Disclaimer for FAA Research Publication

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1 Disclaimer for FAA Research Publication Although the FAA has sponsored this project, it neither endorses nor rejects the findings of the research. The presentation of this information is in the interest of invoking technical community comment on the results and conclusions of the research. H.Kim et al., UCSD 1

2 Impact Damage Formation on Composite Aircraft Structures Hyonny Kim, Gabriella DeFrancisci, Daniel Whisler, Jennifer Rhymer Department of Structural Engineering, University of California San Diego La Jolla, CA Project Description Paper Supporting Presentation Given at Federal Aviation Administration Joint Advanced Materials & Structures (JAMS) 5th Annual Technical Review Meeting July 2009, NIAR/WSU, Wichita, KS Abstract The ongoing FAA research activities at UCSD, summarized herein, are composed of: (i) large-scale blunt impact, (ii) lab-scale blunt impact, and (iii) hail ice impact. The blunt impact studies are focused on understanding the development of the formation of massive internal damage to composite fuselage, when contact is made by ground vehicles/equipment, with little or no external visible detectability. The hail ice impact work seeks to establish a database for the formation of damage by high velocity ice impacts, and to establish models for predicting damage initiation failure thresholds as well the final state of damage produced. H.Kim et al., UCSD 2

3 1.0 Introduction Impact damage resulting from collisions of ground vehicles/equipment with aircraft structural components, as well as from events such as hail and bird strikes, is a significant source of damage to commercial aircraft that has the potential to go by undetected. Impacts by hail and birds can occur at in-flight velocities, thereby posing significant threats to the structure. More commonly occurring, however, are blunt impact threats such as ground maintenance and service vehicles, equipment, etc., as shown in Figure 1. With new all-composite fuselage transport aircraft coming into service, significantly more composite skin surface area is exposed to such impacts. To address the difficulties that exist in being able to predict and detect the damage resulting from blunt impact, and to aid in assessing its effect on structural performance, focused investigation on the development of impact damage is needed. Of particular interest is damage that can be difficult to visually detect from the exterior, but could be extensive below the skin s outer surface. Sub-surface damage (typically delamination and backside fiber failure) usually forms in a panel skin when impacts occur at levels just exceeding the amount needed to initiate failure (Kim et al. [1], Kim and Kedward [2]). This level is referred to as the failure threshold energy. Additionally, damage from blunt impacts to internal stiffeners can be extensive, as well as debonding between the stiffener and the skin. Of critical concern is whether such extensive damage can result in the structure losing limit load capability. H.Kim et al., UCSD 3

4 Figure 1. Maintenance/Service Threat Sources: Ground Vehicles, Luggage Carts, Cargo Containers, etc. The objectives of this research project focuses on impact damage formation by a range of sources, including: (i) low velocity wide-area blunt impact vehicle/ground maintenance collision, and (ii) high velocity hail, bird, and general impact: Low-Velocity High-Mass Wide-Area Blunt Impact: 1. Identify which blunt impact scenarios are commonly occurring and are of major concern to airline maintenance organizations and aircraft manufacturers. H.Kim et al., UCSD 4

5 2. Develop Methodology for Blunt Impact Threat Characterization and Prediction. 3. Experimental identification of key phenomena and parameters governing high energy blunt impact damage formation, particularly focusing on what conditions relate to the development of massive damage occurring with minimal or no visual detectability on the impact side. 4. Damage tolerance assessment of blunt impact damaged structures with focus on conditions related to loss of limit load capability for level of damage incurred, and which types of structural configurations and details are more prone to this loss of capability. High Velocity Hail, Bird, and General Impact: 1. Investigate impact damage initiation and damage formation to composite panels, including those of skin-stiffened and sandwich construction. 2. Develop models capable of predicting impact damage to composite panels. 3. Develop unified treatment methodology for predicting damage initiation by variety of impactor projectile types e.g., bird, hail, tire fragment, runway debris, lost access panel, etc. Accomplishment of these objectives are intended to aid maintenance engineers in assessing whether an incident could have caused damage to a structure, and if so, what sort of inspection technique should be applied to resolve the extent of damage. Furthermore, it is expected that design engineers can make use of the research outcomes to: (i) improve the resistance of composite aircraft structures H.Kim et al., UCSD 5

