steel plates formed against 14.5mm AP projectiles. In this study, the hardness values are normalized with respect to the initial base micro-hardness of the target plate before the ballistic testing. It can be observed in both target craters that as we move close to crater edge, the hardness values increases in comparison to their base hardness. This increase in hardness from base hardness is higher in case of HNS crater in comparison to RHA steel crater. Also, the increase in hardness from base hardness was observed up to longer distance in HNS crater in comparison to RHA steel crater. For example, in HNS crater, the increase in hardness from base hardness (measured at 10 mm from impact surface) is observed up to 20mm from the crater edge (Fig.7b), whereas in RHA steel crater the hardness increase was observed up to 10mm only (Fig.7b). The observation of hardness increase up to higher distance in HNS crater is due to its higher strain hardening rate resulting from activation of deformation twins during ballistic impact (Fig.8). It indicates that larger volume of material is involved in energy dissipation in HNS in comparison to RHA steel. This could also have been an important factor for better ballistic performance of HNS. The formation of deformation twin during ballistic impact on HNS has already been pointed out in an earlier study [42].

3.3 Ballistic performance against deformable AK-47 mild steel core projectile Fig.9 (a-d) show the front face, crater cross-section and rear face of HNS plates after ballistic trials against AK-47 mild steel core projectiles. It can be seen that 5mm HNS plate (Fig.9c) is able to stop the AK-47 mild steel projectile, whereas 4mm HNS plate was perforated by the projectile (Fig.9d). Thus a minimum of 5 mm hot-rolled HNS plate is required to stop the AK-47 mild steel core projectiles. Fig.10(a-c) show the front face, crater

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cross section and rear face of RHA steel plates after ballistic trials against AK-47 mild steel core projectiles. It can be observed that 7mm RHA steel plate (Fig.10b) could stop the projectiles, whereas 6mm thick RHA steel plate was perforated by the projectile (Fig.10c). Hence a minimum of 7mm thick RHA steel plate is required to stop the AK-47 mild steel core projectiles. When rear face damage patterns are compared, it is observed that 7.5mm thick RHA steel plate shows smooth bulge, whereas 7mm thick HNS plates do not show any bulge indicating higher resistance to impact (Fig.9a and Fig.10a). Further, the amount of bulging increased with decrease in HNS plate thickness (6mm and 5mm plates) up to the observation of the perforation in 4mm thickness HNS plates. Table.6 summarizes the ballistic results against AK-47 mild steel core projectiles. It can be seen that HNS plate requires 28% lower areal density in comparison to RHA steel to stop the AK-47 mild steel core projectiles. When a soft deformable projectile such as AK-47 mild steel core projectile impacts the target, the projectile energy is mainly consumed by plastic deformation of the target and by formation of ASB and ASB induced cracks in target [17,37,43]. The energy absorbed by plastic deformation will be high if the flow stress of the material is high and more volume of material is involved in the dissipating the kinetic energy of the projectile. Apart from the target material deformation, some part of projectile energy is consumed by projectile deformation itself. It should be noted that the projectile deformation (blunting) will be more in the case of high strength targets in comparison to lower strength targets. Due to blunting, the projectile will lose its ability to penetrate the target. Hence high strength targets will have better chance of defeating the soft projectile. However, in some cases, high strength targets may be shattered or perforated easily due to activation of low energy failure modes such as brittle cracking, spalling and fragmentation [36,37,43]. Thus high strength targets may not exhibit higher ballistic performance in all the cases. In the present study, both HNS and RHA 16

steels have enough toughness and ductility, hence they do not exhibit brittle cracking behaviour. Also, when the target thickness to projectile diameter ratio (t/d) is close to 1 (which is the case in present study), the main mechanism by which the target fails against deformable projectiles is through ASB induced plugging [17,37,43]. Hence, resistance to ASB formation is an important factor along with dynamic flow stress and toughness in deciding the ballistic performance against the soft projectiles. ASBs generated during ballistic impact are the result of the thermo mechanical instability occurring at higher strain rates. During high strain rate deformation, there is little time available for heat dissipation, due to which there is a temperature rise in certain regions, resulting in localized severe deformation [43]. It is clear from various studies on adiabatic shear band formation [43-46] that the material with high density, high specific heat capacity, higher strain hardening coefficient and strain rate sensitivity, and low thermal softening parameter will have lower propensity to ASB induced plugging. The HNS has higher strain hardening rate, higher strain rate sensitivity than that of RHA steel (Table.4). Hence HNS is expected to show better resistance to the ASB formation in comparison to RHA steel plates. In this work, an attempt has been made to quantitatively calculate the adiabatic shear sensitivity of HNS and RHA steel using an analytical equation derived by Wright et al [45,46] According to Wright [45,46], the susceptibility to ASB formation is given by

