Experimental Studies on Mitigating the Risk of Air Blast Loading

Blast loading events that arise from the detonation of explosives pose a severe threat to the lives of civilians and military personnel alike. Such dangers typical of a detonation event include but are not limited to an intense, sudden initial pressure spike, extreme temperatures due to the burning of gases released by the explosive, and damage to the integrity of surrounding structures. It is therefore the purpose of the studies detailed in this manuscript to investigate various methods of mitigating the dangers posed by shock loading, as well as to investigate a novel impact mitigation device. To address the danger presented by high velocity glass fragments generated by windows that have failed due to shock loading, a study was conducted to evaluate the effectiveness of coated laminated safety glass panels’ ability to contain glass fragments when subject to shock loading over a range of temperature conditions. Using a shock tube apparatus, fully clamped specimens were loaded under room temperature (25 °C), low temperatures (-10 and 0 °C), and high temperatures (50, 80, 110 °C). For each experiment, the incident and reflected shock wave pressure profiles were recorded and three-dimensional Digital Image Correlation was used to analyze high-speed images and compute the full-field deformation, in-plane strains, and velocities during the blast-loading event. A post-mortem study of the sandwich specimen was performed to investigate the effectiveness of such materials under different temperatures to withstand these shock loads. The composite panel showed great endurance during the blast loading for temperatures from 0 to 80 °C, however was unable to contain glass fragments at -10 °C and 110 °C. A new system was designed to mitigate the impact forces during a collision using shock loading. The device consists of a cylindrical composite bladder sealed on one end by an inflation valve and on the other by an aluminum sheet of variable thickness. The bladder is pressurized and as an impactor nears the device, it strikes a striker-needle which ruptures the aluminum sheet, thus producing a shockwave just prior to impact. This produced shock wave decelerates the impactor and creates momentum (impulse) opposing that impulse transmitted from the impactor. Drop weight experiments were performed to show the applicability of this anti-shock device in reducing the momentum of the incoming body. A range of variables including needle length, bladder pressure, impact velocity, and drop mass were tested to better understand the processes involved. Time lapse photography coupled with 2D Digital Image Correlation (DIC) was used to obtain the striker full field motion data during the drop-weight experiments. It was found that the device effectively mitigates impact for higher impact velocities and for higher bladder pressures, decreasing peak loads during impact by up to 58% and energy imparted on the structure by 40%. An experimental study was also conducted to examine the induced pressure from the interaction of a planar shock front and perforated plates under fixed and freestanding boundary conditions using the shock-tube facility. Two series of experiments with variations in the blockage ratio, net hydraulic diameter, shapes, and sizes of the orifices, were conducted. During each experiment, pressure histories caused by the interaction of the incident shock wave with the plates were recorded. During the fluid structure interaction time, the side-view images of the targets were recorded using a single high-speed camera to identify the motion response of each plate configuration. The influence of varying the incident shock wave Mach number on the pressure profile was examined under clamped boundary conditions. The experimental results show that as the blockage ratio of the freestanding perforated plate decreased from 100% to 65%, the reflected peak pressure decreased by 26%, and the maximum impulse imparted to the perforated plate decreased by 33%. During the fluid-structure interaction process, as the blockage ratio of the freestanding perforated plate decreased from 100% to 65%, the plate’s momentum, velocity and kinetic energy decreased by approximately 60%, 61%, and 84% respectively. Finally, investigations were conducted to investigate the performance of different surface roughness (Ra) of 1018 mild low carbon steel panels under blast loading. Specimens were machined to have three different surface finishes of 0.8, 1.4, and 5.0 μm. The shock tube apparatus was utilized to generate controlled blast loadings on simply supported specimens. For each experiment, incident and reflected shock wave pressure profiles were recorded, and three-dimensional Digital Image Correlation was used to analyze high-speed images and compute the full-field deformation, in-plane strains, and velocities during the blast loading event. In addition, another high speed camera was utilized to record the side-view deformation images and this information was used to validate the data obtained from the 3D stereovision DIC technique. The results indicated that the impulse imparted to the plate decreased as the surface roughness increased from 0.8 μm to 5.0 μm. Due to this impulse reduction along with high surface roughness, the plates demonstrated a decrease in back face deflection, in-plane strain and out-of-plane velocity.

In order to better understand some of the topics that will be discussed in this chapter and those that follow, one must understand the difference between active and passive mitigation techniques. An active mitigation system is any mitigation technique that relies on a certain outside stimulus in order to be initiated [1,2]. A basic example of this is an airbag system in an automobile, an impact mitigation system, which relies on input from sensors that detect a crash impact, thereby initiating deployment of the airbag [3][4][5][6][7][8]. Another example would be the water deluge systems utilized in UK offshore oil and gas platforms, which rely on sensors to detect gas leaks, and thereby initiating the release of water droplets into leak areas in order to mitigate the risk of a gas explosion [9]. Current innovations in this area include the work conducted by the Defense Research Council, which has investigated panels filled with cells containing isolated chemical agents which act as a propellant when they interact. Thus, if the panel is subject to a blast loading effect, the cells rupture and the interaction is initiated, thus generating a gas buffer which opposes the blast wave [1]. While they can be effective, it can be seen that the processes and equipment required for active mitigations can be quite extensive and complex, and therefore costly. Furthermore, it is necessary in most cases that the system be initiated immediately upon impact, a feat which is difficult to achieve for most sensors. For these reasons, the more simple passive mitigation techniques are often preferred [2].
Passive mitigation systems do not rely on sensor input or other outside stimulus to be initiated, and instead utilize structural properties built into the design of the system [2]. This makes the passive system simpler and eliminates any reaction time between the event and the response. Common techniques employed in passive systems include impedance mismatching, sacrificial cladding, blast deflection, and blast and shockwave disruption [2]. One noteworthy example of a passive blast mitigation technique which utilizes impedance mismatching techniques consists of elastomeric layers containing air-filled microspheres [10]. Another good example of a passive blast mitigation technique which uses a water cavity between the incident blast wave and the structure to be protected was proposed by Shin et al. The authors identified that placing a water cavity delays the shock front and minimizes the magnitude of initial peak shock pressure by 40% [11].
With regards to blast mitigation, it is important to recognize that the phenomena generated by explosion events which pose a threat to human life are not limited to the high peak pressures that are a result of the initial explosion [12][13][14].
Clearly these peak pressures are harmful to humans and are thus their mitigation a topic of investigations detailed in this work, however succeeding events such as high temperatures produced from burning gasses, secondary debris and fragments, direct impact, and the compromise of structural integrity of surrounding structures also pose a significant threat and thus are also addressed in this work.
Chapter 2 of this manuscript addresses the problem of debris generated by harm to structures during a blast loading event, detailing the investigations and analysis of a passive mitigation technique in the form of sandwich structures designed to contain glass fragments generated from blast-loaded windows. Sandwich structures have proved useful in Naval and Aerospace applications due to their high stiffness to weight ratio and energy absorption capabilities [15][16][17][18][19][20][21][22], and thus it can be considered that similar configurations will have similar qualities when applied to glass structures.
According to Wang et al, when sandwich structures are subjected to high-intensity impulse loadings, the core materials play a crucial role in the dynamic behavior and overall structural response [15]. The core properties assist in dispersing the mechanical impulse that is transmitted into the structure, and thus protect anything located behind it [15]. As concluded by many studies, the majority of human harm caused by shock-loading events is not a result of the shock-wave directly, but of unsecured objects such as glass shards that are propelled at high velocities by the shockwave [23][24][25]. Monolithic glass panes, necessary as they are, are often the culprit in these sorts of injuries due to the fact that they offer little resistance to air blast loads. in Nano-structured porous materials [43] and reticulated foam [44]. While all previous studies revolve around fixed perforated plates, it is the purpose of this work to expand upon previous knowledge by observing the behavior of free-standing perforated plates.
To do this, the identical perforated plates with varying perforation specifications were tested for both fixed and freestanding conditions, and the results observed. It was determined that pressure profiles at various distances from the plate behaved similarly for freestanding plates as with fixed plates.
Finally, Chapter 5 of this manuscript investigates a passive blast mitigation technique which holds the added benefit of being cost effective and easy to implement. This chapter details a series of blast experiments conducted on steel panels with varying degrees of surface roughness. Prior literature has explored the problem of shock wave interaction on the boundary layer over a rough wall in transonic circular bump geometry [45,46]. However, shock being normally loaded or impinged onto a rough surface structure (as is typical with an explosive event) has not previously been investigated even though it is of importance in any treatment of shock wave attenuation. A 2-D numerical model, which is simulating interactions between a blast wave and a V-shaped or a cone-shaped structure, was developed by W. Peng et al. [47]. Normal, oblique, and Mach stem reflections are types of reflection when a blast wave impinges on a surface. The Normal reflection occurs when the incident angle is 0°. The angle between the incident shock front and the reflecting surface of a structure is defined as the incident angle [47].