6 to damage from blunt impacts as well as a variety of other sources such as hailand bird-strikes, runway debris, lost access panel, etc, and (ii) provide critical information on the mode and extent of seeded damage, particularly those not easily detected by visual inspection, resulting from a wide gamut of impact threats i.e., low to high velocity. 2.0 Project Results to Date The results of the three project activities are described in separate subsections. (i) large-scale blunt impact, (ii) lab-scale blunt impact, and (iii) hail ice impact. 2.1 Large-Scale Blunt Impact Test Specimens The large-scale blunt impact experimental activities planned at UCSD are to be conducted over multiple years, as described by the building block pyramid shown in Figure 2. The First year of activity will focus on establishing a basic understanding of key failure modes, how these are excited in relationship to bluntness parameters and incidence angle of the impact, and the establishment of a clean database measuring structural response and failure development. In addition to assessing the mechanisms of how blunt impact damage forms, these data will be critical to the development of modeling methodology and simulation tools for predicting damage. H.Kim et al., UCSD 6

7 Phase III (Year 3) Phase II (Year 2) Phase I (Year 1) Increasing Length Scale, Complexity, and Specificity OEM Hardware - 1/4 to 1/2 Barrel Size - Vehicle Impacts Large Panel - e.g., 5 Bays - Damage Excitation - Damage Thresholds - Model Correlation Basic Elements - Excite Key Failure Modes - Model Correlation Data - Understand Damage Formation & Relationship to Bluntness Parameters Modeling Capability Development & Correlation with Test are Key Aspects at Each Level Scaling, B.C. Effects Scaling, B.C. Effects Dynamics Figure 2. Blunt Impact Testing Building Block During an on-site workshop held on January 23, 2009 at UCSD, participants (approx. 40 persons from industry, agency, and academia) agreed that two configurations of full-scale element-level test specimens be defined for the Basic Elements tests to be conducted as part of Phase I. The first configuration, shown in Figure 3, is primarily focused on damage development to the circumferential frame members and their connection to the skins. The second configuration, shown in Figure 4, is focused on damage formation to the stringers and their connection to the skins. While a recent Working Meeting on-site at UCSD held on June 30 and July 1, 2009, has modified the configuration of the test specimens (to be wider in stringer direction), the drawings shown in these figures remains conceptually correct. H.Kim et al., UCSD 7

8 Frames Shear Ties Skin Stringers Figure 3. Frame Focused Test Specimen Figure 4. Stringer Focused Test Specimen These specimens will be tested in the UCSD Powell Structural Research Labs which are well-suited for conducting large-scale tests. The first year of tests will H.Kim et al., UCSD 8

9 consist of quasi-static representative impacts, which have been shown to be equivalent for low velocities of impact (Wardle and Lagace [3]). The test setup is shown conceptually in Figure 5. The test specimens are presently being designed and will be fabricated by UCSD in conjunction with UCSD s industrial research partner San Diego Composites. Specimen materials will be carbon fiber and toughened epoxy matrix (reflecting current aerospace fuselage materials) provided by UCSD s industrial partner Cytec. Present status of specimens. At present, the test specimen design has been finalized and tooling for the curved skin and c-shaped frame are being fabricated. Tools for the shear tie elements connecting the skin to the frame, and for the stringers will soon be fabricated. Development of details and manufacturing trials are planned prior to fabricating a full-sized part. Figure 5. Conceptual Setup in UCSD Large Scale Test Labs H.Kim et al., UCSD 9