{

( ) √

}

(4)

In this equation (4), χSB is the parameter representing the susceptibility to ASB, n is the strain hardening exponent, m is the strain rate sensitivity and α is the non-dimensional

17

thermal softening parameter defined by α

where σ is the flow stress, T is the

temperature, ρ is the density and c is the specific heat of the material. Higher the χSBvalue, higher is the possibility of shear band formation. In this study, the density and specific heat of both steels are taken as 7800 Kg/m3 and 452 J/Kg K respectively. The ‘n’ values were calculated from quasi-static compression flow curves (Fig.11a). The ‘m’ values of HNS and RHA steel were calculated to be 0.038 and 0.027 respectively (Table.4). The thermal softening values were calculated from the high temperature quasi-static compression flow curves. Fig.11b shows the effect of temperature on flow behaviour of HNS and RHA steel and Fig.11c shows the variations of flow stress with respect to temperature. The thermal softening parameter for HNS and RHA steel is found to be -0.74 MPa/K and -0.97 MPa/K respectively. Using these n, m and thermal softening parameter values, the adiabatic shear band sensitivity parameter of 8mm thick HNS and 11mm thick RHA is calculated as 1.05 and 10.2 respectively. It is clear that HNS is having significantly higher resistance to adiabatic shear localization. In order to further confirm the difference in adiabatic shear localization tendency between HNS and RHA steel, post ballistic microstructural analysis was carried out. Fig.12a shows typical crater cross section microstructure in 7.5mm thick RHA plate tested against AK-47 mild steel core projectile. A prominent ASB (which travel larger distance in thickness direction) is observed in the crater cross section of 7.5mm RHA steel plate (Fig.12a). Fig.12b shows crater cross section microstructure of 5mm thick HNS plate. It can be noticed that even at 5 mm thickness, no prominent ASBs are observed in HNS crater. The HNS is able to homogeneously dissipate the projectile energy. Heavy deformation twinning is observed near the impact location of the crater (Fig.12b). These twins subdivide the grain resulting increased work hardening rate and thereby formation of ASBs will be postponed in the case 18

of HNS. Fig.12c shows the crater cross section of 4mm thick HNS plate which was perforated by the projectile. Traces of adiabatic shear bands can be found in the crater walls (Fig.12c). This confirms that ASB induced plugging is the perforation mechanism in HNS plate also. HNS plates with higher dynamic flow stress and higher ASB resistance perforated only at 4mm, whereas the RHA steel plate with lower dynamic flow stress and ASB resistance was perforated at 6mm thickness itself. In summary, the better ballistic performance of HNS (28% lower areal density required to stop the projectile) in comparison to RHA steel against AK-47 mild steel projectiles can be attributed to its higher dynamic flow stress, higher volume of material involved in energy absorption and higher resistance to ASB induced plugging.

Conclusions In this work, we have studied the ballistic performance of HNS and RHA steels against deformable AK-47 mild steel core and non-deformable 14.5mm AP projectiles. ASB induced plugging and ductile-hole growth were the main target defeat mechanism against AK-47 mild steel core and 14.5mm AP projectiles respectively. HNS showed better ballistic performance against both deformable and non-deformable projectiles. The better ballistic performance of HNS in comparison to RHA steels can be attributed to its higher dynamic flow stress, higher volume involved in energy absorption, higher resistance to adiabatic shear localization and tensile failure. Higher dynamic flow stress and higher resistance to shear localization of HNS is due to a) higher strain hardening and lower thermal softening, resulting from single phase austenitic microstructure (with FCC crystal structure) and b) due to the presence of nitrogen, which increases the thermal part of the flow stress tremendously (higher strain rate hardening) and reduces the stacking energy (which results in activation of 19

mechanical twinning). The better ballistic performance of HNS against commonly used two ammunitions shows that HNS is a potential candidate material for various armour applications.