Introduction
Most post-investigations of incidents involving explosions have cited that the majority of human casualties were not caused by the air blast wave or bomb fragments them-selves. Rather, it was fragments of objects coming from glass windows, walls, and other unsecured bodies that were responsible for the majority of human injury [1][2][3]. Monolithic glass panes, necessary as they are, are often the culprit in these sorts of injuries due to the fact that they offer little resistance to air blast loads. Thus, the development of blast-resistant windows has been a topic of interest among many researchers all over the world.
There are many different types of blast-resistant panels, including clear, wired, tempered, and laminated safety glass (LSG) panels [4]. In comparison to most of the commercial glass panels, LSG panels have been proven to have better blast and impact resistance and hence have been the central focus of research on blast resistant windows [5][6][7] of the interlayer rather than failure at the supports to achieve a desired level of protection. Many researchers have studied the dynamic properties of glass [18,19] and PVB materials [20,21] to help design better glass resistant panels.
A recent investigation by P. Kumar and Shukla reported that a sandwich configuration consisting of a PVB interlayer, two glass panels, with outer protective films was the most effective at mitigating blast-loadings compared to other glass systems, such as tempered glass, wired glass, and LSG [5]. This promising sandwich structure needs more research to investigate the bounds of its utility. Thus, in this study we aim to evaluate the blast performance of this sandwich structure in a variety of temperature environments. A shock tube facility generated the dynamic shock loading imparted to the glass composites while a series of pressure sensors and highspeed cameras were used to obtain pressure profiles and images of the specimen during loading at the different temperatures. Three-dimensional Digital Image Correlation (3D DIC) was used to analyze the high-speed images and compute the full-field deformation, in-plane strains, and velocities during the blast-loading event.
The composite panel showed superior blast mitigating property for temperatures ranging from 0 to 80 °C. The polymeric thin sheet coating and PVB both contribute in containment of the glass fragments and withstand the blast load. The failure of the sandwich panel at -10 °C is attributed to the glass transition of coating material and the failure at 110 °C is likely due to the tearing of the coating by the glass fragments.

Experimental Procedures
In the present study, the performance and dynamic behavior of sandwich specimen under shock loading was evaluated. To verify consistency and repeatability in the experimental results, at least three blast-loading experiments were performed at each temperature environment. Experiments were carried out at temperatures of -10, 0, 25, 50, 80 and 110 °C.

Sandwich Specimen Geometry and Materials
The sandwich specimens used for shock loading experiments were prepared by adhering a protective film (provided by XO ARMOR® with a thickness of 0.279 mm) on both of the outer faces of the laminated safety glass (LSG) panel using a special chemical adhesive XO® bond. Figure 1a shows a detailed schematic for the sandwich specimen configuration. The LSG panel used in this study consists of two clear glass plates (each ply had thickness of 3.14 mm) made out of Soda-Lime-Silica Glass, which were bonded with 0.76 mm-thick transparent PVB interlayer as shown in Fig.   1a, b. The overall dimensions for the sandwich specimens were 305 mm long, 305 mm PVB interlayer was first cut to the correct size and sandwiched between two clear glass plates. The sandwiched glass assembly was then passed along the production line where it moved through an oven while the air was being pressed out. After that, it was heated and pressurized in a computer-controlled autoclave for 24 hours. This treatment process yields good bonding strength and gives the glass plies clear optical properties.
The XO® bond penetrates the glass and forms a permanent, nano-level molecular bond between the glass and the XO ® film.

Blast Loading Under Different Temperatures
The experiments were conducted using the shock tube apparatus, which can produce transient blast loading. The shock tube facility at the University of Rhode Island, shown in Fig.2, had already been developed and has the capability to generate a controllable shock wave loading on the target specimen up to a 2.2 MPa overpressure [22,23]. In the present study, the shock tube had an overall length of 8m.
It consists of three main sections: high pressurized (driver) section, low pressure (driven) section, and muzzle section. Both driver and driven sections had diameter of 0.15 m and they were separated by a diaphragm. The muzzle end was the final section of the shock tube facility and had diameter of 0.07 m.

Fig. 2 Shock tube facility
During the experiments, the driver section was pressurized with compressed Helium gas, creating a pressure difference across the diaphragm that is composed of a variable number of Mylar sheets. Thus the pressure needed to rupture the diaphragm is controlled by the quantity and thickness of the Mylar sheets. When the pressure reaches its critical value as determined by the Mylar, the diaphragm ruptures, forming a supersonic shock wave that travels towards the target specimen. The theoretical assumptions that were used to describe the gas flow in the shock tube and the detailed compressible gas flow equations have been previously established in the literature [24,25].
In order to investigate the actual loading conditions in windows, sandwich specimens were held under fully clamped boundary conditions during blast loading. A detailed fixture support design is shown in Fig. 3a. The inner dimensions of the clamping fixture (window frame) were 267 mm x 267 mm. Rubber supports were placed between the steel metal clamping supports and the sandwich specimen to avoid breaking any glass during the clamping process. The sandwich specimen was placed in the robust supports and positioned 0.1 mm away from the end of the muzzle section.
The shock tube end muzzle was aligned so that its axis coincided with the center of the front face of the sandwich specimen. The blast loading was applied over a central     Fig. 5b. For low temperature testing, the sandwich specimen was cooled down in a freezer for about an hour up to a minimum temperature of (-20 °C). The sandwich specimens were sealed from the freezer's environment so no humidity was accumulated during the cooling process. Once desired temperature was achieved, the sandwich specimen was placed in the support fixture (inside the shock-tube facility) and tested after a certain calibrated waiting time. It is important to note that the sandwich specimen began to warm up after it is removed from the freezer and thus, a series of calibration tests were conducted to ascertain how long the specimen would take to reach the temperatures desired for testing (-10, 0 °C). In order to perform blast loading experiments at −10 and 0 °C, the sandwich specimen was held in room temperature for approximately 2 min. and 5 min., respectively as shown in Fig. 6.  The inter frame time between the recorded images was 50 micro-seconds. This photography system consists of two [Photron SA1] high-speed digital cameras, which were located behind the sandwich specimen (the detailed experimental setup is shown in Fig. 7). The resolution of these cameras was selected to correspond with the exposed blast area (152.4 mm x 152.4 mm) of the sandwich specimen. Two high intensity light sources were used to illuminate the sandwich specimens due to the short exposure times during the experiments. The high-speed photography system first was calibrated to ensure accurate measured results throughout the blast loading event. This was done by using the calibration grid plate and procedures provided by Correlated Solutions. A non-contact 3-D Digital Image Correlation (DIC) technique used images with a speckle pattern from the high speed cameras to obtain the full field deformation, velocities, and strains of each sandwich specimen during loading. Flat paint was used to create a random, high contrast speckle pattern on the surface of each sandwich specimen in order for the DIC technique to work. Figure 7 shows a real captured image of the speckle pattern. The DIC analysis was then performed using VIC-3D software from Correlated Solutions.