10 2.1.2 Boundary Conditions. Boundary conditions are a critical aspect of this project. Of primary importance is that the small sized test specimen to be tested at UCSD has the same stress state and key deformation metrics as a full-sized barrel being impacted by the same conditions. This can be achieved by specifying appropriate boundary conditions on all four sides of the test specimen. A methodology, using FEA models of full-sized barrels and of the test specimens, is shown in Figure 6. The full-barrel model will be analyzed with blunt impact conditions applied, from which key stress and deformation quantities are determined. In parallel, a modified fullbarrel model having a cutout zone representing the test specimen will be interrogated with relevant edge loadings and moments. The rotational and translational stiffness can be determined on all four sides in this manner. These stiffnesses, referred to as k i in Figure 6, will be applied to the small panel specimen model boundaries, and the stress and deformation quantities extracted from this model will be compared with those corresponding values from the fullbarrel model. Iteration of k i is likely needed until these quantities match (within acceptable tolerance). In this manner, the test specimen can be made to represent the stress and deformation state of the full-barrel. These boundary stiffnesses would be implemented by a set of high-stiffness coil springs as illustrated in Figure 5. H.Kim et al., UCSD 10

11 Figure 6. Methodology for FEA-Based Boundary Condition Determination It should be noted that for quasi-static testing, the transverse stiffness (in loading direction) is not needed. This concept is illustrated in Figure 7. For true dynamic testing, the mass of the entire aircraft, or a representative base mass M base, must also be accounted for. As noted earlier, this phase of test activity will conduct tests in quasi-static manner, keeping track of applied indentation displacement as a key metric describing the applied threat. Future phases will address dynamic effects and will involve dynamic testing. H.Kim et al., UCSD 11

12 Figure 7. Dynamic vs. Static Boundary Conditions Finite Element Results Finite element analysis (FEA) of the test specimens has been conducted to determine locations of high stress and to observe deformation states that can drive interlaminar/debonding failures of the test specimen. Two locations of interest were investigated on the Frame Panel specimens, as shown in Figure 8. At each of these locations, an impactor with radius of curvature 3.0 and 12.0 in. were applied under a displacement-control mode of loading. Figure 8. Frame Panel Impactor Locations H.Kim et al., UCSD 12

13 The bending-induced normal stresses (in frame-direction) are summarized in Figure 9 for an indentation depth of 0.7 in. for both impactor radii. The maximum compressive stress occurs on the panel skin on the impactor side, and the maximum tensile stress occurs in the c-shaped frame. Note that these models have linear-elastic material behavior with nonlinear geometry effects and surfaceto-surface contact interactions active. The relative values of stress are more meaningful/insightful than the actual values of these quantities. The smallerradius impactor produces significantly higher compressive stress than the 12 in. radius, largely due to the localized curvature that the smaller radius imposes onto the skin side. However, the tensile stress in the frame is roughly the same in both cases. Therefore, the larger radius impactor has less propensity to produce surface-visible damage (e.g., in-plane compressive failure of outer skin) than the smaller radius impactor. Figure 9. FEA of Indentation at Location 1 H.Kim et al., UCSD 13

14 Figure 10 summarizes a comparison of the normal stress in the shear ties (in direction perpendicular to skin). Tensile values near the skin surface act as driving forces for pull-off of the shear ties from the skin, as well as causing the development of interlaminar tension failure in the radius region of the shear tie where it curves and is mechanically fastened to the skin. Note that the set of shear ties located away from the impactor location also have high tension stress, indicating the likelihood of damage development (pull-off) at locations away from the impactor location. Figure 10. Normal Stress in Shear Ties Driving Force for Pull-Off from Skin and Interlaminar Tension Failure in Shear Tie Radius Figure 11 shows the indentation of the test specimen at location 2 (see Figure 8) which is located directly over a stringer. The impactor first contacts the skin spanning between the stringer walls. The main difference between the two radii conditions is that the 12 in. impactor makes contact with the stringer walls, whereas the 3 in. impactor does not for an indentation depth of 0.7 in. The effect H.Kim et al., UCSD 14