Acknowledgments The author would like to thank Director, DMRL for giving permission to publish this paper. The authors would like to thank DRDO for funding this work. The authors thank EMG, MBG small arms range and high strain rate testing team for their support in the experimental work. The authors thank IIT madras for their help in TEM studies. The raw/processed data required to reproduce these findings cannot be shared at this time as the data also forms part of an ongoing study.

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[18] W.S Lee, T.Tian Su. Mechanical properties and microstructural features of AISI 4340 high-strength alloy steel under quenched and tempered conditions. J Mater Process Tech 87(1999) 198-206. [19] N S Lim,C W Bang,Sanjeev das,H W Jin,R. Ayer, C G Park. Influence of tempering temperature on Both the Microstructural Evolution and elemental Distribution in AISI 4340 steels. Met Mater.Int 18(1)(2012) pp 87-94 [20] J.H Holloman. Tensile deformation,Trans AIME162 (1945) pp.268-29 [21] P. Mullner, C. Solenthaler, P. Uggowitzer and M. O. Speidel. On the effect of nitrogen on the dislocation structure of austenitic stainless steel. Mater Sci and Eng, A164(1993) pp.164-169 [22] B.Karthick, R.Veera babu, DVV Satyanarayana. Effect on aging and oxidation on strain hardening behaviour of nickel free high nitrogen steels.Met.Mater.Int.22(3) (2016) 413-423 [23] G Sun ,A Yu, S Sun,C Ji, Z jiang, J Lian. Plastic deformation and fracture behaviour of high-nitrogen nickel free austenitic stainless steel. Mater Sci.Tech (2017) pp1-10 [24] V.G. Gavriljuk, H. Berns, High Nitrogen Steels, Springer, Berlin, 1999. [25] G.Krauss and D.K Matlock . Effects of strain hardening and fine structure on strength and toughness of tempered martensite in carbon steels. Journal De physique IV 5 (1995) C8-51-60 [26] G.Krauss. Deformation and Fracture in martensitic carbon steels tempered at low temperatures.Mat Trans A, 32A(2001) pp.861-877 [27] Zvi Rosenberg, Roman Kositski. Alon Malka-Markovitz. The effective strength of metallic plates perforated by rigid projectiles. Int.J.Impac.Engg 121(2018)pp 35-43 [28] K.T Ramesh. Nano Materials, Materials and Mechanisms,Springer,Dordrecht.2009 [29] M.A Meyers, Dynamic behaviour of Material. J.Wiley, New York 1994. [30] G.T Gray III.High-Strain-Rate Deformation: Mechanical Behaviour and Deformation Substructures Induced. Annual review of Materials science. 42 (2012) pp.285-303. [31] Wei Q, Cheng S, Ramesh KT, et al. Effect of nanocrystalline and ultrafine grain sizes on the strain rate sensitivity and activation volume: fcc versus bcc metals. Mater Sci Eng, A 381(2004) pp.71–9. [32] P.S.Follansbee,U.F.Kocks, A constitutive description of the deformation of copper based on the use of the mechanical threshold stress as an internal state variable, ActaMetall.36(1988) pp.81–93.