Tensile Testing of Coating Material under Different Temperatures
In order to characterize and correlate the response of the thermoplastic material used for coating the LSG panels, quasi-static tensile tests on XO ARMOR® protective films were performed at the same desired temperature environments as the shock loading experiments. To achieve the desired range of testing temperatures around the film specimens prior to and during testing, an environmental chamber (330 mm long × 381 mm wide × 762 mm high) made of thermal insulation foam was designed and fitted to the tensile testing machine. A schematic of the experimental setup is shown in Fig. 8. For testing the protective film under high temperatures (50, 80, 110 °C), a heating element with variable DC power supply (0-60 V) was utilized to control the amount of heat generated. A circulation fan was used to ensure proper air-circulation.
By controlling the voltage supplied to the chamber, different steady state high temperatures were achieved. Film specimens were heated as shown in Fig. 8a for approximately 40-50 minutes until the desired temperature was reached. For low temperature tensile testing, dry ice (CO 2 ) was placed in small containers to create a large surface area as well as achieve high sublimation rates inside the chamber.
Different steady state low temperatures were achieved by controlling the amount of dry ice and air flow rate of circulation fan. Film specimens were cooled down as shown in Fig. 8c for approximately 12 and 20 min until the desired temperature was reached. In the present study, the temperature was measured (using a K-type thermocouple) directly on samples until their temperature stabilized at the desired level. Typical temperature-time plots for tensile specimens at high and low temperatures are shown in Fig. 8b, d respectively.

XO-ARMOR Response
The effect of temperature on the stress-strain curve for the film coating is shown in Fig. 9. This figure denotes that the XO-ARMOR materials showed both possibilities of ductile and brittle failure like most thermoplastic polymers. The temperature increase leads to a decrease in elastic modulus, reduction in tensile strength, and an increase in ductility. Conversely, brittle fracture was favored at lower temperatures.

Room Temperature Response
In the current work, the applied shock load was experienced by the sandwich specimen during fluid-structure-interaction (FSI) time, the early stage of blast loading.

Effect of Temperature on Sandwich Specimen Response
The temperature effect on stress wave propagation through the sandwich specimen during FSI time (up to t = 200 μs) was investigated in the present study. The stress wave propagates with lower wave speeds at higher temperature environments as compared to room and low temperature environments as shown by Fig. 14a-  The temperature effect on the mean back-face deflection history is shown in Fig. 15.
As discussed earlier in Fig. 13a, the room temperature specimen showed a maximum deflection of 35.3 mm at t = 2000 μs and then began to rebound. For the 0 °C experiment, the composite structure showed a maximum back-face deflection of 29.7 mm at t = 1800 μs and began to rebound. As the temperature decreased to -10 °C, the sandwich panel reached deflection of 14.2 mm and then began to fail at t = 700 μs.
The expectation was that the specimen at lower temperatures responded quicker due to  The full-field of in-plane strains (ɛ xx ) on the back-face of the sandwich specimen at maximum deflection or at failure for temperatures ranging from -10 to 110 °C are shown in Fig. 17. For -10 °C (Fig. 17a), the sandwich specimen was reached 6% in-plane strain and then failed. This is in-line with the results of tensile experiments on the coated film which showed brittle failure at -10 °C when strained to 5%. Figure 17 shows that the in-plane strain of the sandwich specimen increased with increasing temperature. The composite panel showed great endurance during the blast loading for temperatures ranging from 0 to 80 °C (Fig. 17b-e). In-plane strains of 9, 12.8, 16, and 19.3% were experienced by the coating at temperatures of 0, 25, 50, and 80 °C respectively. At 110 °C (Fig. 17f), the panel showed 23% in-plane strain at the time of failure. This strain is much lower than the failure strain observed in the 110 °C tensile experiment on coating film. Thus it can be inferred that this failure in the coating material was due to the tearing of the film by the glass fragments, which is due to the failure of PVB section at higher temperatures.

Fig. 17
Full field in-plane strain (e xx ) at maximum deflection or at failure of sandwich specimen at different environment temperatures The real time images captured during the experiment were studied to further investigate the failure process at 110 °C and -10 °C. From the images (Fig. 18a) we can observe that many small glass fragments were flying out of the sandwich specimen, showing that the PVB interlayer was not able to contain them (PVB material loses strength at high temperatures). The de-bonding of glass fragments from PVB interlayer might have facilitated the rupture of the coating film by the glass fragments. However, as seen in (Fig. 18b) the PVB interlayer at −10 °C was able to hold a substantial amount of the small glass fragments together even after catastrophic failure of the coating material. Figure 19 shows how these glass fragments were contained by the PVB interlayer. This signifies the critical role of PVB interlayer in the performance of the sandwiched structure.

Conclusions
This study focused on the dynamic response of laminated glass composite, which is coated with XO-Film, under combined shock loading and extreme temperatures.
The major findings of this study are as follows: • There was substantial glass fragmentation at all temperatures. However, the coating material and PVB interlayer enabled structural integrity, controlled the damage propagation, and prevented glass pieces from flying away at temperatures ranging from ~0 to ~80 °C. In which case both interlayer and coating materials survived the shock loading with no catastrophic failure.
• PVB interlayer at −10 °C was able to hold a substantial amount of the small glass fragments together even after catastrophic failure of the coating material.
• The failure at 110 °C is likely due tearing of the coating by the glass fragments.
• Due to the temperature-dependent material properties of the protective film coating and PVB interlayer, the sandwich specimen showed an increasing trend in back-face deflection and in-plane strain with increasing temperature.
• The fragmentation process of glass continues to the point of peak deformation.
This is a unique energy dissipation mechanism offered by this construction of sandwich panel.
• For good performance of this type of sandwich structure geometry, both the interlayer and the coating film should have good ductility and strength at the target temperatures.  provide alerts that are intended to assist drivers in avoiding or mitigating a crash. Such systems can reduce the injury and fatal accidents by 36% and 59% respectively [9].
Passive safety systems that are installed in a vehicle, such as the pedestrian airbag system and the active hood lift system, protect a pedestrian's head during contact with a vehicle [10,11]. Choi et al developed a methodology for estimating the safety benefits of integrated pedestrian protection system based on simulations and showed it is capable of reducing pedestrian fatalities by approximately 90% [12].
It is, therefore, the purpose of the work detailed in this chapter to investigate a new method of impact mitigation, the results of which may not only be applied to some future form of civilian protection similar to those mentioned above, but provide insight into the physical phenomena present in the mitigation of impact using opposing shock waves. This paper presents a novel device designed to release a shock wave just moments before an impactor collides with it, and details the experiments conducted in order to understand the device's impact mitigation performance. The device consists of a cylindrical composite bladder sealed on one end by an inflation valve and on the other by an aluminum sheet of variable thickness. The bladder is pressurized and as an impactor nears the device, it strikes a striker-needle which ruptures the aluminum sheet, thus producing a shockwave just prior to impact. This produced shockwave decelerates the impactor and creates momentum (impulse) opposing that transmitted from the impactor. To test the effectiveness of the device, dynamic loading is replicated using an Instron 9210 drop-weight impact tower, with various drop tests using a range of weights and impact velocities were conducted.
Other variables tested were the length of the striker-needle, that is, response time, and bladder pressure. Impact velocity was easily determined for all tests using the Instron´s data acquisition software, and time lapse photography coupled with 2D Digital Image Correlation (DIC) was used to obtain the striker's full field motion data during the drop-weight experiments. The DIC technique is a recent noncontact optical method for analyzing full-field shape, motion and deformation [13]. This technique uses images with a speckle pattern from the high speed camera. The Photron Fastcam Viewer (PFV) software was employed to synchronize the camera and record the images during the experiments. A speckle pattern was placed directly on the striker surface in order for the DIC technique to work. The DIC analysis was then performed using VIC-2D software from Correlated Solutions.
Through experimentation and subsequent analysis, it was determined that the device is indeed effective at mitigating impact. It was found that as the striker-needle length increased, impactor impulse decreased, for example, as the needle length was increased from 6 mm to 18 mm; impactor impulse was reduced by 37%. It was also found that the anti-shock device caused a greater load reduction for higher impact velocities. Furthermore, the device is more effective when the bladder pressure is increased. Finally, the device was able to decrease peak load by 58%, and energy imparted to the structure by 40%.

Experimental Setup and Procedures
A series of low impact velocity experiments were conducted using a drop weight testing machine to understand the influence of the produced shock load on mitigating the impact loading of the striker.

