FIBER REINFORCED THERMOPLASTICS FOR BALLISTIC IMPACT MICHAEL A. MAGRINI SELVUM (BRIAN) PILLAY, COMMITTEE CHAIR JONG-EUN KIM UDAY K.

Transcription

1 FIBER REINFORCED THERMOPLASTICS FOR BALLISTIC IMPACT by MICHAEL A. MAGRINI SELVUM (BRIAN) PILLAY, COMMITTEE CHAIR JONG-EUN KIM UDAY K. VAIDYA A THESIS Submitted to the graduate faculty of The University of Alabama at Birmingham, in partial fulfillment of the requirements for the degree of Master of Science BIRMINGHAM, ALABAMA 2010

2 Copyright by Michael A. Magrini 2010

3 FIBER REINFORCED THERMOPLASTICS FOR BALLISTIC IMPACT MICHAEL A. MAGRINI MATERIALS SCIENCE AND ENGINEERING ABSTRACT Fiber reinforced plastics are gaining exposure as high performance, low weight solutions to high velocity impact resistance. More specifically, fiber reinforced thermoplastics (FRTPs) are being considered over current fiber reinforced thermosets due to superior energy absorbing mechanisms, lower cost, more efficient processing, and recyclability. Three types of materials were analyzed: (a) glass fiber reinforced epoxy, (b) glass fiber reinforced polyethylene, and (c) self-reinforced polyolefins. A fiber reinforced thermoset composite was included in testing in order to better understand the performance and benefits of the thermoplastic matrix system. All samples had areal densities around 5 kg/m 2 (1 lb/ft 2 ) and were tested at m/s using a single stage gas gun and m/s using a universal receiver which fired live ammunition. Static indentation testing was also conducted in order to compare failure modes. The selfreinforced polyethylene (PE) performance was superior to the other materials as energy absorption per unit areal density was 64.5% higher than the next highest material. Back face signatures were approximately 62% lower for self-reinforced polypropylene (PP) materials, indicating that the woven architecture resulted forced more local damage, consisting of mostly fiber shear. The damage of the self-reinforced PE was more significant due to the cross-ply laminate architecture, which allowed the fibers to experience much more strain. Testing with a hard strike plate in front of the selfreinforced PE showed an increase in front face damage for decreased strike plate iii

4 thickness. Increasing projectile KE by 382% led to an increase in front face damage area by 405%. Equations modeling the back face signature cone radius and predicting the exit velocities of projectiles were validated for PP-1B. Predicted exit velocities were within 10-20% of actual velocities of PP-1B. LS-DYNA models utilizing MAT162 for PP-1B were validated with experimental results. iv

5 ACKNOWLEDGMENTS I would like to offer my deepest gratitude to my advisor and committee chair, Dr. Selvum Pillay, for his knowledge, inspiration, and guidance throughout my time conducting this research and for his encouragement through my undergraduate years which instilled in me a desire to further my education. Without his guidance, this research wouldn t have been possible. I would like to thank Dr. Uday Vaidya for offering his wealth of knowledge to assist me in completing this research. His outlook on research is inspiring and has created in me a sense of enjoyment and adventure towards working on the next task. I am greatly appreciative of Dr. Jong-Eun Kim for his patience during my Ls- Dyna learning experience. His expertise with Ls-Dyna made the computer simulations for this research possible, and I am thankful that I had such an excellent resource to consult. I would like to thank everyone in the Composites Group for always helping me load the equipment for range testing, Balaji and Ameya for performing the static indentation testing, and the general encouragement from all of the faculty, staff, and students in the Department of Materials Science and Engineering. v

6 Last, but not least, I want to thank my family for supporting my educational ventures over the years and motivating me to be the best that I can. Without their love and encouragement, I would not be where I am today. vi

7 TABLE OF CONTENTS ABSTRACT... iii ACKNOWLEDGMENTS...v TABLE OF CONTENTS... vii LIST OF TABLES... ix LIST OF FIGURES...x LIST OF ABBREVIATIONS... xiii I. INTRODUCTION...1 II. LITERATURE REVIEW Fiber Failure Primary Yarns Secondary Yarns Cone Formation and Elastic Response Delamination and Matrix Cracking Friction Fiber/Projectile Friction Fiber/Fiber Friction Shear Plugging Ballistic Limit Calculations on Backing Panels Static Indentation for Predicting High Velocity Impact Failure III. OBJECTIVES IV. MATERIALS AND METHODS Materials and Processing vii

8 4.1.1 Glass Fiber/Epoxy matrix Glass Fiber/Ductile Matrix Polyolefin Fiber/Ductile Matrix Testing Methods Gas Gun Apparatus Universal Receiver Apparatus Static Indentation Testing Procedure V. RESULTS AND DISCUSSION Gas Gun Testing Results Glass Fiber/Epoxy Matrix Glass Fiber/Ductile Matrix Polyolefin Fiber/Ductile Matrix Universal Receiver Testing Results Glass Fiber/Ductile Matrix Polyolefin Fiber/Ductile Matrix NIJ Standard Testing Universal Receiver NIJ Level III-A NIJ Level III NIJ Level IV mm AP Test Static Indentation VI. MODELING Back Face Signature Modeling Static Indentation vs. Ballistic Impact Impact Simulations VII. CONCLUSIONS VIII. REFERENCES viii

9 LIST OF TABLES Table 1. Testing results for impact velocities of m/s with 7.62 mm spherical projectile Table 2. Testing results for impact velocities of m/s with 7.62 mm ogival projectile Table 3. Experimental vs. predicted exit velocity values from static indentation data Table 4. Comparison of simulated and observed values for PP-1B ix

10 LIST OF FIGURES Figure 1. Schematic of the primary and secondary yarns undergoing ballistic impact....7 Figure 2. (a) Projectile impacting a single fiber and (b) the transmission of the stress waves....8 Figure 3. Pyramidal shape of back face signature formed during impact in plain weave fiber reinforced materials Figure 4. Schematic showing the volume of the bulge at the back of the ceramic and the equivalent volume bulge on the back face of the backing plate. Adapted from Ravid et al [24] Figure 5. Common projectile shapes encountered during ballistc impacts: (a) hemispherical, (b) flat-faced cone, (c) cylindrical, (d) spherical, and (e) ogival Figure 6. Universal Receiver testing setup Figure 7. Energy absorbed/areal density for 7.62 mm spherical projectile at m/s. * - Projectile did not penetrate Figure 8. Glass/Epoxy post-impact delamination area on the back face of the sample Figure 9. Back face signature height per unit thickness for samples impacted at m/s with a 7.62 mm spherical projectile. * - Projectile did not penetrate Figure 10. PP-2 impact damage (left) and exit damage (right) Figure 11. Side views of impact damage to PP-1A (left) and PP-1B (right) Figure 12. Fiber wrinkling on exit side of PP-1A sample Figure 13. Impact face damage of PE-1A (left) and PE-1B (right) Figure 14. Elliptical shape on exit side of cross-ply laminated sample of PE-1A x

11 Figure 15. Comparison of back face signature cone volume for 7.62 mm spherical projectile at m/s. * - Projectile did not penetrate Figure 16. Energy absorbed per unit areal density for 7.62 mm ogival projectile at m/s Figure 17. Primary fiber failure from front side (left) and side view (right) in Glass/PE Figure 18. Back face signature cone height per unit thickness for samples impacted at m/s with a 7.62 mm ogival projectile Figure 19. Comparison of back face signature cone volume for 7.62 mm ogival projectile at m/s. * - See note from Figure Figure 20. Similar damage on entrance (left) and exit (right) sides of PE-1A Figure 21. Relationship between back face cone volumes for samples tested with 7.62 mm spherical projectile at ~ 350 m/s and samples tested with 7.62 mm ogival projectile at ~ 710 m/s Figure 22. Side view of PP-1A showing back face signature and inter-laminar shear/delamination after NIJ III-A test Figure 23. Front face of PE-1B showing radial fiber wrinkling from the point of impact after NIJ III-A test Figure 24. Back face signature of PE-1B following NIJ III-A test Figure 25. Back face signature of PE-1A following NIJ III test Figure 26. Static indentation data for PP-1B Figure 27. Top view of the mesh used to simulate impact of PP-1B Figure 28. Left: Model of PP-1B damage at 350 m/s. Right: Actual PP-1B damage at 350 m/s Figure 29. Left: Model of PP-1B damage when impacted at 683 m/s. Right: Actual damage of PP-1B at 683 m/s xi

12 Figure 30. Left: Computer simulation cross-section exhibiting PP-1B response to ballistic impact at 867 m/s. Right: Cross section of PP-1B showing projectile displacement and material delamination at impact velocity of 867 m/s xii

13 LIST OF ABBREVIATIONS FRTP PP PE BFS Fiber Reinforced Thermoplastic Polypropylene Polyethylene Back Face Signature xiii