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[33] E Lach, A Bohmann, M.Scharf, E Werner. Deformation behavior of nitrogen-alloyed stainless steels at high strain rates. Advanced Engg Mater 2(11) (2000) pp.750-752. [34] S. Frechard , A. Redjaımia,,E. Lach , A. Lichtenberger. Dynamical behaviour and microstructural evolution of a nitrogen-alloyed austenitic stainless steel. Mat Sci Eng 480(2008) pp.89-95 [35] V Pare ,S.Modi ,K Jonnalagadda . Thermo-mechanical behavior and bulk texture studies on AA5052-H32 under dynamic compression. Mater Sci Engg A 668(2016) pp38-49 [36] M E Backman and W Goldsmith. The mechanics of penetration of projectiles into targets.Int.J Engng Sci.16 (1978) pp.1-99 [37] RL.Woodward, The interrelation of failure modes observed in the penetration of metallic targets. International Journal of Impact Engineering 23(2) (1984) pp. 121-129 [38] Z Rosenberg , E Dekel . Revisiting the perforation of ductile plates by sharp-nosed rigid projectiles. Int J Solids Struct 47 (2010) pp 3022–33. [39] MJ Forrestal , VK Luk, NS Brar . Perforation of aluminum armor plates with conicalnose projectiles. Mech Mater 10 (1990) pp.97–105. [40] Masri R. Ballistically equivalent aluminium targets and the effect of hole slenderness ratio on ductile plate perforation. Int J Impact Eng 80 (2015) .pp45–55 [41] G. E.Dieter. Mechanical Metallurgy (SI metric edition). Mcgraw-Hill Bookcompany. New York.1976. page no-49 [42] H Berns, S Riedner, V Gavriljuk, Y Petrov, A Weihrauch. Microstructural changes in high interstitial stainless austenitic steels due to ballistic impact. Mat SciEngA528 (2011) pp.4669-4675. [43] RL.Woodward, The penetration of metal targets which fail by adiabatic shear plugging.Int. J. Mech. Sci. 20(1978) pp.599-607 [44] R.S. Culver: in Metallurgical Effects at High Strain Rates, R.W. Rohde, B. M. Butcher, J. R. Holland, and C. H. Karnes, eds., Plenum Press, New York, NY, 1973, p. 519. [45] T W. Wright. Approximate analysis for the formation of adiabatic shear bands. J. Mech. Phys. Solids, 38(4) (1990) pp.515-530. [46] T.W. Wright, The Physics and Mathematics of Adiabatic Shear Bands, Cambridge Press NY, USA, 2002. 277. .

23



20µm

20µm

γ (111) γ (200)

γ (220) γ (311) γ (222)

Fig.1 Optical micrographs of (a) 50mm thickness HNS plate (b)8 mm thickness HNS plate (c) XRD patterns of the 50mm and 8mm thickness HNS plate showing the peaks of FCC austenite

24

10µm

10µm

500nm

Fig.2 Scanning electron micrographs of (a) 50mm thickness RHA steel plate (b) 11mm thickness RHA steel plate showing precipitates (white color) in ferrite laths (c) Transmission electron micrograph of 50 mm thickness RHA steel plate showing carbides in intra and inter lath locations d) XRD patterns of the 50mm and 11 mm thickness RHA steel plate showing bcc peaks of ferrite phase

25

RHA steel_11mm (n=0.13)

Fig.3 Log –log plot showing plastic strain vs.true stress data of the HNS and RHA steel plates during tensile testing

26

b

a

d

c

Fig.4 Compression stress-strain curves of HNS and RHA steel plates tested at quasi-static and high strain rates

27

a

b

Fig.5 Work hardening rates of a) HNS and b) RHA steel plates tested at quasi-static and high strain rates

28

Rear

a

b

Front face

Crater Cross section

Rear face

Fig.6 Front face, crater cross-section and rear face of 50mm thickness a) HNS and b) RHA steel plates tested against 14.5 AP projectiles

29

10mm Impact direction

50 mm

25 mm

Hardness measurement

Crater endeeend

Fig.7 (a) Schematic showing locations at which hardness measurements are made (b) Normalized hardness variation across the crater cross section of 50 mm thick HNS and RHA steel plates tested against 14.5mm AP projectiles

30

20µm

Fig.8 Crater cross section microstructure of 50 mm thick HNS plates tested against 14.5mm AP projectiles showing heavy mechanical twinning

31

a

b

c

d

Front face

Crater Cross-section Rear face channel

Fig.9 Front face, crater cross-section and rear face of HNS plates tested against AK-47 projectiles (a) 7mm thick plate(b) 6mm thick plate (c) 5mm thick plate (d) 4 mm thick plate

32

a

b

c

Front face

Crater cross-section channel

Rear face

Fig.10 Front face, crater cross-section and rear face of RHA steel plates tested against AK47 projectiles (a) 7.5mm thick plate (b) 7mm thick plate c) 6mm thick plate