Anti-Shock Production Set-up
The proposed system design consists of a special cylindrical thin-walled rubber tube that has a 76.2 mm outer diameter (OD) and 254 mm length as shown in Fig 1. The rubber tube composite material used in the present study was provided by PEGA Air Springs Company. The system device preparation and the main parts are shown in where, r is the radius of the circular end of the cylinder and L is the length of the cylinder. The calibrated pressure values and the potential energy stored in the device are listed in Table 1.
The striker-needle design shown in Fig. 2 was applied to penetrate the aluminum foils as the impactor nears the pressurized device, thus initiating the opposing shock, which resists the impact load. The needle was designed to have an adjustable length of penetration and the impactor had a flat surface with circular cross section of 76.2 mm diameter.

Experimental Approach
The first series of drop weight experiments was conducted to understand the influence of the needle penetration length on the anti-shock wave production. In these experiments, the impactor was given the same initial height and the bladder's pressure was held constant. After the determination of the onset of the shock release of the device, the second set of experiments was performed to investigate the effectiveness of this device in mitigating different impact energies. In all these experiments, the internal pressure in the anti-shock device was kept constant. Lastly, the third set of experiments was loaded with the same impact energies, but the bladder's pressure differed in order to produce different levels of anti-shock. The experimental details are listed in Table 2.

Drop Weight Impact Experiments
The dynamic loading was implemented by a drop weight impact tower apparatus (Instron 9210), as shown in Fig. 2  The mitigation device was held inside the testing machine enclosure and was fully clamped during experimentation as shown in Fig. 2. The positioning of the device inside the enclosure allowed it to be oriented in such a way that a high speed photography system could be employed during testing.
During the experiments the total mass of the impacting system has to be taken into account. Table 3 shows the mass of all components contributing to the impact.
The needle lengths negligibly contributed to the impactor mass.
where E is the desired potential energy, m is the mass of the drop weight, and g is the acceleration due to gravity.
After the drop height was determined, the drop tower velocity was determined.
The device was placed in the testing machine enclosure and the cross head was lowered until it came into contact with the device. Figure 2 shows the device placed in the supports with the striker/needle assembly in contact with the device. The velocity sensor must be adjusted so that the velocity flag attached to the crosshead is in line with the bottom of the sensor. With the sensor adjusted, the number indicated on the scale was taken as a datum point and the calculated drop height was set from the datum. The crosshead was raised to the appropriate height and the device was removed. A velocity test was then completed to ensure that the proper velocity was reached. Impact velocities were checked against a calculated velocity determined by Before experiments were performed, the data acquisition system was configured. Each tup has a calibration factor that must be inputted into the software.
The system was configured using the correct calibration factor for the 10,000 lb tup.
After the calibration factor was entered, the sampling rate was properly chosen. The sampling rate determines if the entire event is captured. The data acquisition system records 8192 data points regardless of the sampling rate, therefore, it is important to know the duration of the impact event. For the given study, the event duration was approximately 50 ms. A sampling rate of 81.92 KHz was chosen as this corresponds to 100 ms allowing for a proper margin of safety. For more information on the test machine description, parts functions, running preparations, and the operational specifications, see reference manual [14].

Effectiveness of the Produced Mitigating Shock
In an effort to better understand the various factors that affect this device's ability to mitigate impact loading, a range of factors were independently tested.

Loading
The first series of drop weight experiments investigates the optimal release time of the shock (needle penetration length) on the mitigation of impactor energy.
Three needle lengths were utilized (ℓ 1 = 6 mm, ℓ 2 = 12 mm, and ℓ 3 = 18 mm) as shown in Fig. 3a in order to control the onset of the shock release. The smallest needle was designed to release the shock late in the impact history after the impactor has already made contact with the bladder; the second longest needle was designed to release the shock quickly after the impactor had contacted the bladder and lastly the longest needle was designed to release the shock prior to the impactor contact. The drop tower was outfitted with a mass (m 1 ) of 8.58 kg and drop height (h 1 ) of 0.3 m (see Fig. 3a).
Initially, the impactors were given the same potential energy (E o ) of 25.3 J. The bladder's pressure (P o ) was kept constant at 345 KPa. The impulse reduction from the anti-shock mechanism is shown in Fig. 3c.
This was simply obtained by integrating the pressure-time data of the first impact pressure (see the dashed box in Fig. 3b). As discussed above, when the needle length was increased from 6 mm to 18 mm, the maximum impulse resulting from the first impact decreased from 3500 Pa.s to 2200 Pa.s. Therefore, the anti-shock device under these conditions reduces or mitigates the impactor impulse on the bladder by approximately 37%.

Influence of Impact Velocity
After determining the optimal needle length such that the anti-shock is released prior to impact, the second set of experiments was performed to investigate the influence of impactor velocity on the effectiveness of the mitigating shock device.  increased under a constant inner system's pressure condition.
With higher impact velocities, the anti-shock mechanism has a greater load reduction. This is due to the impactor approaching the pressure release at higher velocities. Thus, getting closer to the pressure release sooner and experiencing higher emitted shock pressures. This load reduction can be better seen represented by the impulse as shown in Fig. 4e. In experiment (a 2 ), the impactor impulse on the bladder was reduced by approximately 37%, whereas, experiments (b 2 , and c 2 ) showed load reduction of about 12%.

Influence of Inner System's Pressure
As discussed earlier, the mitigating pressure release property of the device at lower impact velocities had little effect in mitigating the incoming body (see experiment c 2 ) as compared to shocks generated at higher impact velocities, which showed greater load reduction. One way to increase the effectiveness of the produced mitigating shock at low impact velocities is to increase the internal pressure of the device. To test this, experiments were carried out at different inner system pressures in comparison to previous experiments to produce higher levels of anti-shock.
Three inner system's pressures (P o ) were selected to be 345, 517, and 690 KPa  The load reduction from the anti-shock mechanism is shown in Fig. 5d.
Experiments (a 3 , b 3 , and c 3 ) showed a load reduction of 5%, 32%, and 58%, respectively. In other words, when the internal pressure of the device was increased by 100% (experiments c 3 ), the anti-shock mechanism showed a greater load reduction of 58%. Therefore, mitigating the effect of normal impact loadings under lower impact velocities was achieved by producing higher levels of anti-shock.

Impactor Energy Mitigation
The applicability of the anti-shock device in mitigating the velocity and kinetic energy of the incoming body (impactor) is discussed in this section. Experiment c 3 is given as an example to explain the mitigation process. The velocity and kinetic energy of the impactor as a function of displacement are plotted in Fig. 6. A 2D-DIC technique was used as validation during the dynamic event and the data is presented as solid circles in Fig. 6a, b. For purposes of comparison with experimental values, the velocity and corresponding kinetic energy of the impactor unhindered by the antishock is presented as a solid black line in Fig. 6a, b. The real time images of the incoming impactor are shown by Fig. 6c.
Recall that in experiment c 3 the anti-shock device had a stored potential energy of 800 J (see Table 2). The impactor impacted with a velocity of v e = 1.  Table 4.

energy mitigation as shown by series III. This mitigation effect increases linearly with
increasing the potential energy of the device at a constant kinetic energy of the impactor as shown in Fig. 7b. This means that as the potential energy of the device increases, the mitigation effect also increases by the same percentage. The first is the impactor velocity. With a higher impactor velocity, the percentage of mitigation increases nonlinearly. The second is the potential energy of the device.
With a higher potential energy of the device, the increase in the percentage of mitigation is proportional. Therefore, this device is ideal for situations where the impact velocity is high and/or the potential energy of the device is high.

Conclusions
A new system was designed to mitigate the impact forces during a collision using shock loading. Drop-weight impact experiments were performed to determine the efficacy of this new mitigation technique. This technique has been found effective in mitigating the kinetic energy imparted on to the structure and in reducing the impactor peak load during collision. The following is the summary of the investigation: • The first series of drop weight experiments investigated the optimal needle length (such that the anti-shock is released prior to collision) on the mitigation of impactor load/energy. For the experiments conducted, the device mitigates impact more effectively as needle length increases indefinitely.
• For the second series of experiments, which investigated the effect of impact velocity on the effectiveness of the device, it was found that with higher impactor velocities the anti-shock mechanism produced a greater load reduction.
• Mitigating the effect of normal impact loadings under lower impact velocities was achieved by producing higher levels of anti-shock. The results indicated that when the stored potential energy of the device was increased from 400 J to 800 J, the maximum impulse imparted to the impactor by the produced shock increased by approximately 160%. In addition, the results indicated that the anti-shock device with 800 J of stored energy delays the time of impact loading three times more as compared to anti-shock device with 400 J.
• Overall, the device effectively mitigates impact for higher impact velocities and for higher bladder pressures, decreasing peak loads during impact by up to 58% and energy imparted on the structure by 40%.