14 I. INTRODUCTION The use of fiber reinforced thermoplastic composites (FRTPs) has been expanding in recent years to include a large share in the area of military applications. This is driven by the excellent properties obtained from these lightweight materials, the wide range of uses, and ease of manufacturing and repairing. One application where fiber reinforced plastics have found a niche is in the area of impact resistance. When compared to current armor grade materials, such as metals, the benefits of high performance fiber reinforced composites become significantly more appealing. The main drawback of using metals lies within the weight and the generation of dangerous spall when penetrated. In order to achieve an acceptable level of ballistic protection in tanks and personnel carriers commonly experiencing high lethality threats, large quantities of traditional materials such as rolled homogeneous armor steel, aluminum, and titanium are needed which tend to push the total operational weights to between tons [1]. The high hardness of these materials, usually between on the Rockwell C scale, is used to initially fracture and/or blunt the projectile. FRTPs have been shown to excel in impact protection [2, 3, 4]. This provides a very compelling reason for the military to turn towards advanced composites for use in high mobility vehicles. The military is currently employing hard armor panels composed of a phenolic resin matrix reinforced with S-2 glass fibers in High Mobility Multipurpose Wheeled Vehicles [5]. The military began utilizing fiber reinforced composites for soldier helmets in the early 1980s [6]. The helmets were constructed of layers of Kevlar fiber reinforcement 1

15 with a phenolic resin matrix. This was a significant improvement from the earlier helmets fielded that were constructed of steel; these helmets were relatively heavy with limited performance compared to current fiber reinforced helmets. Another use of fiber reinforced composites for ballistic applications is for body armor. The majority of light body armor is constructed of high tensile strength fibers, such as aramid or ultra high molecular weight polyethylene fibers, and is suitable to stop lower velocity projectiles such as those from handguns. When higher threat level projectiles are encountered, these vests are supplemented with a ballistic plate insert, which is usually a ceramic material. In these applications, fiber reinforced composites are the most suitable due mainly to a high level of performance and weight reduction. In an integrated armor system, a material of high hardness is frequently used to absorb the initial impact of the projectile, which degrades and fractures the projectile and also distributes a fraction of the load over a larger area [3, 7, 8]. Strike face components are commonly manufactured from materials such as ceramics and metals. The degradation of the projectile significantly decreases the probability of penetration through the rest of the armor and is critical for the success of the armor. The backing plate serves the purpose of providing structural integrity to the armor system and also provides the system with a reinforcement to catch the projectile and/or spall and act as a further energy absorbing mechanism [3]. Hazell and Appleby-Thomas [9] performed a study to validate the addition of woven aramid fibers to the back of carbon fiber reinforced plastics to improve the integrity of aircraft structures when faced with projectile impact. The stiffness of the components was held constant while the impact resistance increased noticeably with only a small increase in weight. The need to engineer these backing plate 2

16 materials to improve performance and reduce weight is crucial as the lives of soldiers depend on these materials. One of the main reasons for using FRTPs lies in the properties, which consist of high levels of energy absorption, specific strength, and specific stiffness [10]. The drawing process of thermoplastic fibers for reinforcement provides them with strengths commonly reaching 200 MPa and up to 20% strain-to-failure. The ideal fiber is one that has a high tensile strength and a high strain-to-failure. This allows the fiber to accept a large load and also deform before it fails, absorbing the maximum amount of energy. When multiple fiber bundles are combined, they all act in accordance with each other to resist the projectile. Adding a resin matrix to these fibers allows load transfer to the fibers and gives the component structural integrity. An important factor that the military has to consider with troops and vehicles is mobility. The ability to rapidly deploy units provides a tactical advantage and can have a significant impact on the outcome of an event. The mobility of a vehicle on the battlefield is also crucial to its survival, thus fueling the desire to achieve the lightest possible armor. A benefit to using fiber reinforced composites in armor applications is the large reduction in weight. When comparing traditional homogenous rolled alloy steel armor with composite armor of the same ballistic resistance, the composite armor will be approximately 50% of the weight [11]. This allows for quicker and more efficient movement as well as savings in fuel consumption and a decreased wear rate on vehicular components. 3

17 The benefit to using a thermoplastic matrix is the fact that complex shapes can be achieved with relative ease. The thermoplastic matrix can be positioned in any orientation, exposed to heat, and will permanently take the mold shape upon solidification. This can be done under a vacuum, under high pressure, or in a mechanical press. Thermoplastics are also recyclable, which allows reintegration of material at the end of a component s lifespan or reprocessed in the case of scrap material or in the event of an improper processing run. If a component becomes damaged in service, repair is also not difficult as the damaged part can be reconstructed and reprocessed. The main focus of this research will be on the testing and evaluation of FRTPs for impact protection. Impact responses and resulting failure mechanisms in various materials will be studied to determine the damage progression and energy absorption. Understanding the failure modes and fracture mechanics of FRTPs under high strain rates will lead to a better understanding of why thermoplastics would be suitable for ballistic impact. 4

18 II. LITERATURE REVIEW The energy absorption of FRTPs under high strain rates can be characterized through many different modes, namely tension and fracture of the primary yarns, deformation of secondary yarns, cone formation in the material, delamination, shear plugging, matrix cracking, friction, stress wave transmission, and fiber pull out/matrix debonding [2, 7, 10, 12, 13, 14]. Different research groups assemble these energy absorbing mechanisms into varying categories, but the generalized groups most often seem to be elastic response, fiber engagement/failure, cone formation, delamination [7, 15]. The total energy absorbed in FRTPs is the difference between the initial and exit velocities and can be approximated by Equation 1 [7, 15]. E Abs = ½m(V i -V f ) 2 = E F + E D + E CF + E ER (Eq. 1) where E Abs is total energy absorbed, V i is initial velocity, V f is exit velocity, E F is energy absorbed by fiber failure, E D is energy absorbed by delamination, E CF is energy absorbed by cone formation, and E ER is energy absorbed by the elastic response of the material. If the final velocity of the projectile is zero and full penetration did not occur, then the material absorbed all of the kinetic energy of the projectile. Theoretically, the point at which the projectile completely penetrates with a final velocity of zero is called the ballistic limit, or V 50. The energy absorption by the individual constituents varies for different materials [16]. Numerous factors such as projectile geometry, size and composition, impact 5

19 velocity, and laminate thickness influence energy absorption methods [8]. In many studies, the fiber failure appears to play the largest energy dissipation role, while the addition of a matrix system results in more failed fibers [14]. Failure modes are ultimately dependent upon the material properties, weave architecture, plate thickness, and projectile velocity. For example, at an impact velocity below V 50 for a given material thickness, a thinner plate will undergo more bending and deflection, resulting in higher percentages of delamination. In the same conditions, a thicker plate, which will act stiffer upon loading, will experience significantly more fiber engagement and less delamination. A higher extent of deformation in an impacted material generally leads to a higher value of absorbed energy [17]. 2.1 Fiber Failure In this section, the progression of the projectile and energy absorption through the composite will be observed with respect to the yarn/projectile interaction and the yarn/yarn interaction. Many variables such as projectile shape and velocity, fabric structure, number of plies, and friction must all be accounted for during a high velocity impact [4]. In the case of projectile shape and velocity, higher velocity projectiles with aerodynamic geometries produce a higher chance of penetration. The fabric structure also affects the composite through friction and interaction of the fibers [30]. At the area of initial contact between the projectile and the front face of the composite, two classifications of the fibers are developed, namely primary and secondary yarns. The primary yarns encompass the fibers directly in contact with the projectile, while secondary yarns are described as those that intersect the primary yarns [4, 10, 18]. 6

20 Figure 1 shows a schematic of the primary and secondary yarns during an impact event [9, 18]. In order to understand the interaction that takes place, it is assumed that the impact event is limited to a single fiber. Figure 1. Schematic of the primary and secondary yarns undergoing ballistic impact. Note: From Ballistic Impact Behaviour of Woven Fabric Composites: Parametric Studies by N. K. Naik, P. Shrirao, and B. C. K. Reddy. Materials Science and Engineering: A. Vol. 412, Issues 1-2, 5. p Copyright 2005 by Elsevier. Reprinted with permission. 7

21 2.1.1 Primary Yarns As a primary fiber is impacted, a longitudinal wave and a transverse wave are generated. The main energy dissipating characteristic of a fiber is attributed to its ability to withstand high loads and transfer them to surrounding fibers before local failure occurs [12]. The longitudinal wave moves along the fiber axis; and as the tail of the wave passes, material from the fiber flows toward the impact point [10, 18]. The movement of material through the thickness is known as the transverse wave and deflects in the direction of the penetrating projectile [4]. Figure 2 shows the impact of a projectile and the transmission of the resulting stress waves. Figure 2. (a) Projectile impacting a single fiber and (b) the transmission of the stress waves. Note: From High performance fibers for lightweight armor in AMPTIAC Quarterly. Volume 9, Number 2. p. 4. Copyright 2005 by AMPTIAC. Reprinted with permission. 8