15 of this is more load transfer from the 12 in. radius impactor into the frame, thereby resulting in significantly higher bending stress in the frame flange with still-lower skin surface compressive stress than the 3 in. impactor case. Therefore, for this condition, the larger-radius impactor has greater likelihood for producing internal damage with less driving force for producing surface-visible damage (e.g., due to compressive failure in skin). Figure 11. FEA of Indentation at Location 2 Out-of-plane stresses are not predicted by shell elements in FEA, which were the element types used in these models. The side-view of Location 1 loading, shown in Figure 12, illustrates deformation states in the skin and stringers that would drive the debonding/delamination for the stringers from the skins. This shows skin-stringer debonding to be a key damage mode of interest, with two stringers being debonded for shear-tie located impacts (i.e., at location 1). H.Kim et al., UCSD 15

16 Figure 12. Outward Bulging of Skin Between Stringers Likely Driving Force for Stringer Debonding The contact force developed for an indentation depth of 0.7 in. at the locations 1 and 2 of the frame panel provide insight into the differences between smaller vs. larger radius impactors. These forces are summarized in Table 1. For both locations, the smaller-radius impactor develops a lower contact force than the larger-radius impactor. This is logical in the context of the degree of local deformation developed. The contact force at location 2, however, is almost 2X higher for the 12 in. impactor due to the fact that the larger-radius impactor develops contacts with stiffer internal components more earlier than the smaller radius impactor (if these contacts even can develop at all). A few points can be draw from this observation: (i) the larger radius impactor develops more global stress state by involving more of the surrounding structure in the vicinity of the impactor, (ii) the forces developed can be much higher for given amount of indentation depth which could influence development of damage both at the impact location as well as at internal reaction points further away, (iii) the larger radius impactor develops a more spread out contact zone on the exterior surface which can be related to lower likelihood for leaving visible markings/damage due to lower contacting pressures and lower compressive bending stresses developed. H.Kim et al., UCSD 16

17 Table 1. Contact Force for 0.7 in. Indentation at Locations 1 and 2 Impactor Radius 3 inches 12 inches Location lbs 4750 lbs Location lbs 3250 lbs Test Plan The test specimens will be loaded incrementally as illustrated in Figure 13, with the intention of gathering elastic-response (i.e., no damage) data for various impactor conditions, as well observing the growth of various damage modes as increasing level of indentation displacement is applied. As illustrated in Figure 13, the test specimen will be unloaded following detection of initial failure, after some intermediate level of damage development, and after severe level of load drop indicated massive/final failure. Determination of the damage state by visual and nondestructive (e.g., portable in-situ c-scan) methods will catalogue the damage state at each level. These data are particularly important for subsequent modeling activity focused on prediction of damage initiation and growth. H.Kim et al., UCSD 17

18 Figure 13. Test Specimen Incremental Loading Future Activity and Expected Outcome In the immediate near term, the stringer panel test specimens will be fabricated in August 2009, and tested in September-October Frame panels will be fabricated in the fall and tested in early These compose the basic tests of the first phase of the activity. Subsequent phases of activity will involve larger, more specific tests specimens. Expected outputs of these studies will be: (i) experimental based description of blunt impact damage formation mechanisms, (ii) database on the structural response of large test panels to blunt impacts (iii) modeling capability for predicting blunt impact response, blunt impact damage initiation and modes, and H.Kim et al., UCSD 18