Introduction
Blast mitigation techniques such as perforated structures are the focus of numerous studies because of their relevance in attenuating blast waves and reducing its pressure and damaging effects. Britan et al. [1,2] showed that the shock wave attenuation can be attained by orifice plates and grids, the inclination angle of the barrier, its width, opening ratio, and the incident shock wave Mach number determine the shock wave attenuation. Ram and Sadot conducted numerical studies to develop constitutive relations and validate experimental results [3,4]. Zhou and Hao carried out numerical simulations to study the effectiveness of blast barriers for blast load reduction. The results showed that a blast barrier not only reduces the peak reflected pressure and impulse on a building behind the blast barrier, but also delay the arrival time of the blast wave. An approximate formula was derived to estimate the reflected pressure-time history on a building behind a blast barrier [5].
The use of mild steel perforated plates as a passive mitigation system has been widely studied. Langdon et al. [6] observed that the effect of the blast loading on a structure is reduced when a perforated plate is placed on the path of the blast wave traveling down a tube, especially when the plate is well spaced apart from the target. Moreover, blockage ratios of 25% and 50% may not affect the target plate's mid-point deflection, however, with blockage ratio above 50%, the target plate mid-point deflection for a given impulse as well as the threshold impulse of the target plates are reduced. In addition, the separation distance has been found to influence the performance of perforated plates in blast wave shielding. Langdon et al. [8] reported that the mid-point deflection is reduced with an increase in separation distance. They also found that when the blockage ratio is increased to 75% above, a significant decrease in the target plate mid-point deflection is obtained for a given impulse and irrespective of the separation distance. Similar pressure-mitigating behaviors have been observed in Nano-structured porous materials [9] and reticulated foam [10].
Furthermore, it has been demonstrated that fixed barriers can reduce both static pressure and stagnation or reflected the pressure of the shock wave, but free obstacles such as dust reduces static pressure, however increases the reflected pressure [11].
The purpose of the study detailed in this chapter was to expand on previous Upon analysis of experiments on freestanding, it was seen that as the blockage ratio of the fixed perforated specimen decreased from 100% to 65%, the reflected peak pressure decreased by 26%, and the maximum impulse imparted to the specimen decreased by 33%. Similar results were observed in freestanding specimens. This represents the data obtained from the pressure transducer located 15 mm from the specimen. However, pressure transducers located further away (55 mm and 175 mm) observed similar trends. Furthermore, tests were conducted on specimens with hydraulic diameters (D h ) ranging from 4 mm to 7.1 mm. It was determined that the hydraulic diameter did not have any effect on the initial reflected peak pressure. Of the factors tested, only the blockage ratio affects reflected peak pressure. Finally, as the perforation perimeter of the specimens' perforations decreased, it was observed that higher pressures were present in the build-up regions with respect to the level of the initial peak reflected pressure.

Experimental Procedures
In the first part of this study, shock wave loading experiments on free standing perforated plates were conducted to evaluate the reflected pressure pulses and the motion response for all the targets with and without perforations. The second part of this study was on clamping all the targets inside the shock-tube to evaluate the reflected and transmitted pressure pulses for all the targets. To verify consistency and repeatability in the experimental results, at least three experiments were performed on each plate configuration.

Perforated Plate Specimens
In the present study, the blockage ratio (BR) is simply defined as the area ratio of the solid area of the perforated plate to the exposed area. Three different blockage ratios were selected for this study. A laser perforating technique was utilized to create a uniform round, square, and triangular perforations on the plate specimens. The numbers of holes in the plate, the size of these holes, and the distance between them have been used as design variables to obtain a fixed blockage ratio. The hydraulic diameter (D h ) is defined as (D h = 4A/P) where A is the total open area in the plate and P is the total perimeter of the holes in the plate. The areal density of all the targets was 38 kg/m 2 . A schematic of all targets used in this study is shown in Fig.1 and a summary of their perforations geometry and properties is listed in Table 1.

Application of Shock Load
The shock tube facility at the University of Rhode Island, shown in Fig.2 a, has already been developed and has the capability to generate a planar shock wave with a controlled overpressure level on a target plate. The shock tube has an overall length of 8 m. It consists of three main sections: high pressurized (driver) section, low pressure (driven) section, and muzzle section. Both the driver and driven sections had a diameter of 0.15 m and were separated by a diaphragm. The muzzle end is the final section of the shock tube facility and has a diameter of 0.0381m, which is also the diameter of the loading area. During experiments, the driver section was pressurized with compressed Helium gas, which created a pressure difference across the diaphragm. When the pressure reaches a critical value, the diaphragm ruptures, forming a shock wave that travels towards the target plate. In the present study, the diaphragm was made up of 1 ply of Mylar sheet which has various thickness of (5, 7, and 10 mil), depending on the peak shock pressure required. The typical incident peak pressures were 0.3, 0.41, and 0.55 MPa, respectively.

Shock Tube for Free-Standing Conditions
The focus of this study was to evaluate the reflected pressure pulses and the motion response for all the targets with and without perforations. The plate specimen was inserted inside the muzzle end perpendicular to the shock tube axis and positioned flush to the muzzle exit (as shown in Fig.2 b). It should be noted that the diameter of the plates was made exactly same as the inner diameter of the shock tube muzzle.
In order to investigate the pressure characteristics of the incident and reflected shock waves (upstream from the perforated plates), three pressure transducers (PCB CA102B) were flush mounted at the muzzle exit. These sensors are denoted as 1, 2, and 3 in Fig.2    where, u x is the x direction velocity and ds is the areal element of the plate.

Shock Tube Modification for Clamped Conditions
The shock tube was modified to measure the pressure history of the incident, reflected, and transmitted waves upstream and downstream from the perforated plates.
A detailed fixture design for clamping the perforated plate inside the shock-tube is shown by Fig. 5a.   In order to ensure that no reflection from the shock occurs due to the modified fixture, blank tests were carried out by firing the shock tube and recording the pressure profiles. For this purpose, the apparatus was tested without porous samples. A diaphragm configuration of one stacked ply of Mylar sheet (thickness of 0.178 mm) was used to generate an incident peak pressure of about 0.4 MPa. As the incoming wave leaves the end muzzle and enters the modified fixture, it is clear that there is no reflected shock peak as recorded by pressure sensors at CH2 and CH3 (see Fig. 7a).
This can be better represented by the incident pressure profiles and their calculated areal impulse during the first millisecond as shown in Fig. 7b, c respectively.

Experiments with Free Standing Boundary Conditions
The focus of this study was to evaluate the reflected pressure pulses and the motion response for all the targets with and without perforations. To verify consistency and repeatability in the experimental results, at least three experiments were performed on each plate configuration.

Effect of Plate's Blockage Ratio on the Reflected Pressure Profile
In this study, the effect of the plate's blockage ratio on the reflected pressure   The overall time period of the load acting on the plates (reflected pressure) was the same for different blockage ratios. The time period is defined as the time at which the pressure becomes equal to the atmospheric pressure. The time period of the reflected pressure was about 6.0 ms. The pressure reduction can be better seen as represented by the areal impulse imparted to the plate as shown by Fig.8 (b) For further distances at sensors located 55 mm and 175 mm away from the perforated specimens, the effect of plate's blockage ratio on the reflected peak pressure was still significant as shown by Fig. 9a, b respectively. This indicates that the effect of plate's blockage ratio is highly non-localized, and that effect was not vanished at greater distances upstream. There is no doubt that the net opening area in the plates has an effective influence on the reflected peak pressure.