22 The longitudinal stress wave undergoes attenuation as it travels down the axis of the fiber. This is measured by a stress wave transmission factor, and a higher transmission factor leads to a larger transmission of stress to the other fibers [10, 18]. This transmission of stress causes higher elastic strain in the fiber bundle, which is then transferred to proximal fiber bundles, and therefore higher energy absorption of the composite as a whole. At the instant of impact, an initial compressive strain is created on the surface of the material [8]. This strain is given by the equation: V ε = (Eq. 2) c where V is the impact velocity, and c is speed of sound in the fiber, given by: 1 2 E c = (Eq. 3) ρ where ρ is the fiber density and E is the tensile modulus. In the case of an impact where the maximum fiber strength exceeds the maximum projectile energy, the waves in the fibers will have sufficient time to travel and disperse that energy over a larger area and the fiber will be able to deflect more, leading to a greater retardation of the projectile s penetration [20]. If the projectile s energy exceeds the maximum fiber strength, the principal yarns will fail before the longitudinal waves can disperse energy, leading to the successful penetration of the projectile. Zee and Hsieh [21] performed a study on Kevlar, carbon, and polyethylene fibers and found that 9

23 most of the primary fibers involved in the impact event were broken and the strained, unbroken fibers contribution was relatively insignificant Secondary Yarns The secondary yarns come into play as the principal yarns deflect, pushing or pulling secondary yarns out of plane. The methods of energy dissipation by the secondary yarns occur through stored strain energy and matrix cracking and delamination [10, 18]. According to Cheeseman et al. [4], the secondary yarns deform and exhibit a strain wave similar to those in the primary yarns and in turn, deform the yarns intersecting them. In situations where the impact velocity is high but not significant enough to completely penetrate the material, the secondary yarns on the front most layers will fail while the back layers will deform elastically. Also for higher energy impacts, some of the secondary yarns in close proximity to the primary yarns will fracture in the first stages of the impact. This occurs because these yarns experience strain almost equal in magnitude to that by the principle yarns [4]. The deformation of the secondary yarns through the thickness of the material contributes to the mechanism of cone formation and also promotes energy dissipation through elastic deformation. 2.2 Cone Formation and Elastic Response As stated earlier, deformation of the yarns initiates the formation of a cone. The cone, also known as backface signature (BFS), is a direct response of the material to the projectile. When a projectile impacts the fibers, there is a large increase in the transverse wave velocity which initiates the cone formation [10]. The fibers at the point of impact fail in tension, while the fibers on the back of the plate behave elastically and elongate, 10

24 although some fail from fiber pullout [4]. The back face signature shape is highly dependent on the fiber weave structure. Gu [22] showed that fiber reinforced laminates with a [0/90] n structure will exhibit a base area geometry with a major axis in line with the fiber axis of the last ply, whereas a plain weave fiber structure exhibited a pyramidal shape such as that in Figure 3. The front fibers deflect into the back layers elastically, failing in tension if the energy is great enough. Once the fibers fail, they rebound and form a cone on the front surface with an apex at the point of impact [4]. Figure 3. Pyramidal shape of back face signature formed during impact in plain weave fiber reinforced materials. Note: From Analytical Modeling for the Ballistic Perforation of Planar Plain-Woven Fabric Target by Projectile by Bohong Gu. Composites Part B. Vol. 34. p Copyright 2003 by Elsevier. Reprinted with permission. 11

25 Throughout the impact event, projectile displacement and cone height are taken to be equal and cone surface radius versus time is roughly linear [10]. The mass of the cone is given by M 2 = πr hρ (Eq. 4) where r is the radius of the cone base, h is the height of the cone, and ρ is the density of the material. Equation 4 can be applied to Equation 5 to determine the energy absorbed by the cone, where the velocity of the cone is the same as that of the projectile at the given time interval. 1 MV 2 2 EKE = (Eq. 5) where M is the mass of the cone and V is the velocity of the projectile and cone. The progression of the cone diameter can also be modeled using the following equation ( 1+ ε ) p p σ p 1 dσ ct = dε ρ ρ dε ε 0 (Eq. 6) where σ p and ε p are the stress and strain at which plastic deformation begins, respectively, and ρ is the density of the material. The cone radius after a certain time interval, t i, is given by r ti n i = = n=0 c tn t (Eq. 7) 12

26 Important variables that affect the cone characteristics were observed by van Hoof [23]. The author found that elastic modulus (both through-the-thickness and in-plane), through-the-thickness compressive strength, and in-plane strain-to-failure were critical in decreasing cone size, specifically that a low through thickness elastic modulus and high values for the other parameters produced a smaller cone. One phenomenon that takes place in the cone formation process in laminated materials is delamination of plies, which is a major energy absorbing mechanism. When considering the combination of fiber reinforced plastics with a ceramic strike face, Ravid et al. [24] found that the volume of the bulge at the back of the ceramic plate from the projectile and ceramic plate combined equals the backface signature of the backing panel. This observation is evident in Figure 4, which shows the angles between the projectile axis and bulge edge and corresponding volumes of the bulges. Figure 4. Schematic showing the volume of the bulge at the back of the ceramic and the equivalent volume bulge on the back face of the backing plate. Adapted from Ravid et al [24]. 13

27 Ravid and coworkers [24] also found that the resisting force of each layer of the backing panel can be determined via Equation 8 assuming force is constant over each area that is affected within the plies, which allows for the prediction of appropriate backing plate thickness when subjected to a certain threat level [24]. F i η i R = ( Eε i ) Y ψ S cosθ i (Eq. 8) sin ψ where E is Young s modulus, εiis strain, Y is the number of yarns of fabric per unit length, S is the cross section of yarn, θi is the angle between projectile direction and force in yarn, η i R is the distance from projectile axis to edge of affected area in the laminate, and ψ is the angle between projectile axis and edge of back face bulge. 2.3 Delamination and Matrix Cracking Delamination in a laminated composite occurs when the initial strain in a ply exceeds the fiber-matrix interface strength and the matrix cracks [10]. The extent of the matrix cracking is dependent on the matrix volume fraction and the fiber-matrix interface strength. The occurrence of matrix cracking also depends on the ductility of the polymer. Another way a composite absorbs energy is through inhibiting crack propagation by the presence of fiber reinforcement [25]. A strong fiber-matrix bond is required for structural load bearing. For high velocity impact events, a weaker fiber matrix interface is desirable, which allows delamination to occur and therefore allows the fibers adequate time to transfer loads and strain-to-failure [4]. Too high of an interfacial bond strength will produce a more brittle behavior upon impact [25]. Polyethylene materials reinforced 14

28 with polyethylene fibers are one example of FRTPs that exhibit good ballistic performance, yet also have good interfacial and fiber-matrix bonding [7]. As mentioned earlier when discussing cone formation in laminates, delamination occurs in parts of the cone. The matrix cracks first, followed by separation of the plies. The delamination shape depends mostly on the weave architecture and the ply stacking sequence. According to Harel et al. [7], the delamination energy is given by ny 2 πγ U delam = (Eq. 9) 2 where n is the number of delamination planes, y is the diameter of the delamination zone, and γ is the fracture surface energy of the matrix. Harel et al. [7] showed that the delamination energy contributed to nearly all of the energy absorbed by the plate, which led to the conclusion that delamination was the dominating mechanism for energy absorption. In the case of multiple impacts, the projectiles after the first impact encounter less resistance with each succeeding impact, due to the decreasing amount of delamination area. This can be linked to experiments by Sheikh et al. [16] where it was theoretically postulated that multiple, spaced-out thinner plates would perform better than one consolidated panel because the thinner plates absorbed more energy per unit area. The setup by Sheikh et al. [16] consisted of a 20 mm projectile impacting one large laminated sample and multiple separated laminates of [0/90] E-glass/epoxy laminates at 500 m/s. 15

29 For the trials involving multiple laminates, the gap between each laminate was varied from 10mm to 80mm. The authors determined experimentally that the energy absorption by multiple, spaced-out plates was less than that of a single, thicker plate. 2.4 Friction Fiber/Projectile Friction Another important mechanism of energy absorption in fiber reinforced composites incorporates the friction between the projectile and the fibers. It was found that energy was absorbed in the event of a projectile displacing fibers on a single fabric ply [26]. Fiber displacement is governed by the size and shape of the projectile. It can be hypothesized that a projectile would have a higher chance of penetration. This is supported by Tan and coworkers [27], where they observed that conical and ogival projectiles pushed the fibers away from each other, while flat-faced projectiles sheared the fibers, and hemispherical projectiles caused fibers to pull out. Different projectile shapes commonly utilized for high velocity impact testing can be seen in Figure 5. 16

30 Figure 5. Common projectile shapes encountered during ballistc impacts: (a) hemispherical, (b) flat-faced cone, (c) cylindrical, (d) spherical, and (e) ogival. The displacement of fibers by a projectile is also governed by the weave architecture. Chitrangad [28] determined that the area covered by a fabric should be between 0.6 and 0.95 percent of the total area to be effective at resisting a projectile. A tighter weave pattern provides more resistance to a projectile than a loose one, but fiber coverage area that exceeds the range will allow extremely limited fiber elasticity, resulting in a decreased amount of energy absorbed. In the case of projectiles displacing the fibers, the gap left by the projectile is typically smaller than the projectile itself [20]. It should be noted that a higher resin content constrains the fibers from slipping, therefore causing more projectile-fiber interaction and a higher energy absorption. In the case of highly oriented thermoplastic fibers such as polyethylene or polypropylene, it might be considered that the elevation of temperature at the instant of impact may have an adverse effect on the ballistic resistance of a composite. Alcock et al. [29] showed that the variation in temperature through the glass transition temperature 17