19 extent of blunt impact damage, (iv) methodology for conducting blunt impact tests and analyses, and (v) guidance on the scaling of blunt impact test results to larger-scale/full aircraft, including issues of proper boundary condition representation. 2.2 Lab-Scale Blunt Impact Overview The effects of bluntness of an impactor are of interest as this is related to both the external visual detectability of an impact event, as well as the development of any internal damage in the laminate. The objective of this investigation is to determine the effect of impactor radius on the initiation of damage to composite panels. A low-velocity pendulum impactor was used to strike 200 mm square woven glass/epoxy composite plates of 3.18 and 6.35 mm thickness. Hemispherical steel impactor tips of radius 12.7mm to mm were mounted to a piezoelectric force sensor which measures the contacting force history during the impact event. Distinct threshold energy levels for the onset of delamination and backside fiber breakage have been measured. These threshold energy levels increase significantly with increasing impactor tip radius Experimental Setup The pendulum impactor is shown in Figure 14. With a pendulum arm length of 1.402m and a total mass of 5.5kg, the impactor is capable of energy levels of up to 150J. Position control is achieved through an optical encoder with 0.1º resolution attached to the pivot of the arm. A steel reinforced test fixture provides H.Kim et al., UCSD 19

20 the mount for two 12.7mm thick aluminum picture frame fixture with a 165 x 165 mm window (see Figure 14). Force measurements are achieved via a piezo-electric sensor with a 0-5V range output. Two different sensors are available: a 19mm diameter, 22.24kN max force Dytran model 1050V6 sensor and a 50.8mm diameter, 111.2kN max force Dytran model 1060V5. Both sensors accept custom-shaped tips. Three different spherical tips (see Figure ) with radius 12.7mm, 50.8mm, and 152.4mm are used in this study to investigate increasing levels of bluntness. The material tested is FR4 woven fiberglass/epoxy. Panels are cut to approximately 200 x 200 mm and are either 3.18mm or 6.35mm thick. The material properties are E=18.6GPa, ν=0.18, and ρ=1860kg/m Test Methodology The matrix provided in Error! Reference source not found. shows the six different test scenarios available for three impactor tips and two panel thicknesses. For each case, two different test protocols are used: "sweeping" and "bracketing". Sweeping is a single panel tested at multiple increasing energy levels until it fails. Bracketing tests a single panel at energy levels just below, above, and right at failure. To be clear, failure is defined in this research project based on two distinct mode of damage. Failure threshold energy 1 or FTE1 is the projectile kinetic energy to just initiate delamination, localized to the impact H.Kim et al., UCSD 20

21 site. For the FR4 glass/epoxy specimens, the delamination is observable as an internal white region. Failure threshold energy 2 or FTE2 is related to the onset of fiber breakage and is confirmed by the presence of cracks visible on the panel surface. Figure 14. Pendulum impactor test setup H.Kim et al., UCSD 21

22 Large Impact Tips R152.4 R101.6 R50.8 R25.4 Large Force Sensor Small Impact Tips R6.35 R12.7 Small Force Sensor Figure 15. Impactor tips and force sensors; tip radius dimensions in mm Table 2. Test Matrix Number of Panels Tested for each Thickness T, Impactor Tip Radius R R 12.7mm R 50.8mm R 152.4mm T 3.18mm T 6.35mm Each test run is conducted as follows. The fiberglass panel is clamped in place and centered in the test fixture. Simple energy based calculations convert the desired kinetic energy at impact to a specific angular position (i.e., potential energy level of raised impactor mass). The pendulum arm is then raised accordingly and a pneumatic locking mechanism secures it in place. Carbon and H.Kim et al., UCSD 22

23 graph paper sheets are layered in front of the panel directly in the path of the impactor. At release, air pressure opens the pneumatic mechanism and the impactor swings to target. The tip strikes the graph paper, carbon paper, and the plate. The impact on the graph paper records the size of the contact area, the sensor just behind the tip records the force, and the optical encoder records the position of the impactor. Velocity just before impact is calculated by finding the slope of the position data through a linear regression curve fit of the encoder absolute position data. This angular velocity is converted into an incoming velocity of the impactor just prior to impact Results A representative force time history plot, position time history plot, and contact area capture is shown in Figures 16, 17, and 18, respectively. Both force and arc position time history plots are actually an overlay of three separate tests. Given an identical mass, tip, energy, and target, the force and position measurements are observed to exhibit good repeatability. Oscillations in the force history data during impact are due to the panel vibrating upon contact and striking the sensor. The arc position data is intentionally offset from zero, and does not affect data processing since this signal is used only for determining the impactor velocity. Unfortunately, it does have a relatively high signal to noise ratio. Filtering the data with software is possible but unnecessary since the primary extracted parameter, velocity, was found to be the same when using either raw or filtered results. Contact areas are circular for all energy levels up to FTE2, as shown in H.Kim et al., UCSD 23