Size of Perforations Effects
The upstream pressure histories that were measured by the pressure transducers (1, 2 and 3) are shown in Fig. 10a-   specimens, the existence of the pressure buildup region was still significant as shown in Fig. 10b. For large distances upstream of the perforated plates (175 mm away), the specific perforation sizes of the plates had no influence on the build-up region (vanished) as shown in Fig. 10c. This shows that the effect of different perforation sizes is highly localized, and that effect vanishes at greater distances upstream. These results can be better seen as represented by the areal impulse imparted to the plates as shown by Fig. 11. For the sensor located 15 mm away from the perforated specimens (see Fig. 11a), perforated plates (4, 5, and 6) obtained a maximum areal impulse of 1500, 1434, and 1373 Pa.s from the shock loading respectively. The results indicated that the plate with smaller circular perforations (plate 6) showed 9% decrease in the maximum areal impulse as compared to plate 4 which had larger circular perforations. Therefore, there was no doubt that, the net hole perimeter (or the hydraulic diameter) of the perforated plates had an effective influence on the reflected pressure build-up region.

Shape of Perforations Effects
The perforated plate (   These results can be better seen as represented by the areal impulse imparted to the plates as shown by Fig. 13. For the sensor located 15 mm away from the perforated specimens (see Fig. 13a), perforated plate (6, 7, and 8) obtained a maximum areal impulse of 1373, 1289, and 1199 Pa.s from the shock loading respectively. The results indicated that the plate with triangular perforations (plate 8) showed a 13% decrease in the maximum areal impulse as compared to plate 6 which has circular perforations. Therefore, there is no doubt that, the net hole perimeter (or the hydraulic diameter) of the perforated plates has a strong impact on the pressure build-up region.

Motion Response of Perforated Plates
The high speed side-view images during the fluid structure interaction time of plates (1, 2, and 8), which have different blockage ratio are shown in Fig. 14. The shock wave is propagating from the right side of the image to the left side. Time t= 0 ms corresponds to the beginning of the event where the shock wave impinged on the plates and the final image for each experiment represents the plate's motion at t= 0.6 ms. Since the compressibility of the gas only has a significant effect on the pressure profile at the beginning of the shock wave loading process [12]. The fluid-structure interaction process was almost over by the time of (t= 0.6 ms) because the gas pressure has dropped to about 40% of its peak value (see Fig. 8a). After this time, most of the load on the plates comes from the wind (movement of the gas particles) behind the shock wave front.

Experiments with Clamped Boundary Conditions
Inside a straight conduit, it has been known that when a planar shock wave collides with a perforated obstacle two processes take place at once. One part of the shock wave was reflected head-on from the perforated plate and the other part was transmitted through the open area of the plate. The first part referred to shock wave reflection from obstacle and the second part as shock wave diffraction through the perforation generating a non-steady flow behind it [1]. A set of experiments were conducted with clamped boundary conditions to evaluate the reflected and transmitted pressure pulses of the same samples of porous medium (Plates 4-8) using the modified shock tube described in (Section 2.2.2). The effect of the incident shock wave Mach number on the pressure history was also studied. Recall that, a total of four pressure transducers were used to measure the induced pressure histories. The first two pressure sensors (CH1, CH2) were used to investigate the "upstream flow" of both incoming and reflected waves, and the other two pressure transducers (CH3, CH4) were used to measure the "downstream flow" of the transmitted waves as they pass through the barrier.

Effects of Perforation Sizes and shapes
The pressure traces upstream and downstream of the perforated plates (4-8) are shown by Figs. 18 and 19 respectively at a constant incident shock wave Mach number of 1.76. Figure 20 shows the transmitted pressure profiles for further distance downstream from the perforated plates. The main observation from these pressure profiles was drawn in detail in the following sections. For better understanding, the results were also represented by the areal impulse.

Reflected pressure Profiles
The generated shock wave had an average incident peak pressure of approximately 0.3 MPa and an average reflected peak pressure of about 0.75 MPa (see The initial reflected peak pressures were the same (0.75 MPa) for different plate's perforation sizes and shapes, and were a function of only the blockage ratio of the plate. Afterwards, a pressure build-up region on the perforated plates was observed up to t = 1.2 ms due to the high resistance to the fluid flow (see Fig. 18a, b). The resulted pressure build-up region for plates (4, 5, 6, 7, and 8)  Therefore, there was no doubt that the net hole perimeter (or the hydraulic diameter) of the perforated plates has an effective influence on the pressure build-up region.

Transmitted Pressure Profiles
The When the shock wave passes through the barriers, a transmitted peak pressure of approximately 0.2 MPa was observed (see Fig.19a, b). Recall that the generated shock wave had an incident peak pressure of 0.3 MPa. This means that as the transmitted wave exits the perforated plate, its strength reduces compared to the initial incident shock wave. In this case, the shock wave was reduced (mitigated) by 33%. It should be noted that the data of the blank test (no plate or BR=0%) was presented as a dotted black line in Fig. 19a, b for purposes of comparison with experimental values. In comparison with the transmitted shock waves measured at CH4, Fig. 20 indicates that the transmitted shock wave had the same initial peak pressure of 0.2 MPa, but it showed a longer exponential decay period of approximately 1.5 ms. This perforated plates, the transient pressure showed a tendency to approach steady state and to reach equilibrium amplitude. This was consistent with previous findings of Ben-Dor et al who reported that for larger distances downstream the specific shape of the perforated plate had practically no influence on the recorded pressures as long as the porosity is kept at a constant value [1].

Effect of Incident Shock Wave Mach Number
In order to study the effect of the incident shock wave Mach number on the pressure history, experiments have been carried out at three shock Mach numbers, i.e., 1.76, 2.0, and 2.4. Figure 21a shows the incident pressure profiles with different incident peak pressures of 0.3, 0.41, and 0.55 MPa. The reflected pressure profiles measured for plate 1 (BR=100%) at different incident shock levels are shown in Fig.   21b. These experiments (Fig. 21) were used as a baseline for the performance of other specimens as a function of incident Mach number (or incident pressure). The experiment for plate 6 (BR=65%) was taken as a typical example to explain the effect of the incident shock wave Mach number on the reflected and transmitted pressure history. Figure 22 shows the reflected pressure profiles at CH2 for different incident shock levels and Figure 23 presents the transmitted pressure profiles at CH3.   The influence of varying the incident shock wave Mach number on the transmitted pressure profile was better represented by the areal impulse (Fig. 23b)

Effect of Boundary Conditions
In order to understand the effect of boundary conditions on the reflected shock pressure history, Fig. 24a  It can be seen in Fig. 24 that the imparted impulse for the freestanding plate was significantly less than that for the corresponding fixed plate. For a free-standing plate, the impact of a blast wave caused the plate to recede. The receding motion of the plate relieved the pressure experienced by the plate and resulted in a decrease in the impulse transmitted to the plate.

Conclusions
In the first part of this study, shock wave loading experiments on freestanding perforated plates was conducted to evaluate the reflected pressure pulses and the motion response for all the targets with and without perforations. The second part of this study was on clamping all the targets inside the shock-tube to evaluate the reflected and transmitted pressure pulses for all the targets. Through these experiments, the following major conclusions have been drawn:

Introduction
Blast loading events that arise from the detonation of explosives pose a severe threat to the lives of civilians and military personnel alike [1,2]. Hence, there has been extensive research on how to mitigate blast wave effects. There currently exists a variety of blast and shock-proofing methods that are implemented in a variety of applications including passive mitigation systems. These mitigation techniques leverage four approaches towards blast mitigation, including impedance mismatching, sacrificial cladding, blast deflection as well as blast and shockwave disruption [3]. The employment of plastically deforming metal plates and sandwich structures to absorb energy for impact and blast loading applications are reported widely in literature [4][5][6][7][8].
Extensive studies have been conducted on various core configurations from closed cell, lattice corrugation, pyramidal truss, honeycomb and composite functional graded structures [9][10][11][12][13][14][15]. However, these methods often involve complex and costly materials and/or fabrication methods. Thus, it is the aim of this work to investigate the blast mitigation effectiveness of a relatively simple and cost-effective treatment: the application of surface roughness to the face of a normally loaded material.
Prior literature has explored the problem of shock wave interaction on the boundary layer over a rough wall in transonic circular bump geometry. However, shock being normally loaded or impinged onto a rough surface structure (as is typical with an explosive event) has not previously been investigated even though it is of importance in any treatment of shock wave attenuation. According to A. Kumar et al and J. Mendonca et al, shock wave interaction on the boundary layer causes many noticeable features on the local as well as the whole flow field such as flow separation and reattachment, boundary layer thickening, shock unsteadiness etc. [16,17]. When the surface roughness is considered as a flow parameter, an added complexity in the physics of transonic flow is integrated [17]. Understanding the fluid structure interaction behavior during blast loading plays a major role in blast mitigation and helps in evaluating the blast performance of structures and consequently benefits in the design of higher blast resistance structures [18].
The blast wave propagation and reflection from structures with various shapes has been widely studied . A 2-D numerical model simulating interactions between a blast wave and a V-shaped or a cone-shaped structure was developed by W.
Peng et al. [22]. Normal, oblique, and Mach stem reflections are types of reflection when a blast wave impinges on a surface. The Normal reflection occurs when the incident angle is 0°. The angle between the incident shock front and the reflecting surface of a structure is defined as the incident angle [22]. The authors reported that when the incident angle is greater than 40°, Mach stem reflection occurred. The Mach stem reflection resulted in noteworthy decrease of reflected pressure, as well as the transmitted impulse. The influence of Mach stem reflection in reducing blast wave impact decreased with the decrease of blast intensity [22].
To expand on the aforementioned work already conducted in this field which investigates only the effect of surface roughness on transverse shock loading, specimens made of 1018 mild/low carbon steel were manufactured with a variety of surface roughness values and subjected to normal blast loading. Using a shock tube to initiate the shock and DIC techniques to analyze full field deformation, it was determined that blast mitigation improves as surface roughness increases, insofar as damage to the structure itself is concerned. That is to say, specimens with higher surface roughness values deformed less than specimens with lower surface roughness values when subject to the same loading. Furthermore, pressure sensors were used to obtain the pressure profiles as a function of time at various distances from the specimen. It was found that the reflected peak pressure measured 23 mm from the specimen decreased as surface roughness increased. Peak pressures measured further away from the specimen (183 mm) were determined to be unaffected as surface roughness varied from specimen to specimen.

Experimental Setup and Procedures
Shock tube experiments were performed to establish the dynamic behavior of 1018 mild/low carbon steel panels at different surface roughness of (0.8, 1.4, and 5.0) µm. At least three experiments were performed for each case to ensure repeatability.

Specimen Geometry and Preparation
The material used in this experimental study was 1018 mild/low carbon steel,

Shock Tube
A shock tube apparatus was used to generate a concentrated in air shockwave that is imparted on the panels (Fig. 2a). The total length of the shock tube is 8 m and is composed of three separate sections: driver section, driven section, and reduced diameter muzzle section. A Mylar diaphragm separates the driver and driven sections, while the driver section is pressurized using helium gas. Under critical pressure the diaphragm bursts, releasing a high pressure wave. The high pressure wave travels down the length of the driven section and develops into a shock wave front. The shockwave then travels through the muzzle section and the pressure of the event is captured by two dynamic pressure transducers (Fig. 2b). The pressure transducers capture the magnitude of the pressure while velocity of the wave can be inferred from the time between pressure histories and established positions of the sensors. The shockwave then leaves the muzzle, impacts the specimen and the pressure from the impact is reflected back into the muzzle. The reflected pressure is the pressure that loads the specimen. A detailed dimension of the muzzle section is shown by Fig. 2b. contact is only on the edge of the back face. The specimen was then secured to the simple support using Nickel-Chrome wire with a diameter of 0.05 mm, which easily breaks during shock loading and deformation.

High Speed Photography Systems
High speed photography coupled with 3D Digital Image Correlation (DIC) was utilized to provide full field displacements, strain, and velocities of shock tube specimens during in air shock loading. Two high speed cameras were setup to view the back face of the specimen for 3D DIC while a side view camera was positioned to view the out of plane deflection as shown by Fig. 3. The cameras used during experimentation were Photron FastCam SA1. Images of the event from all cameras were taken at 50,000 fps for an inter-frame time of 20 μs.

Experimental Results and Discussion
The experimental data obtained were carefully analyzed to obtain the impulse imparted to the specimen, back-face deflections, in-plane strains, and the energy used in deforming the specimens.

Pressure Profile and Impulse
The pressure history for each experiment is shown in Fig 4. Recall that channel 1 is the closest to the specimen (23 mm away) and channel 2 is the furthest from the specimen (183 mm away) (see Fig 2b). For each experiment, the incident shock wave For the same incident peak pressure, reflected peak pressures were not the same for different plates' surface roughness (see Fig. 4a).  As shown in Fig. 5, the impulse imparted to the specimens at higher surface roughness was significantly less than that at fine surface finish which resulted in lower peak deflections. The manner in which the shock wave was reflected by a surface is dependent on the microscopic shape characteristics of the surface. A smooth surface may reflect the shock wave in a single direction, while a rough surface will tend to scatter the shock wave in various directions. Consequently, for the same applied pressure, specimens with higher surface roughness deformed for a shorter time and reached lower deflections than the fine surface roughness specimen.

Digital Image Correlation Analysis
The deflection, in-plane strain (ε yy ), and the velocity of the back face for each experiment were generated using 3-D Digital Image Correlation (DIC) technique. To verify the performance of the DIC method, the back-face mid-point deflections obtained using the proposed DIC technique were compared with the deflections measured from the side-view images (Fig. 6). The comparison of the deflections obtained from the DIC and the side-view images for the fine and milled surface finish experiments is shown by Fig. 7. For both experiments, the deflections obtained from the DIC matched very well with the deflections measured from the side-view images (maximum difference at peak deflections is less than 2%). The results confirm the accuracy of the DIC technique. The surface roughness effect on the mean back-face deflection history is shown in Fig.   9. All of the curves exhibit a maximum elastic-plastic deformation in the first peak, followed by elastic reverberations and damping. For all experiments, the specimens deflected in a similar manner prior to reaching their maximum deflections and beginning to reverberate. As discussed earlier in Fig. 6a, the fine surface finish specimen (R a = 0.8 µm) showed a maximum deflection of 67.0 mm at t = 5400 μs and then began to reverberate. As the surface roughness increased to R a = 1.4 and 5.0 µm for the same incident pressure loading, the panel reached maximum back-face deflections of 60.0 mm and 49.0 mm at 5000 μs and 4000 μs and then began to reverberate respectively. It should be noticed that the specimen with higher surface finish began to reverberate ahead of time, 400 μs and 1400 μs (for grinded and milled surface cases respectively), as compared to fine surface experiment. At t = 8000 μs, the specimen which had a surface finish of R a = 5.0 µm settled with 43% less out-ofplane deflection as compared to the specimen of R a = 0.8 µm surface finish. The full field out-of-plane deflection (W) images for different surface finishes, with a scale of 0 mm (purple) to 67.0 mm (red), are shown in Fig. 10. It is important to note here that these deflections were obtained for much longer times than shown in The mid-point in-plane strains (ε yy ) on the back-face of the specimen are plotted in Fig. 11. It can be observed from the figure that all the specimens exhibited similar in-plane strain profile of about 0.27% up to t = 400 μs. After t = 400 μs, the specimens showed significant bending, which resulted in higher in-plane strain values.
At surface roughness of R a = 0.8 µm, a maximum in-plane strain of 7.0% was observed at t = 5400 μs. As the surface roughness increased, the specimens exhibited lower deflections (less bending, see Fig. 6), causing a decrease in the in-plane strain values. For R a = 1.4 µm and R a = 5.0 µm experiments, the specimens showed maximum in-plane strain of 5.5% and 4.2% at 5000 μs and 4000 μs, respectively. In comparison to the fine surface roughness experiment, the specimens with R a = 1.4 µm and R a = 5.0 µm showed a decrease in ε yy of 22% and 40%, respectively.