31 of polyolefin fibers during impact has a very small effect on the amount of energy absorbed even at temperatures well below the glass transition temperature. Testing consisted of plain weave polypropylene (PP) tape laminates impacted with a 9 mm projectile and a 5.39 mm fragment simulating projectile while the test temperatures varied from -40 C to 120 C. This suggests that highly oriented fibers are able to withstand transient thermal gradients and will not degrade to a point of lowering the ballistic resistance Fiber/Fiber Friction A significant amount of research has been conducted on comparing dry fiber with lubricated fiber of the same material. Studies show that at a constant velocity, a wet laminate will be perforated with laterally displaced yarns, while the dry laminate arrests the projectile [26]. This wetness can also be taken into consideration for friction between the fibers themselves. Fibers that do not slip against each other bear the load and therefore assist in energy dissipation. Briscoe et al. [30] performed a study on yarn-toyarn friction in three types of Kevlar fibers, namely as received, cleaned of all sizing agents, and chemically treated with polydimethylsiloxane. In this experiment, each type of Kevlar fiber was tested to determine its frictional value. As expected, the polydiethylsiloxane treated fibers had the lowest friction and the scoured fibers had the highest friction, therefore promoting more interaction between the fibers. This led to higher energy absorption that correlated to the fibers absorbing higher energy when subjected to ballistic testing. 18

32 2.5 Shear Plugging When a blunt shaped projectile (a cylindrical projectile, for example) cannot displace the fibers laterally through sliding and friction and the projectile s energy is greater than the maximum strength of the fibers, they will tend to shear or pull out. When the fibers shear in this manner, the material equivalent to the shape of the projectile cross-section will be pushed through. This is known as shear plugging and is most commonly observed with flat-faced projectiles. Projectiles of other geometries will cause fiber slip or fiber pull out [27]. Shear plugging occurs more frequently in brittle materials such as carbon/epoxy composites than in thermoplastic composites. This phenomenon is rarely witnessed at low velocities or with rounded projectiles. Sharp edges and high velocities also lead to the shearing of the fibers. Naik [10] found that the energy absorbed by shear plugging is based on the sheared distance, shear plug strength, and the area over which the stress is applied. This is given by Equation 10. E = Nh Sπ dh (Eq. 10) l where N is the number of plies sheared, S is the shear plugging strength, h l is the layer thickness, d is projectile diameter, and h is plate thickness. 2.6 Ballistic Limit Calculations on Backing Panels Specific models have been developed to predict the ballistic limit of fiber reinforced plastics subjected to impact events. Wen [31] developed equations for predicting V 50 for projectiles with geometries consisting of cylindrical, spherical, and ogival shapes. Florence [32] developed an analytical model for a hard ceramic strike face 19

33 backed by an aluminum backing plate. The author observed that the ceramic would crush in a conical shape, therefore applying a circular area load on the backing panel. In order to keep up with the progress of lightweight armor development, some variations evolved from Florence s model [24, 33, 34, 35]. Hetherington and coworkers [33] found that the Florence model provided a general idea of energy absorbed per unit areal density depending on the ratio of thicknesses of the ceramic face and the ductile backing panel. These authors applied the following equation to a ceramic plate combined with a glass fiber reinforced plastic backing panel to obtain theoretical values to compare against their experiments. V p 0.91 ( ) ε cs = f a M p 1 2 (Eq. 11) where M is projectile mass, ( a) p f is M p /( M p + (h 1 d 1 + h 2 d 2 ) π a 2 ) π a 2, a is a p + 2h 1, a p is the projectile radius, h 1 is the ceramic thickness, h 2 is the backing panel thickness, d 1 is the density of the ceramic, d 2 is the density of the backing panel, S is UTS* h 2, UTS is ultimate tensile strength of backing panel, and ε c is the breaking strain of backing plate. Hetherington and coworkers [33] conducted a series of tests to compare the initial and residual velocities of a 12.7 mm projectile when impacting 4mm, 6mm, 9mm, and 18mm thick ceramic tiles combined with 5mm, 8mm, and 10mm thick glass fiber reinforced plastic panels. All combinations were impacted at roughly m/s and gave a close correlation to the predicted values. Using the same equation as 20

34 Hetherington and coworkers, Navarro et al. [36] tested ceramic strike plates in combination with Kevlar and polyethylene backing plates and determined that the theoretical exit velocity values were rather conservative. 2.7 Static Indentation for Predicting High Velocity Impact Failure The factor of static penetration energy to ballistic impact has also been considered. Static indentation produces a larger amount of localized damage in a material. At very high velocities, impact damage is very localized [37]. At velocities below a fiber reinforced plastic s ballistic limit, Gama and Gillespie [37] found that damage closely resembled that caused by a shear punch test and could accurately be predicted from observing static indentation testing results [37]. Sun and Potti [38] developed a general equation predicting the exit velocity in ballistic testing utilizing the static penetration energy, initial projectile velocity, and projectile mass: V f 2 = V 2 i ESP (Eq. 12) m where V f is the exit velocity, V i is the initial velocity, m is the projectile mass, and E SP is the static penetration energy. A multitude of work has been completed regarding the high velocity impact response of fiber reinforced composites. Material response modes such as elastic deformation, fiber failure, delamination, matrix cracking, and cone formation were understood. The previously discussed concepts can be applied to fiber reinforced materials with either a thermoset or a thermoplastic matrix. For this study, a focus will be placed on characterizing those materials with thermoplastic constituents and an 21

35 understanding of the failure modes will provide insight into the suitability of FRTPs for high velocity impact resistance. 22

36 III. OBJECTIVES I. Understand the failure mechanisms and therefore energy absorption methods and damage progression within FRTPs subjected to high velocity impact. II. Evaluate performance and characterize damage within the FRTP plate and at a strike face/backing plate interface for projectile penetration and spall effect on the backing plate. III. Apply existing computational models to predict the damage progression in the backing plate. 23

37 IV. MATERIALS AND METHODS 4.1 Materials and Processing The materials chosen for this research included various fiber reinforced thermoplastics in both woven and unidirectional architectures to observe the effect that reinforcement and matrix properties have on energy absorption. Although the main focus of this research involved fiber reinforced thermoplastics, a fiber reinforced thermoset composite was also included in the testing in order to better understand and compare these systems. Therefore, three classifications of materials were analyzed: (a) glass fiber reinforcement in an epoxy matrix, (b) glass fiber reinforcement in a ductile matrix, and (c) polyolefin fiber reinforcement in a ductile matrix. The areal densities of the samples are given in kg/m 2. Due to confidentiality agreements, the specific names of the materials will not be disclosed Glass Fiber/Epoxy matrix Included only for comparison purposes to fiber reinforced thermoplastics, this thermoset composite consisted of glass fibers with an epoxy matrix. S-2 Glass was used as the fiber reinforcement and the matrix was an epoxy (SC-15). This composite was processed by Vacuum Assisted Resin Transfer Molding (VARTM). This material had a plain weave architecture and will be referred to as Glass/Epoxy. 24

38 4.1.2 Glass Fiber/Ductile Matrix This category included S-2 glass fiber reinforcement in a PE matrix with a crossply structure. Testing this combination showed how a projectile interacts with brittle fibers. The ductile matrix should permit more mobility of the fibers and therefore should allow the fibers more time to transfer the load. This material was composed of 85% fiber by weight and will be referred to as Glass/PE Polyolefin Fiber/Ductile Matrix High strain-to-failure fibers such as highly drawn polypropylene and polyethylene fibers were utilized. In the case of the self-reinforced PP and PE, the matrix material was present from the fiber processing as the fiber composition is composed of amorphous and crystalline constituents. Self-Reinforced PP. Two types of self-reinforced twill weave fiber architecture PP laminated composites were considered for this study. The first type of self-reinforced PP was supplied in two configurations one which was processed at a lower temperature and pressure and therefore had a lower bond strength (PP-1A) and one which was processed at a higher temperature and pressure that resulted in a stiffer, higher degree of consolidation material (PP-1B). The second type of self-reinforced PP was only supplied in high bond strength condition (PP-2). The processing of these samples was done by the manufacturer and PP-1 samples were supplied in approximately 5, 9.75, and 14.7 kg/m 2 (1, 2, and 3 lb/ft 2 ) configurations. 5, 9.75, 14.7, and 58.7 kg/m 2 (1, 2, 3, and 12 lb/ft 2 ) configurations were tested in this study, with the 58.7 kg/m 2 configuration achieved by bonding four 14.7 kg/m 2 panels with 3M acrylic VHB pressure sensitive adhesive tape. 25