24 Figure 18a, with no observable changes at FTE1. At FTE2, the contact area changes from a circle to a "peanut" like shape (see Figure b). This progression is confirmed by Davies and Zhang [4], who attribute the interlaminar delaminations to high bending strains. The failure threshold energy is observed to increase with increasing panel thickness and tip radius. FTE2 is not reached by the 152.4mm impactor at energy levels of 40J. See Table 3. It is apparent that for a given energy level, two impactors differing only in their degree of bluntness may not initiate the same damage mode. It is possible then, that a structure with visible signs of damage may be identified and repaired accordingly, while another structure with no visible signs of damage may not be repaired, even though they were both subject to the same impact energy. Also, to initiate a given damage mode, a blunter impactor needs more energy than a sharper impactor. However, as the energy level increases, the entire test frame is increasingly involved in the target dynamic response, thereby affecting the build up of local contact forces on the panel. At higher energy levels, the frame experiences a substantial amount of deflection as it attempts to react the impactor energy. It is possible then, for two structures with visible signs of damage that one may have only local damage in immediate area whereas the other may have global damage in locations away from the impact. H.Kim et al., UCSD 24

25 3500 Contact Force vs Time Force (N) Time (s) x 10-3 Figure 16. Force time history plot for radius 12.7mm, panel thickness 3.18mm, and energy 7.5J H.Kim et al., UCSD 25

26 3.2 Arc Position vs Time Arc Position (m) Time (s) Figure 17. Arc position time history plot for radius 12.7mm, panel thickness 3.18mm, and energy 7.5J (a) contact shape for impacts below FTE2 (b) shape for impacts above FTE2 Figure 18. Contact area for tip radius 50.8mm, panel thickness 6.35mm H.Kim et al., UCSD 26

27 Table 3. Failure Threshold Energy Summary FTE1 for each panel thickness T, impactor tip radius R R 12.7mm R 50.8mm R 152.4mm T 3.18mm 2.5J 5J 12J T 6.35mm 7J 8.5J 12J FTE2 for each panel thickness T, impactor tip radius R R 12.7mm R 50.8mm R 152.4mm T 3.18mm 8J 12J N/A T 6.35mm 19J 30J N/A The force time history plots corresponding to FTE1 (see Table II and Figure ) show the peak force (critical force at FTE1) to be increasing significantly with larger tip radius for the 3.18mm thick panel. FTE1 for the 6.35mm thick panel, however, shows a more gradually-increasing peak force with tip radius (see Figure ). This can in part be explained by the larger difference in FTE1 values for the thinner panel across the range of tip radii, than what is observed for the thicker panel. The thicker panel is therefore less affected by impactor geometry than the more compliant thinner panel. H.Kim et al., UCSD 27

28 Contact Force vs Time FTE1 for T 3.18mm R 12.7mm R 50.8mm R 152.4mm 5000 Force (N) Time (s) x 10-3 Figure 19. Force time history plots at FTE1 for all radii and panel thickness 3.18mm Contact Force vs Time FTE1 for T 6.35mm R 12.7mm R 50.8mm R 152.4mm 5000 Force (N) Time (s) x 10-3 Figure 20. Force time history plots at FTE1 for all radii and panel thickness 6.35mm H.Kim et al., UCSD 28

