Impact Damage Formation on Composite Aircraft Structures

H.Kim et al., UCSD 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.


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 92093-0085

Project Description Paper Supporting Presentation Given at Federal Aviation Administration Joint Advanced Materials & Structures (JAMS)

5th Annual Technical Review Meeting 21-22 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.


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.


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.


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


to damage from blunt impacts as well as a variety of other sources such as hail-

and 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

2.1.1 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.


Basic Elements- Excite Key Failure Modes- Model Correlation Data

- Understand Damage Formation &Relationship to Bluntness Parameters

Large Panel- e.g., 5 Bays

- Damage Excitation- Damage Thresholds- Model Correlation

OEMHardware- 1/4 to 1/2Barrel Size

- Vehicle Impacts

Scaling,B.C. EffectsDynamics

Scaling,B.C. Effects

Increasing LengthScale, Complexity,and Specificity

Phase III(Year 3)

Phase II(Year 2)

Phase I(Year 1)

Modeling CapabilityDevelopment & Correlationwith Test are Key Aspects

at Each Level

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.


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

Skin

Stringers

Frames

Shear Ties


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


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 full-

barrel 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 ki 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 full-

barrel model. Iteration of ki 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.


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 Mbase, 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.


Figure 7. Dynamic vs. Static Boundary Conditions

2.1.3 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


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 surface-

to-surface contact interactions active. The relative values of stress are more

meaningful/insightful than the actual values of these quantities. The smaller-

radius 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


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


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).


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.


Table 1. Contact Force for 0.7 in. Indentation at Locations 1 and 2

3250 lbs1700 lbsLocation 2


12 inches3 inches

Impactor Radius



12 inches3 inches

Impactor Radius

2.1.4 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.


Figure 13. Test Specimen Incremental Loading

2.1.5 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 2009. Frame panels will be

fabricated in the fall and tested in early 2010. 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


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

2.2.1 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 152.4 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.

2.2.2 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


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/m3.

2.2.3 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


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


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 9 10 8 T 6.35mm 9 7 7

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

R6.35 R12.7

R152.4 R101.6 R50.8 R25.4

Small Impact Tips

Large Impact Tips

Large Force Sensor

Small Force

Sensor


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.

2.2.4 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


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.


0 2 4 6 8 10 12

x 10-3

0

500

1000

1500

2000

2500

3000

3500Contact Force vs Time

Time (s)

Forc

e (

N)

Figure 16. Force time history plot for radius 12.7mm, panel thickness 3.18mm, and energy 7.5J


-0.1 -0.08 -0.06 -0.04 -0.02 0 0.02 0.04 0.06 0.08 0.12.3

2.4

2.5

2.6

2.7

2.8

2.9

3

3.1

3.2Arc Position vs Time

Time (s)

Arc

Positio

n (

m)

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


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.


0 2 4 6 8 10 12

x 10-3

0

1000

2000

3000

4000

5000

6000

7000

Time (s)

Forc

e (

N)

Contact Force vs Time

FTE1 for T 3.18mm

R 12.7mm

R 50.8mm

R 152.4mm

Figure 19. Force time history plots at FTE1 for all radii and panel thickness 3.18mm

0 2 4 6 8 10 12

x 10-3

0

1000

2000

3000

4000

5000

6000

7000

Time (s)

Forc

e (

N)

Contact Force vs Time

FTE1 for T 6.35mm

R 12.7mm

R 50.8mm

R 152.4mm

Figure 20 . Force time history plots at FTE1 for all radii and panel thickness 6.35mm


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


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.


0 2 4 6 8 10 12 14 16 18 200

100

200

300

400

500

600

700

800

900

Energy (J)

Conta

ct

Are

a (

mm

2)

Contact Area 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 FTE1+

Figure 21. Contact area as a function of energy for all radii and panel thickness 3.18mm

0 2 4 6 8 10 12 14 16 18 200

50

100

150

200

250

Energy (J)

Conta

ct

Are

a (

mm

2)

Contact Area 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 FTE1+

Figure 22. Contact area as a function of energy for all radii and panel thickness 6.35mm


0 2 4 6 8 10 12 14 16 18 200

1000

2000

3000

4000

5000

6000

7000

8000

Energy (J)

Forc

e (

N)

Contact Force 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 FTE1+

Figure 23. Contact force as a function of energy for all radii and panel thickness 3.18mm

0 2 4 6 8 10 12 14 16 18 200

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Energy (J)

Forc

e (

N)

Contact Force 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 FTE1+

Figure 24. Contact force as a function of energy for all radii and panel thickness 6.35mm


0 2 4 6 8 10 12 14 16 18 200

20

40

60

80

100

120

140

160

180

200

Energy (J)

Avera

ge P

ressure

(M

Pa)

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 FTE1+

Figure 25. Average contact pressure as a function of energy for all radii and panel thickness 3.18mm

0 2 4 6 8 10 12 14 16 18 200

50

100

150

200

250

300

350

400

Energy (J)

Avera

ge P

ressure

(M

Pa)

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 FTE1+

Figure 26. Average contact pressure as a function of energy for all radii and panel thickness 6.35mm


0 1 2 3 4 5 60

500

1000

1500

2000

2500

3000

3500

4000

4500

5000

Impactor Position (mm)

Forc

e (

N)

Contact Force vs Impactor Position

FTE1 for T 3.18mm

R 12.7mm

R 50.8mm

R 152.4mm

Figure 27. Contact force as a function of impactor position for panel thickness 3.18mm

0 0.5 1 1.5 2 2.5 3 3.5 40

1000

2000

3000

4000

5000

6000

7000

Impactor Position (mm)

Forc

e (

N)

Contact Force vs Impactor Position

FTE1 for T 6.35mm

R 12.7mm

R 50.8mm

R 152.4mm

Figure 28. Contact force as a function of impactor position for panel thickness 6.35mm


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


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 T800/3900-2 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 impact-

induced damage by various NDI methods.

Table 4. Ice Impact Test Matrix

8 plies [0/45/90/-45]_s 9 9 9 9 9 9 9

16 plies [0/45/90/-45]_2s 10 10 10 10 10 10 10

24 plies [0/45/90/-45]_3s 9 9 9 9 9 9 9

Total 28 28 28 28 28 28 28 196

* Matrix set-up: 3 specimens for structural/NDI testing; 3 specimens for trial impact calibration tests; 3 specimens retained for

NDI use

Low

Veloc

Dia 3

6x10

Quasi-Isotropic

Layup

Number of Panels Needed for Each Condition *

Hail

Dia 1

12x12

Hail

Dia 2

12x12

Hail

Dia 2

(angle)

12x12

Hail

Dia 3

12x12

Low

Veloc

Dia 1

6x10

Low

Veloc

Dia 2

6x10

Panel Thickness


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. 1278-1288.

2. 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. 25-41.

3. 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. 103-121.

4. Davies, G. A. O. and X. Zhang. 1993. "Impact damage prediction in carbon

composite structures," Int. J. Impact Engineering. 16(1):149-170.

Documents

Impact Damage Formation on Composite Aircraft Structures