39 Since the supplied PP-2 was thinner than the other materials, the pressure sensitive adhesive tape was utilized to achieve an areal density as close to 5 kg/m 2 as possible. This is the only configuration that was tested for PP-2. Self-Reinforced PE. The self-reinforced PE was constructed of ultra high molecular weight polyethylene. The material was supplied in two different configurations from the manufacturer as a cross-ply rolled mat. The variation in the two different configurations was due to dissimilar strain-to-failure values. The material with the lower strain-to-failure will be labeled as PE-1A, while the material with the higher strain-to-failure will be labeled as PE-1B. The samples for testing had areal densities of 5, 9.75, and 19.5 kg/m 2 and were fabricated by bonding individual plies in a heated press. The matrix content of the self-reinforced polyolefins is a function of the amorphous polymer constituent in the fibers and is difficult to quantify due to the variations in fiber processing and final laminate processing. 4.2 Testing Methods Impact testing methods of the fiber reinforced thermoplastics involved high velocity testing. This was accomplished by two methods: a gas gun propulsion system and a Universal Receiver that fires live ammunition [40]. The materials were tested to obtain the energy absorption at various velocities, and some materials were also tested based on the National Institute of Justice Standards [39]. Static indentation tests were conducted to determine if a failure mode correlation could be made to high velocity impact results. 26

40 4.2.1 Gas Gun Apparatus The gas gun uses compressed nitrogen for the actuator and compressed nitrogen or helium for projectile propulsion. The gun fires projectiles up to 38.1 mm in diameter at velocities up to 400 m/s using helium as the propulsion gas. Smaller projectiles require the use of a sabot, which is constructed of low density polyurethane foam. The sample mounting area is in a 12.5 x 12.5 x 72 chamber constructed of low-carbon steel. For these experiments, the samples were 6 x 6 in size and constrained on all four sides during testing. Oehler Model 35 chronographs placed before and after the sample registered the initial and exit velocities through polycarbonate windows in the chamber Universal Receiver Apparatus The Universal Receiver utilizes live centerfire ammunition rounds to simulate actual ballistic threats. The capabilities available in-house include the ability to fire the following projectiles: 9 mm,.44 Magnum, 7.62 mm, 12.7 mm, and 12 gauge shot shell. Projectile types and velocities used correspond to the NIJ Standards requirements [39], although numerous projectiles and velocities can be tailored to meet specific testing requirements. The fixture for testing can accommodate samples ranging from cm x cm to 50.8 cm x 50.8 cm in size. For this research, all samples tested were cm x cm in dimension and were positioned 49 feet from the receiver. The panels are constrained on all four sides by a fixture with infrared velocity detectors spaced out before the sample to record initial velocity. Chronographs are also placed after the sample to obtain an exit velocity if penetration occurs. The Universal Receiver testing setup can be seen in Figure 6. 27

41 Figure 6. Universal Receiver testing setup Static Indentation Testing Static indentation was conducted with a ½ diameter hemispherical indentor. PP- 1A and PP-1B were tested in 5 kg/m 2 areal densities according to ASTM D6264 test. Testing was conducted for a simply supported boundary condition in order to investigate a correlation to the ballistic testing results. 4.3 Procedure One experimental procedure involved subjecting the fiber reinforced thermoplastics of comparable areal density to high velocity impact with a spherical projectile at a consistent velocity in the gas gun. Four samples of each material PE-1A, PE-1B, PP-1A, PP-1B, PP-2, Glass/PE, and Glass/Epoxy were impacted with a 7.62 mm diameter spherical steel projectile with nominal mass of 2.04 g at velocities ranging from m/s. The reasoning behind this experiment was to demonstrate the performance of each material and determine which exhibited superior performance at a 28

42 given impact velocity. As the highest velocity achievable for a 2.04 g spherical projectile in the gas gun is around 400 m/s, it was initially predicted that not all of the samples would arrest the projectile. The failure modes for each sample were visually examined and compared to gain a qualitative understanding of the energy absorption. In order to better understand material response and failure modes during ballistic impact events, a second experimental procedure was conducted that involved testing at velocities well above the ballistic limit of each material. One sample from each of the six different materials was tested: PE-1A, PE-1B, PP-1A, PP-1B, PP-2 and Glass/PE. As with the first experiment involving the gas gun, the areal density of each sample was kept around 5 kg/m 2. The velocity was chosen to be between 680 m/s and 740 m/s to ensure the complete perforation of even the highest performing material. However, in this case, a 7.62 mm ogival projectile that consisted of a lead core and copper jacket was utilized. The mass of the projectile was 9.33 g. The purpose of this experiment was to obtain exit velocities which would determine the impact duration in each material, which could then be applied to equations to predict back face signature characteristics. Due to the applicability of the equations to the woven structure of PP-1B, only this material will be considered in the mathematical and computer modeling. The delamination area, BFS cone volume, and BFS cone height of each sample was measured with a caliper and angular calculations to determine the validity of the equations for fiber reinforced thermoplastics. Many of these materials were also tested according to the NIJ standards either as constituent laminates for Level III-A and III or combined with a hard strike face to 29

43 evaluate performance against armor piercing ammunition (Level IV and 12.7 mm AP). The main focus of these tests involved the exposure of thermoplastic composites to real world threats and the evaluation of failure mechanisms leading to projectile defeat. The test levels were as follows: Level III-A g.44 Magnum semi-jacketed hollow point projectile; Level III g 7.62 mm M80 projectile; Level IV g 7.62 mm M2 AP armor piercing projectile; and the 12.7 mm AP test used a g AP M2 armor piercing projectile. PP-1A, PP-1B, PE-1A, and PE-1B were subjected to NIJ Level III-A and III as stand-alone solutions. PE-1A was combined with a hard strike face, subjected to NIJ Level IV and 12.7 mm AP tests, and then underwent destructive evaluation to determine the effect a hard strike face had on the damage progression throughout the backing plate material. Both the front and back faces of the FRTP were observed and damage areas were analyzed with Image Pro Plus. In the case of PP-1A and PP-1B, static indentation was also conducted with ½ diameter hemispherical indentor. An attempt was made to find a correlation between static indentation performance and ballistic impact performance of PP-1B. PP-1B was chosen as extensive testing has been performed on this material and many properties are known. 30

44 V. RESULTS AND DISCUSSION 5.1 Gas Gun Testing Results These samples were impacted with a 7.62 mm diameter spherical steel projectile at velocities ranging from 295 m/s m/s. Table 1 shows the results for each sample tested. The average energy absorbed is given for four samples of each material type. Table 1. Testing results for impact velocities of m/s with 7.62 mm spherical projectile. Material Average Energy Absorbed/Areal Density (J/kg/m 2 ) Average BFS Delamination Area (cm 2 ) Average BFS Cone Height per Unit Thickness (mm/mm) Average BFS Cone Volume (cm 3 ) PE-1A* PE-1B* PP PP-1A PP-1B Glass/Epoxy N/A N/A Glass/PE *- All samples of that material arrested projectiles 31

45 Figure 7 shows the energy absorbed normalized to areal density to allow a fair comparison of the materials and their responses to the impact testing. The self-reinforced PE samples provided the most energy absorbed per unit areal density. Figure 7. Energy absorbed/areal density for 7.62 mm spherical projectile at m/s. * - Projectile did not penetrate Glass Fiber/Epoxy Matrix The average energy absorption of the Glass/Epoxy samples was calculated to be J. When normalized with respect to areal density, the energy absorption was found to be J/kg/m 2 (52.15 J/lb/ft 2 ). This value was approximately 3% higher than the 32

46 PP-2 material but well below the performance of the other fiber reinforced composites which initially showed a very similar performance to the PP-2 material. Upon visual observation of the samples, the Glass/Epoxy showed a significantly greater delamination area compared to a very localized damage area on the PP-2. The average area of delamination in the Glass/Epoxy samples was ± 52.4 cm 2. The delamination in the Glass/Epoxy sample can be seen in Figure 8. The BFS cone was almost nonexistent as most damage took place in-plane. Glass fiber reinforced epoxy materials typically exhibit more fiber breakage than fiber reinforced thermoplastics during high velocity impact events because the matrix does not deform as much as a thermoplastic matrix. This, in conjunction with the woven architecture, restricts the ability of the fibers to undergo deflection, hence energy absorption is mainly through fiber breakage and delamination. Figure 8. Glass/Epoxy post-impact delamination area on the back face of the sample. 33

47 5.1.2 Glass Fiber/Ductile Matrix The Glass/PE samples exhibited an average energy absorption of J/kg/m 2 (82.56 J/lb/ft 2 ), which was 58% higher than the Glass/Epoxy. Significant fiber breakage made up the majority of the front face damage with fibers protruding back away from the material. This material also exhibited the highest amount of secondary fiber damage than any of the tested materials. Back face damage was considerable with extreme fiber pull out and the samples showed an average cone volume of 6.9 cm 3, which was surpassed only by the PE-1 samples. The extent of damage that occurred was explained by the fact that the cross-ply structure allowed a higher distribution of energy across the sample. However, when the back face signature cone height was normalized to thickness, results showed that the Glass/PE sample BFS cone heights were higher than any other material. This could be explained by the tendency of thinner materials to undergo more bending than the thicker materials. Figure 9 shows back face signature height per unit thickness. 34