29 The contact area for both panel thicknesses appears to be roughly linearly proportional to impact energy. See Figures 21 and 22 for the 3.18 and 6.35 mm thick panels, respectively. The closed symbols indicate measurements corresponding to no damage being formed (i.e., lower energy levels below FTE1), while the large open symbols indicate measurements following damage initiation. As expected, the larger radius impactor tips develop higher contact areas. The contact area is dramatically higher for the thin panels and for a given tip radius. This observation can be attributed to the ability of the thin panel to locally deform and conform to the impactor tip geometry more easily than the thick panel. Contact force is found to vary proportionally, but not linearly, with energy. Since stiffness is a characteristic of the panel system, the peak contact forces for all radii for a given panel thickness fall on the same curve, as shown in Figures 23 and 24. The contact forces generated for impacts onto the thick panel is significantly larger than for the thin panel, due to the relatively higher transverse stiffness of the thick panel. Greater peak contact forces are developed, in general, for more rigid impact conditions. Dividing the force by contact area to obtain the average contact pressure and plotting this quantity against energy shows trends not readily visible in either the force or area plots examined separately. Plotted in Figures 25 and 26, the average contact pressure for a given radius is observed to increase until it peaks around FTE1, and then decrease thereafter. Contact pressure is substantially lower for the 152.4mm tip when compared with the other two radii. Considering H.Kim et al., UCSD 29

30 the very low contact pressure, it is highly likely that a blunter impactor will not induce visible surface dents, even at high energy levels. However, for a wide area impact, the global damage at locations away from the contact point may be even more substantial despite the local damage being minimal or nonexistent. Impactor position during the impact event is available through double integration of the force time history data and using the initial velocity and known impactor mass parameters accordingly. It should be noted that because the position is derived from the force data, the values shown by the plots are the total motion that the impactor which includes deformation from panel, support structure, and pendulum arm. In fact, the flexibility in the pendulum is sufficient to prevent using impactor position data directly from the absolute encoder. Thus, without displacement time history plots of impacted components, extracting additional data, such as panel stiffness and midpoint panel deformation, is not possible. Limitations aside, the force versus impactor position plots shown in Figures 27 and 28 for the 3.18mm and 6.35mm thick panels, respectively, are insightful. These plots correspond to FTE1. The overlapping curves for a given panel thickness shows a non-linear force versus displacement response that is relatively independent of impactor bluntness. The overlapping loading and unloading paths indicate that little or no damage has accumulated as a result of these impacts. H.Kim et al., UCSD 30

31 Contact Area (mm 2 ) R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE1+ Contact Area vs Energy T 3.18mm Energy (J) Figure 21. Contact area as a function of energy for all radii and panel thickness 3.18mm Contact Area (mm 2 ) R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE1+ Contact Area vs Energy T 6.35mm Energy (J) Figure 22. Contact area as a function of energy for all radii and panel thickness 6.35mm H.Kim et al., UCSD 31

32 R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE1+ Contact Force vs Energy T 3.18mm Force (N) Energy (J) Figure 23. Contact force as a function of energy for all radii and panel thickness 3.18mm Force (N) R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE1+ Contact Force vs Energy T 6.35mm Energy (J) Figure 24. Contact force as a function of energy for all radii and panel thickness 6.35mm H.Kim et al., UCSD 32

33 Average Pressure (MPa) Average Contact Pressure vs Energy T 3.18mm R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE Energy (J) Figure 25. Average contact pressure as a function of energy for all radii and panel thickness 3.18mm Average Pressure (MPa) Average Contact Pressure vs Energy T 6.35mm R 12.7mm No Dam R 12.7mm FTE1+ R 50.8mm No Dam R 50.8mm FTE1+ R 152.4mm No Dam R 152.4mm FTE Energy (J) Figure 26. Average contact pressure as a function of energy for all radii and panel thickness 6.35mm H.Kim et al., UCSD 33