Deformation Energy Evaluation
According to Wang and Shukla [24], the total work done by the gas to deform the plate is defined as the deformation energy. The deformation energy was calculated by evaluating the deflection-time data from the side-view images (Fig. 6) and the force-time data (Fig. 5) from the reflected pressure profile. Combining the deflection-time data and the force-time data resulted in force-deflection data. The deformation energy could then be obtained by integrating the force-deflection data using Equation The deformation energy for different plate's surface roughness is plotted in Fig. 13. The maximum deformation energy of 34 J was observed for the plate with milled surface finish. The plates, which have grinded and fine surface finish showed maximum deformation energies of 40 J and 45 J, respectively. The deformation energy for the plate with R a = 5.0 µm was approximately 25% lower than that for the plate with R a = 0.8 µm. This indicates that the plates with higher surface finishes absorbed significantly less energy during the shock loading process. As stated in the previous section, the impulse was higher for lower plate's surface roughness. This allows for greater deflections for a given shock load, resulting in higher work done on the specimen during fluid structure interaction.

Conclusions
Based on results from shock loading experiments of 1018 mild low carbon steel panels with specific machined surface roughness, the following conclusions were drawn: • In order to decrease the reflected pressure, thus reducing the transmitted impulse and improving the blast wave mitigation, the surface roughness of structures subjected to blast waves should be considered.
• The impulse imparted to the plate decreased by 15% as the surface roughness (R a ) increased from 0.8 µm to 5.0 µm.
• The effect of different plate's surface roughness on the reflected peak pressure is highly localized, and that effect vanished at greater distances upstream.
• The maximum in-plane strain showed a decrease of 40% (from 7.0% to 4.2%) as the plate's surface roughness (Ra) increased from 0.8 µm to 5.0 µm.

CHAPTER 6: RECOMMENDATIONS AND FUTURE WORKS
Upon completing this dissertation, research gaps have been identified and could be addressed during future works. Based on the results from the shock loading and mitigation research studies, the following recommendations are made:

Temperature Loadings
• For a strong blast performance of coated laminated glass panels, both the interlayer and the coating film should have high ductility and strength at the target temperatures. Different coating materials designed for more extreme temperature ranges (below -10 °C and higher that 110 °C) should be considered.

Novel Impact Mitigation Technique using Shock Loading
• The impactor design and independent trigging mechanism should be improved so the capacity of a bladder as well as the trigger mechanism increases in mitigation efficiency.
• An optimization analysis should be performed to find the ideal needle length and shock release time so the energy mitigation is maximized.
• This device could be investigated towards reducing the effects from shrapnel.

Blast Performance of Perforated Structures
• Quantitative 3D-density measurements of the transmitted shock waves could be evaluated using a Background Oriented Schlieren (BOS) technique with high-speed photography.

The Influence of Surface Roughness on Blast Mitigation
• An increased range of surface roughness could be investigated to the (R a ) profiles of each texture and to obtain the optimal roughness and reduce the imparted impulse.
• A finite element model could validate experimental results and be used to investigate the fluid structure interaction and how the surface roughness changes this interface.
• A post-mortem study could be undertaken using the Scanning Electron Microscope (SEM) imaging to look at any surface roughness alteration.  Are all the screw connections TIGHT?
 Have you put the PRESSURE SENSORS in the shock tube?
 Is the shock tube PERPENDICULAR to the supports, and centrally located?
Dump Tank Related:  Is the specimen in the right position and secure?
 Is the shock tube close enough to the specimen, 1/16 in?
 Have you put the side doors on the dump tank?
 If no back side DIC system is required, have you installed a good protection for the back side LEXAN windows ?
DIC System Related:  Have you double checked the camera CABLE CONNECTIONS?
 Is the back view camera system PERPENDICULAR to the back side window?
 Is the side view camera PERPENDICULAR to the side window and specimen?
 Have you set the right FRAME RATE?
 Have you done the BLACK CALIBRATION (SHADING)?
 Is the viewing area acceptable (make the view of the specimen as large as it can)?
 Have you FOCUSED on the specimen (use iris 2.8 or smaller)?
 Have you increased the IRIS to at least 5.6?
 Have you set SYNC MODE for all slave cameras (E-SYNC)?
 Have you done the correct CALIBRATION (grids used to track particle distance)?
 Have you set the right TRIGGER MODE?
 Have you double checked that the cameras can take the desired images?

ENERGY ANALYSIS
This manual is designed for the steps to use the energy and impulse analysis code.
(1) Obtain the original data The original data of shock tube experiments are from the Tektronix oscilloscope (TDS3014 or 3014C). The data must have following name: First channel: TEK00000.csv Second channel: TEK00001.csv Normally, there are two columns in these files. The unit of the first column is second (s). The unit of the second column is voltage (v).Please copy these files into the folder named "experimental data backup".
(2) Analyze the original data. This step is to analyze the original data to obtain the shock wave velocity, the peak pressure and the modified pressure profiles.
This step is carried out in the folder named "original data analysis". In this folder, the m file named profile_analysis.m is necessary. Other files can be deleted or replaced.
You must copy the original data files into this folder and then run the code. The running process is as follow, (1) The code will first ask you how many plys you use in the experiment. This information is only for your record. It does not matter the analysis process.
(2) The code will ask you the sensitivity of the sensors. This value is given in the box of the sensors. This value means how many milli-voltages related to 1 psi.
(3) Then, a figure with two plots of the pressure profiles will be given. Look at (4) Then the code will inform you as follow, "The peak and velocity data have been saved into the file, which isnamed peak&velocity.txt and in the same folder of this code.
There are two more pressure data files in this folder: inc_sp.dat ref_sp.dat They can be used for energy and impulse evaluation.
The code will give some plots to verify your data.
Please double check them very carefully.
press any key to continue" (5) After pressing any key, the code will give four images: The first two files will be used to analyze the energy and impulse. There are two column data in these two files. The first column is time with unit second (s). The second column is pressure data with unit psi.
The last file records the physical parameters, which needs to be input in the energy and impulse analysis.
(7) Please cut these three data files into the folder named "experimental data backup" and delete all of these files in the current folder.
(3) Analyze the incident and remaining energy and impulse This step is to use the data obtained in step 2 to analyze the incident and remaining energy and impulse in a shock tube experiment.
This step is carried out in the folder named "gas energy and impulse analysis". In this folder, seven m files are necessary. They are, Other files can be deleted or replaced. You must copy the data files: inc_sp.dat and ref_sp.dat, into this folder and the correlated blank test data file, inc.dat, from the folder named "blank test data". The code running process is as follow, (1) The code will first show the format and unit of the data. Please be sure that the data should be the exact format.
(2) Then the code will give the total number of the data and ask you how many Skip two points (3) Then the code will ask you to input physical parameters obtained in step 2 (saved in peak&velocity.txt).
(4) The code will calculate the physical parameter profiles. This process is automatic.
(5) Then the code will integrate the parameter profiles to obtain the energy components. This process is automatic.
(6) Finally, the code will ask you to input the name of the file which you want to save data into. Then all of the data will be saved into Please copy and save these data into a safe folder.
(4) Analyze the deformation energy of the gas, momentum and kinetic energy of the specimen This step is to use the data obtained in step 2 and the high-speed side-view images to analyze the deformation energy of the gas, momentum and kinetic energy of the specimen in a shock tube experiment.
This step is carried out in the folder named "specimen momentum and energy". In this folder, only one m files, Deformation_momentum_kinetic_photron.m, are necessary.
Other files can be deleted or replaced. You must copy the data files: ref_sp.dat, and a series of high-speed images into this folder. The code running process is as follow, (1) The code will first show the format and unit of the data. Please be sure that the data should be the exact format.
(2) load the time series of the images You will have three ways to load the time series of the images.
(i) The time between two frames is same. You can input total number of frames and time between two frames. Then the code will generate the time series automatically.
(ii) The time between two frames is not same. You can input total number of frames and input time between two frames frame by frame.
(iii) The time between two frames is not same. The time between two frames is not same. Then you can just load that time series data file.
You can choose anyone and following the instruction.
(3) Length calibration. You can choose any image for length calibration. You will need to choose two points on this image and the vertical distance between these two points will be used to calibrate the length. Therefore, you need to know one real vertical scale between two points on the image.
For example: (i) the span of the supports is 6 inches (ii) the outer diameter of the shock tube is 5 inches The process will repeat three times. Thus, totally you will pick six times.
Please follow the instruction. (5) The code will ask you to input the mass of the specimen.
(6) The code will ask you to input the name of the file which you want to save you data into.
17. Once sample has been blasted, shut off gas and proceed to checking the scope and camera for data 18.