48 3.50 BFS Height per Unit Thickness (mm) PE-1A* PE-1B* PP-2 PP-1A PP-1B Glass/PE Figure 9. Back face signature height per unit thickness for samples impacted at m/s with a 7.62 mm spherical projectile. * - Projectile did not penetrate. Two samples contained intact fibers from middle layers that protruded out of the exit orifice. This appeared to agree with theory that fiber pullout would occur when a fiber reinforced material was impacted with a spherical projectile as edges are not present on the projectile, thus preventing shear forces [27]. The fibers were engaged, avoided failure, and were pulled out due to the spherical shape of the projectile. This could be explained by understanding that the material had an 85% fiber weight fraction, therefore not enough matrix material was present to hold the fibers in place during the impact event. In comparison, the self-reinforced materials had adequate fiber-matrix interfaces, thus reducing the probability of fiber pullout. 35

49 5.1.3 Polyolefin Fiber/Ductile Matrix Self-Reinforced PP. The PP-2 samples exhibited an average energy absorption of J and a normalized energy absorption of J/kg/m 2 (50.55 J/lb/ft 2 ). PP-2 absorbed the least amount of energy per unit areal density among all the tested materials. Visual observation after impact showed very localized damage with the exit aperture very similar in size to the entrance, as shown in Figure 10. The back face signature was taken to be a cone shape and had the lowest volume of all samples tested at 0.53 cm 3. Figure 10. PP-2 impact damage (left) and exit damage (right). The PP-1A samples had an average energy absorption of J with a normalized energy absorption of J/kg/m 2 (86.58 J/lb/ft 2 ) and the PP-1B samples had an average energy absorption of J with a normalized energy absorption of J/kg/m 2 (86.35 J/lb/ft 2 ). These values corresponded to approximately 71% and 70% more energy absorbed per unit areal density than the PP-2 material. The damage between these two types of PP was rather similar with the exception of more fiber 36

50 breakage on the exit side of the PP-1B samples, as can be seen in Figure 11. The impact point showed fiber shear due to the woven architecture and the resulting fiber displacement away from the frontal plane which is also shown in Figure 11. Figure 11. Side views of impact damage to PP-1A (left) and PP-1B (right). Both the PP-1A and the PP-1B samples produced very minimal BFS cones along with significant fiber failure. The reason for the low damage lies within the woven fiber architecture making it more difficult for the transverse waves to travel. The average volumes of the deformation cones were 1.28 cm3 and 1.32 cm3, which were 142% and 149% higher than the values for PP-2. In some cases on the distal side of the sample most noticeably in the PP-1A samples, as shown in Figure 12 fibers showed evidence of fiber wrinkling and primary fiber damage away from the exit point. 37

51 Figure 12. Fiber wrinkling on exit side of PP-1A sample. The energy absorption of PP1-A and PP1-B was approximately 21% less than the PE-1A and PE-1B which led to the postulation that these materials might provide better post-impact structural performance when compared to the other materials due to very localized damage and better than average energy absorption. Self-Reinforced PE. All of the PE-1A and PE-1B samples successfully stopped the projectiles at impact velocities as high as 352 m/s, exhibiting superior impact resistance when compared to the other tested materials. Impact point analysis on both types of PE-1 samples revealed typical fiber breakage, but wrinkling was also evident along the axis of the primary yarns. The wrinkling extended approximately 3 cm to 3.5 cm from the point of impact on all samples, as can be seen in Figure

52 Figure 13. Impact face damage of PE-1A (left) and PE-1B (right). In order to simplify the back face signature data, all the bases of the cones were taken to be of an elliptical shape as in Figure 14. This shape was due to the cross-ply fiber structure within the laminate and had a major axis parallel to the fiber direction in the surface ply on the exit side. Figure 14. Elliptical shape on exit side of cross-ply laminated sample of PE-1A. 39

53 The back face signature on both types of self-reinforced PE samples had cone base areas typically between 25 cm 2 and 34 cm 2, but the cone heights on the PE-1B samples were 2.83 mm higher on average. The average back face signature cone volume was 6.76 cm 3 for PE-1A and cm 3 for PE-1B. These values were 1230% and 1872% higher than that of PP-2. A comparison of the back face signature cone volumes can be seen in Figure 15. In the PE-1A and PE-1B samples, the projectile displacements (measured from the strike face to the front of the projectile) were approximately 8.51 mm and mm, respectively Back Face Cone Volume (cm 3 ) PE-1A* PE-1B* PP-2 PP-1A PP-1B Glass/PE Figure 15. Comparison of back face signature cone volume for 7.62 mm spherical projectile at m/s. * - Projectile did not penetrate. The strength and ductility of the self-reinforced PE samples demonstrated excellent resistance to high velocity impact through the modes of primary fiber engagement and strain, straining of the secondary yarns, delamination, and BFS cone 40

54 formation. The cross-ply laminated structure allowed the fibers to adequately distribute the loads and also strain extensively without failure as the spherical shape of the projectile did not create any areas for stress concentration. 5.2 Universal Receiver Testing Results Each sample in this experiment was impacted once with a 7.62 mm M80 projectile between 680 m/s m/s. The cost and availability of the materials limited the amount of testing, so more samples would improve the statistical significance of the results. The results from the testing are summarized in Table 2. Table 2. Testing results for impact velocities of m/s with 7.62 mm ogival projectile. Material Energy Absorbed/Areal Density (J/kg/m 2 ) BFS Delamination Area (cm 2 ) BFS Cone Height per Unit Thickness (mm/mm) BFS Cone Volume (cm 3 ) PE-1A PE-1B PP PP-1A PP-1B Glass/PE

55 Figure 16 compares the energy absorbed per unit areal density of the samples. The results show that the self-reinforced PE fibers were able to dissipate the most energy, as in previous testing conducted at m/s. 30 Energy Absorbed/Unit Areal Density (J/kg/m 2 ) PE-1A PE-1B PP-2 PP-1A PP-1B Glass/PE Figure 16. Energy absorbed per unit areal density for 7.62 mm ogival projectile at m/s Glass Fiber/Ductile Matrix The Glass/PE sample showed an energy absorption of J and a normalized energy absorption of J/kg/m 2 (79.07 J/lb/ft 2 ). The energy absorbed per unit areal density was 36% higher than the PP-1B sample. Through visual observation, it was concluded that more fibers broke at the impact point in this material than any other. Primary fiber breaking can be seen in Figure 17. This was due in part to the high velocity 42

56 of the projectile, but also to the fact that these fibers had a lower strain-to-failure than the thermoplastic fibers. The back face of the sample initially showed little evidence of penetration other than fractured fibers, but upon further inspection revealed a back face signature with a large elliptical base diameter and very low BFS cone height. The back face signature had a volume of 6.28 cm 3. This value was considered an outlier as the initial material thickness is one-third that of the other fiber reinforced composites and thus may have experienced bending during impact. This could have been the cause of the large delamination area which, unlike any of the other materials, increased from that in the lower velocity tests. Also, the projectile had displaced many fibers and had not fully engaged all primary fibers. This fiber displacement was an excellent example of how projectile shape and fiber orientation interact during impact events. Figure 17. Primary fiber failure from front side (left) and side view (right) in Glass/PE. 43

57 5.2.2 Polyolefin Fiber/Ductile Matrix Self-Reinforced PP. The PP-1B sample absorbed the least amount of energy per unit areal density at J/kg/m 2 (57.97 J/lb/ft 2 ). The PP-1A sample absorbed 20% more energy at J/kg/m 2 (69.66 J/lb/ft 2 ). Visual inspection confirmed that the impact damage on both types of PP very closely resembled the damage on both types of samples tested at 360 m/s. Both types absorbed less energy when subjected to an ogival projectile at around 700 m/s as opposed to spherical projectiles at 360 m/s. This was as expected as the projectile shape contributed to higher probability of penetration and less transferred energy. The volumes of the back face signatures of PP-1A and PP-1B were 0.41 cm 3 and 0.31 cm 3, which were among the lowest values for samples tested with the ogival projectile. Back face signature cone heights per unit thickness were observed to be the lowest of all materials. A comparison can be seen in Figure

58 1.20 BFS Height per Unit Thickness (mm) PE-1A PE-1B PP-2 PP-1A PP-1B Glass/PE* Figure 18. Back face signature cone height per unit thickness for samples impacted at m/s with a 7.62 mm ogival projectile. * - Glass/PE value can be considered an outlier as the initial material thickness is onethird that of the other fiber reinforced composites and thus may have experienced bending during impact. The PP-2 sample absorbed about 6% more energy than the PP-1B sample. In the previous case of the gas gun testing, the PP-1B absorbed substantially more energy than the PP-2 material. The energy absorption of the PP-2 material actually increased from J/kg/m 2 (50.55 J/lb/ft 2 ) to J/kg/m 2 (61.54 J/lb/ft 2 ) as the impact velocity increased from 337 m/s to 735 m/s. This material was the only one to exhibit this behavior. The impact damage imparted upon the sample at 735 m/s was identical to the damage found on the samples tested at 337 m/s, with the back face signature volume only 0.15 cm 3 less than at the lower velocity. This volume was roughly 8% lower than PP-1A and 22% higher than PP-1B, whereas during the lower ballistic velocity tests, PP-2 had a 45