34 R 12.7mm R 50.8mm R 152.4mm Contact Force vs Impactor Position FTE1 for T 3.18mm Force (N) Impactor Position (mm) Figure 27. Contact force as a function of impactor position for panel thickness 3.18mm R 12.7mm R 50.8mm R 152.4mm Contact Force vs Impactor Position FTE1 for T 6.35mm 5000 Force (N) Impactor Position (mm) Figure 28. Contact force as a function of impactor position for panel thickness 6.35mm H.Kim et al., UCSD 34

35 2.2.5 Conclusions Laboratory testing of composite panels highlight several trends in impact damage formation with low velocity, blunt impactors. To create damage, a blunted impactor requires significantly more energy than a sharper impactor. Thicker panels are less affected by bluntness than thinner panels. A method for measuring contact area is described and has been found to show an approximately linearly relationship between contact area and energy. The contact area is much higher for thinner panels which are able to locally deform and therefore develop more contact with the impactor tip. Contact area measurements allow the determination of an average contact pressure. As expected, the peak average contact pressure is measured to be significantly lower for a blunted impactor. When plotted against impact energy, the contact pressure reveals an inflection point corresponding to the onset of FTE1. This implies a softening of the contact interaction between the impactor and target panel as delamination damage is formed. The inflection point, which is not readily visible in either force or contact area plots, can be used for identifying the onset of FTE1 in the smaller radius impact tips. For the largest radius tip (152.4 mm) no inflection in the average contact pressure versus energy relationship was observed. 2.3 Hail Ice Impact UCSD and Sandia Labs (point of contact: Dennis Roach) are collaborating on a project focused on development of damage to carbon/epoxy composite panels of H.Kim et al., UCSD 35

36 unidirectional tape construction. Of interest is particularly the initiation of damage, thereby identification of failure threshold energies, and the visibility and detectability (both exterior visible as well as by advanced NDI) of the damage produced by ice. Table 4 below outlines the test matrix to be investigated, which includes conditions for impact by both high velocity hail ice and low velocity pendulum (instrumented) impactors. The material being used in this study is the Toray T800/ toughened resin system composite being used in construction of the Boeing 787 fuselage. Panel specimen fabrication has begun and initial tests are presently under way at UCSD. Tested panels will be sent to Sandia for inspection. Select panels will remain with Sandia for use in future studies involving the detection of real impactinduced damage by various NDI methods. Panel Thickness Quasi-Isotropic Layup Table 4. Ice Impact Test Matrix Hail Dia 1 12x12 Number of Panels Needed for Each Condition * Hail Low Low Hail Hail Dia 2 Veloc Veloc Dia 2 Dia 3 (angle) Dia 1 Dia 2 12x12 12x12 Low Veloc Dia 3 12x12 6x10 6x10 6x10 8 plies [0/45/90/-45]_s plies [0/45/90/-45]_2s plies [0/45/90/-45]_3s Total * Matrix set-up: 3 specimens for structural/ndi testing; 3 specimens for trial impact calibration tests; 3 specimens retained for NDI use H.Kim et al., UCSD 36

37 3.0 References 1. Kim, H. and Kedward, K. T., Modeling Hail Ice Impacts and Predicting Impact Damage Initiation in Composite Structures, AIAA Journal, Vol. 38, No. 7, 2000, pp Kim, H., Kedward, K.T., and Welch, D.A., Experimental Investigation of High Velocity Ice Impacts on Woven Carbon/Epoxy Composite Panels, Composites Part A, Vol. 34, No. 1, 2003, pp Wardle, B.L., Lagace, P.A., On the use of Quasi-Static Testing to Assess Impact Damage Resistance of Composite Shell Structures, Mechanics of Composite Materials and Structures, Vol.5, No. 1, 1998, pp Davies, G. A. O. and X. Zhang "Impact damage prediction in carbon composite structures," Int. J. Impact Engineering. 16(1): H.Kim et al., UCSD 37