33

b

a

c

Fig.11 a) Log (plastic strain) vs.log (true stress) plot of 8mm thick HNS and 11mm thick RHA steel plates at quasi-static compression b) Effect of temperature on plastic flow behaviour of 8mm thickness HNS and 11mm thickness RHA steel plates c) Thermal softening parameter of 8mm thick HNS and 11mm thick RHA steel

34

aa SB

100µm

b

10µm

c

Traces of Adiabatic shear bands

Fig.12 Crater cross section microstructure of a) 7.5mm thick RHA steel plate b) 5mm thick HNS plate c) 4mm thick HNS plate tested against AK-47 mild steel core projectile 35

Table.1 Chemical composition of the HNS and RHA steel plates (in weight %) Material

Ni

Mo

V

Cr

Mn

N

Si

S

P

C

Fe

HNS

0.2

0.011

-

19.3

20.6

0.54

0.35

0.007

0.02

0.076

Bal

RHA steel

1.7

0.35

0.15

1.5

0.6

-

0.25

0.01max

0.015max

0.32

Bal

Table.2 Details of the Small arms ammunitions used in present study

Core Sl. No.

Ammunition

1

AK-47

2

14.5 AP

Caliber (mm)

Velocity (m/s)

Length

Dia

Weight

(mm)

(mm)

(gms)

7.62x39mm

7.62

740

20

6

5.6

Mild steel

14.5x 114 mm

14.5

1000

52.7

12.36

40.1

Hard steel

Dimensions

36

Material

Table.3 Mechanical properties of the HNS and RHA steel plates

Thickness Material

UTS YS (MPa)

(mm)

(MPa)

% Elongation

%Reductio n in Area

n

CVN (J)

CVN@

Micro Hardnes s

-40˚C

(VHN)

(J)

300g

(%RA)

HNS

50

525

905

61

75

0.4

280

273

250

RHA Steel

50

860

947

20

65

0.16

100

90

280

HNS

8

930

1072

33

70

0.3

-

-

370

RHA Steel

11

980

1052

17

67

0.13

-

-

320

Table.4 Various measured and calculated parameters from compression stress-strain curves of HNS and RHA steel plates.

10%flow stress at Quasi-static strain rate Thickness Material

(0.002/s)

10%flow stress at high strain rate (3300/s or 3000/s)

Strain rate sensitivity parameter

(mm) (m) (MPa)

(MPa)

Work hardening rate at quasi-static strain rate (0.002/s) (MPa)

Work hardening at high strain rate (3300/s or 3000/s (MPa)

HNS

50

726

1570

0.054

1622

64

50

855

1460

0.037

22

-883

8

1040

1780

0.038

1126

-43

11

1077

1580

0.027

33

-967

RHA Steel HNS RHA Steel

37

Table.5 Ballistic results of HNS and RHA steel plates against 14.5mm AP projectiles Minimum Areal density required to stop the projectile

Thickness S.No

Material

Ammunition

Velocity (m/s)

No of shots

(mm)

Rear face observation

(critical thickness×density) (kg/m2)

50

3

Smooth bulge

3

Smooth bulge

43

3

Smooth bulge

40

3

Perforation

50

3

Smooth bulge

3

Smooth bulge

3

Perforation

45 1

2

HNS

RHA steel

14.5 AP

14.5 AP

335

990-1010

48

990-1010

45

38

374

Table.6 Ballistic results of HNS and RHA steel plates against AK-47 mild steel core projectiles Minimum Areal density required to stop the projectile

Thickness S.No

Material

Ammunition

Velocity (m/s)

No of shots

(mm)

Rear face observation

(critical thickness×density) (kg/m2)

1

2

HNS

RHA steel

7

753,751,745

3

No bulge

6

735,739,738

3

Smooth bulge

5

738,738,736,740, 740,736

6

Smooth bulge

4

743,742,734

3

Perforation

7.5

732,745,735,732

4

Smooth bulge

7

746,748,744,744, 753,737

6

Smooth bulge

6

748,739,747

3

Perforation

AK-47

AK-47

39

39

54.6