59 significantly lower back face signature volume than both of the other PP-1 samples. A comparison of the back face signature cone volumes can be seen in Figure 19. The resulting decrease in damage in the self-reinforced PP materials from increasing the projectile velocity was attributed to more fiber shear behavior as the woven fibers could not undergo the amount of strain as the cross-ply materials such as the Glass/PE and the self-reinforced PE. Back Face Cone Volume (cm 3 ) PE-1A PE-1B PP-2 PP-1A PP-1B Glass/PE* Figure 19. Comparison of back face signature cone volume for 7.62 mm ogival projectile at m/s. * - See note from Figure 18. Self-Reinforced PE. The PE-1A and PE-1B samples showed normalized energy absorptions of J/kg/m 2 ( J/lb/ft 2 ) and J/kg/m 2 ( J/lb/ft 2 ), respectively, which amounted to 124% more than the PP-1B sample. The impact point on both types of PE samples showed fiber breakage from the high velocity of the 46

60 projectile. Back face signature was observable on both types of PE-1 samples and appeared to have a quasi-elliptical shaped base, although thicker samples would yield a more defined back face signature as more fibers would be able to interact during the impact event to resist penetration and transfer the load to surrounding fibers. The PE-1A sample had a back face signature volume of 1.91 cm 3 and the PE-1B sample had a volume of 0.81 cm 3. The impact and exit sites were very similar and both showed evidence of fiber breakage but also showed how the fibers were displaced around the ogival projectile in Figure 20. Figure 20. Similar damage on entrance (left) and exit (right) sides of PE-1A. When the volumes of the back face cones were compared between the samples impacted at a lower velocity with a spherical projectile and the samples impacted at higher velocity with an ogival projectile, multiple trends were observed. Figure 21 shows that increasing velocity and varying the projectile geometry from spherical to ogival resulted in a significant decrease in back face cone volume for PE samples. This test 47

61 performed exactly as expected due to the higher chance of penetration from changing projectile geometry from a spherical shape to an ogival shape and also from increasing the initial velocity which reduced the contact time between the projectile and the material. On the other hand, the self-reinforced PP samples exhibited rather small changes in back face cone volume size. The low amount of damage correlated to less interaction between the primary and secondary fibers, but this performance is expected at a velocity much higher than the V 50 due to inadequate time for the interaction to occur. A comparison of the Glass/PE results between velocities was inconclusive as the Glass/PE samples were almost one-third the thickness of the other fiber reinforced composites, thus leading to a higher tendency for bending and the possibility of skewed results. For this reason, the results from the Glass/PE testing were left out when comparing BFS cone volume size between the differing velocities. Qualitatively, as both velocities were above the ballistic limit for the Glass/PE, primary fiber failure appeared relatively consistent between the samples except for evidence of fiber displacement for the sample impacted at a higher velocity with an ogival projectile, thus reducing the amount of fiber engagement. 48

62 Back Face Cone Volume (cm 3 ) PE-1A PE-1B PP-2 PP-1A PP-1B Impact Velocity (m/s) Figure 21. Relationship between back face cone volumes for samples tested with 7.62 mm spherical projectile at ~ 350 m/s and samples tested with 7.62 mm ogival projectile at ~ 710 m/s. The initial high velocity testing results showed that the woven laminates sustained less damage than the cross-ply laminates. Also, the cross-ply laminates generally absorbed more energy than the woven laminates because the structure of the cross-ply laminates allowed the fibers to experience more strain, therefore increasing the time until failure leading to a more efficient load transfer. The fibers in the woven laminates were subjected to more shear forces at the weave intersections, forcing quicker fiber failure. Also note that the fiber tows in the self-reinforced PP materials PP-1A, PP-1B, and PP- 2 appeared to have more of a tape-like geometry, whereas the fibers in the other materials seemed to have roughly circular cross-sections. 49

63 5.3 NIJ Standard Testing Universal Receiver NIJ Level III-A This test called for a.44 Magnum Semi Jacketed Hollow Point projectile at 436 m/s. Unlike the previously used projectiles, this projectile was meant to expand upon impact. This provided insight into the synergistic effects of a continuously changing load area due to material interaction. The testing results were obtained for PP-1A, PP-1B, PE- 1A, and PE-1B. The 9.75 kg/m 2 PP-1A sample was impacted twice and failed to arrest the projectile both times. Both projectiles created similar damage to the previously tested PP-1A samples. Due to the hollow pointed geometry of the projectile, a portion of the material was sheared out at the strike face. Both PP-1A and PP-1B with areal densities of 14.7 kg/m 2 were impacted five times and no penetration occurred. Delamination was comparable for both types and the back face signatures from each impact were relatively small with cone heights of around 11.5 mm. Figure 22 shows the back face signature from the multiple impacts and also shows inter-laminar shear that aided in the delamination. There was a small amount of interaction between damage zones as the points of impact were at least 5 cm apart. 50

64 Figure 22. Side view of PP-1A showing back face signature and inter-laminar shear/delamination after NIJ III-A test. Both PE configurations that were tested had areal densities of 5 kg/m 2, and each configuration successfully arrested the projectile. The PE panels were impacted only once as the delamination and back face signature were significant enough that any following impact tests would have lead to inconclusive data. Strike face damage was considerable as the primary fibers interacted with the secondary fibers causing large strains in the fibers before failure, which led to the creation of a large depression. Figure 23 shows the depression left in the strike face and the fiber wrinkling on the strike face around the point of impact. Material in the shape of a cone with a 2 cm radius also flowed back out away from the material. The back face signature was extreme and can be seen in Figure 24. Accurate measurements of the back face signature could not be obtained as the damage was constrained by the mounting fixture. 51

65 Figure 23. Front face of PE-1B showing radial fiber wrinkling from the point of impact after NIJ III-A test. Figure 24. Back face signature of PE-1B following NIJ III-A test NIJ Level III This test called for a 7.62 mm M80 projectile at 847 m/s. PP-1A and PP-1B were impacted five times, while PE-1A was impacted once. Both types of PP-1 had areal densities of kg/m 2 (12 lb/ft 2 ) while the PE-1A had an initial areal density of 17.1 kg/m 2 (3.5 lb/ft 2 ). The PP-1A successfully arrested two projectiles, but allowed the following three to penetrate. The exit velocities for the third and fourth projectiles were 52

66 118 m/s and 229 m/s, respectively. There was no reading on the exit velocity of the fifth projectile, but the previous data leads to the assumption that the residual velocity was well above 229 m/s. The increase in the exit velocities was attributed to the increase in delaminated layers after each shot. As with the NIJ Level III-A tests of the PP-1A, this test exhibited numerous delaminated layers, a gradually sloped back face signature that was constrained by the mounting fixture, and it arrested the projectile. At the point of impact, the failed fibers slightly protruded out away from the material. This was due to the stress wave traveling through the fibers and fiber flow back toward the impact point forcing material out. Apparently, as the projectile s displacement increased, the extent of deformation also increased. This test also showed that the projectile s pointed geometry caused a slight change in angle during the continuous engagement of primary fibers. The impact progression changed from a higher fraction of displaced fibers to a higher fraction of engaged fibers, hence the greater extent of fiber breakage and delamination towards the end of the projectile s path. This phenomenon would not occur as extensively if a steel core projectile were used as the core would be much more detrimental to the soft fibers. The PP-1B sample successfully arrested all five projectiles with damage almost symmetrical to the damage in the PP-1A sample. The PE-1A sample (initially 17.1 kg/m 2 ) was penetrated on the first test with an exit velocity of 641 m/s, but was reprocessed to approximately 19.6 kg/m 2 (4 lb/ft 2 ), and subsequently stopped two projectiles at 873 m/s and 877 m/s. There was significant back face damage, though it was noted that the damage to the strike face was remarkably 53

67 similar to the previous 5 kg/m 2 (1 lb/ft 2 ) PE-1 panels tested with the same projectile. The point of impact showed a cone of damage with a 34.4 mm radius protruding away from the material a distance of 3.43 mm. This led to the assumption that front face damage was not dependant on thickness for this material. The back face signature can be seen in Figure 25. The main back face cone had a height of 50 mm, while the second bulge interacted with the constraint fixture and could not be accurately measured. There was moderate interaction between the two damage zones although the projectiles impacted 13 cm apart. Figure 25. Back face signature of PE-1A following NIJ III test NIJ Level IV This test called for a 7.62 mm M2AP projectile at 878 m/s. The PE-1A samples were expected to undergo significant cone-shaped damage at the interface between the strike plate and the backing plate. Backing plate front face damage areas varied greatly, 54

68 but a general trend was observed that a 61% larger damage area was present when strike plate thickness was decreased by 42.5%. This could be attributed to greater strike plate material deflection for the thinner material versus less deflection in a thicker material. This theory was reinforced by the fact that the cone damage depth was 141% higher, inferring that the thinner strike plate material deflected into the backing plate. Also, a difference in projectile yaw and production of spall during the impact event would affect the damage area. Due to the previously mentioned variables, a more in-depth study regarding projectile interactions with the strike plate material would provide more concrete evidence of performance. This damage depended on the thickness of the strike plate; however, the actual entry hole diameter for all samples with the same strike plate material varied only slightly from 9.0 mm to 11.5 mm with no noticeable dependence on the thickness of the strike plate. This variation was most likely due to projectile yaw produced during the impact event. In one case with a harder strike plate material, an entry hole diameter of 13.9 mm almost twice the projectile s original diameter was observed. This large entry hole could again be accredited to more extreme projectile deflection caused by the increased material hardness. The average back face signature cone height was 71% higher for the same thickness backing plate coupled with the thinner strike plate. The larger back face signature with a thinner strike plate corresponded to the previously mentioned larger front face damage leading to the conclusion that more of the backing plate was engaged by the projectile/spall. 55

69 mm AP Test This test used the 12.7 mm APM2 projectile at 870 m/s, and PE-1A was again used as with the NIJ Level IV test. When compared to the Level IV test with the same thickness strike plate, the front face damage area of the backing plate was 405% larger. This value correlated well to impact energy as it experienced a 382% increase. Depth of the front face damage on the backing plate was 495% larger than that of the 7.62 mm AP test, which was again expected. Back face signature of the backing plate could not be accurately quantified as interference from the mounting fixture took place. 5.4 Static Indentation Damage characterization from the static indentation results revealed a generalized trend that could be applied to ballistic impact. The less consolidated material, PP-1A, showed a tendency to more readily undergo delamination and bulging; whereas PP-1B, the more rigid material, had less delamination. This observation agreed with the results from the 7.62 mm spherical, high velocity impact where PP-1A had a slightly larger back face cone radius and lower cone height compared to PP-1B, which had a smaller cone radius but higher cone peak. Also, out of plane bulging on the front face from static indentation corresponded with high velocity impact as PP-1A had a more gradual drop from the impact point, while PP-1B had a sharp decline from the impact point. The energy absorbed by PP-1B with a hemispherical indentor was calculated to be 78.7 J. 56

70 VI. MODELING In order to simplify some of the complications with modeling various types of fiber reinforced thermoplastics, PP-1B will be the focus of the following models. This material was chosen as it has a woven structure, which allows more interaction between primary and secondary fibers. Also, penetration occurred in both cases of ballistic impact where it exhibited relatively similar failure modes. 6.1 Back Face Signature Modeling Equation 6, presented by Naik et al. [10], was used to predict the cone base radius on the back of PP-1B tested at 684 m/s. ( 1+ ε ) p p σ p 1 dσ ct = dε ρ ρ dε ε 0 After determining the transverse wave velocity was 238 m/s, the predicted cone radius was found to be mm. The actual cone radius was mm. Predictions for back face signature base radius after testing with a spherical projectile at approximately 350 m/s showed that the base radius should be around mm. The actual base radius was measured to be approximately mm. While not as close as the prediction for testing at the higher velocity, this value is reasonable considering the number of experimental variables such as plate bending and elastic material response. 57

71 6.2 Static Indentation vs. Ballistic Impact Utilizing Equation 12 from Sun and Potti [38], which uses the energy of static penetration for PP-1B with the initial velocity of the gas gun testing, a relationship between the predicted and experimental values was found. The graph of the load versus displacement curve for a 5 kg/m 2 sample indented with a hemispherical indentor is shown in Figure kg/m 2 PP-1B Simply Supported 16, , , Load (N) 10, , , , , Displacement (mm) Figure 26. Static indentation data for PP-1B. The ability to accurately predict a material response to ballistic impact through static indentation testing allows more efficient testing. The predictions and experimental values for PP-1B can be seen in Table 3, following Equation 12, recalling that the 58

72 projectile mass was kg and the static penetration energy was 78.7 J. Predictions for intermediate velocities were also included, although no experimental data was collected. V f = V 2 i 2 E m SP Table 3. Experimental vs. predicted exit velocity values from static indentation data. Sample Projectile Experimental Initial Velocity (m/s) Predicted Exit Velocity (m/s) Experimental Exit Velocity (m/s) PP-1B mm sphere PP-1B mm sphere PP-1B mm sphere PP-1B mm sphere Theoretical 7.62 mm sphere Theoretical 7.62 mm sphere Theoretical 7.62 mm sphere PP-1B-1A 7.62 mm ogival It is evident that the predictions overestimated the exit velocities of the tests by 10-20%. It should be noted that as the projectile geometry is altered, the initial load versus displacement curve would change for static indentation. In this case, the higher velocity tests were modeled from the load versus displacement curve in Figure 26. The difference between the predicted exit velocity and experimental exit velocity for testing at 684 m/s with an ogival projectile was attributed to the absence of an ogival-shaped indentor to use for static indentation. Given the fact that precisely similar samples will 59

73 always produce slightly different results, this prediction method was within an acceptable level for these tests. 6.3 Impact Simulations A model for PP-1B impacted at 350 m/s with a spherical projectile was analyzed using LS-DYNA. PP-1B was modeled using MAT162. This material model was chosen because it is for a woven or unidirectional laminated fiber reinforced composite that accounts for material softening after damage initiation. The specific properties of the material could not be disclosed due to confidentiality agreements. The projectile was modeled using MAT020, the model for a rigid material, as the projectile would experience negligible deformation, and hence energy absorption, from impacting a thin fiber reinforced thermoplastic. A top view of the mesh can be seen in Figure 27. The plate was constrained on all four sides to simulate the actual sample fixture for impact testing. The plate had 12,500 total 3-D hexahedral elements with each element around the point of impact having a size of The square area of impact contained 625 elements on the surface with 2,500 total elements through the thickness of the impact area. The element size increased from the edge of the area of impact to a maximum of 8.14 at the edge of the plate. Four elements were used through the thickness of the plate. The element size was 1.02 for the spherical steel projectile with 677 total elements. The mass of the projectile was 2.04 g and the diameter was 7.62 mm. The contact interface between the projectile and plate was that of an eroding single surface. The damage in the model matches the back face damage of the sample accurately, which can be seen in Figure

74 Figure 27. Top view of the mesh used to simulate impact of PP-1B. Figure 28. Left: Model of PP-1B damage at 350 m/s. Right: Actual PP-1B damage at 350 m/s. 61

75 It should be noted that the failed primary fibers protruding out of the actual sample would not show on the simulation image. The BFS cone height in the model was approximately 3.10 mm while the average actual height for the PP-1B samples was recorded to be 2.91 mm. The cone radius at the base of the model was mm, while the actual radius was approximately mm. In the model, the exit velocity of the projectile was 240 m/s, which was 35 m/s higher than the actual observed exit velocity. A simulation for the Universal Receiver testing of PP-1B at 684 m/s was also analyzed and it was observed that the simulated BFS cone height was 1.90 mm, while the actual measured BFS cone height was 1.64 mm. The radius of the cone at the base of the simulation was 9.90 mm while the actual cone radius was mm. A comparison between the model and the actual sample can be seen in Figure 29. Figure 29. Left: Model of PP-1B damage when impacted at 683 m/s. Right: Actual damage of PP-1B at 683 m/s. 62

76 The visual difference between the images in Figure 28 is due to the failed primary fibers protruding out of the sample. This simulation appeared to reinforce the conclusions obtained from the previous simulation at 350 m/s. The exit velocity in the simulation was found to be m/s while the actual exit velocity was m/s, which shows that this model was extremely accurate with respect to the exit velocity. The simulated BFS cone heights in both models slightly overestimated the actual measurement by 0.2 to 0.3 mm. The only variable changed between the models was the velocity, so this model could be applied to numerous events with considerable confidence. Table 4 shows the numerical values predicted and observed. Table 4. Comparison of simulated and observed values for PP-1B. Projectile Simulated Observed Simulated BFS Observed BFS Simulated Observed Exit Velocity Exit Velocity Cone Height Cone Height BFS Cone BFS Cone (m/s) (m/s) (mm) (mm) Radius Radius (mm) (mm) 7.62 mm sphere mm ogival For the NIJ Level III tests, an existing model was compared to the actual test results. The model accurately showed the front face bulge that occurred as the material rebounded away from the point of impact and also showed the gradually sloping back face signature. Delamination was visible in the simulation initially along the projectile 63

77 path. As the projectile decelerated through the material, the primary fibers transferred more energy before failure. This allowed more secondary fibers to act in the energy absorption process which led to a larger extent of delamination. The extensive delamination can be seen occurring in the lower 1/3 of the simulation in Figure 30. Figure 30. Left: Computer simulation cross-section exhibiting PP-1B response to ballistic impact at 867 m/s. Right: Cross section of PP-1B showing projectile displacement and material delamination at impact velocity of 867 m/s. 64