DOT/FAA/AR-04/40,P3 

Office of Aviation Research 

Washington, D.C. 20591 

Explicit Finite Element Modeling of 

Multilayer Composite Fabric for 

Gas Turbine Engine Containment 

Systems 

Part 3: Model Development and 

Simulation of Experiments

November 2004 

Final Report 

This document is available to the U.S. public 

through the National Technical Information 

Service (NTIS), Springfield, Virginia 22161. 

U.S. Department of Transportation 

Federal Aviation Administration
NOTICE 

This document is disseminated under the sponsorship of the U.S. 

Department of Transportation in the interest of information exchange. The 

United States Government assumes no liability for the contents or use 

thereof. The United States Government does not endorse products or 

manufacturers. Trade or manufacturer's names appear herein solely 

because they are considered essential to the objective of this report. This 

document does not constitute FAA certification policy. Consult your local 

FAA aircraft certification office as to its use. 

This report is available at the Federal Aviation Administration William J. 

Hughes Technical Center's Full-Text Technical Reports page: 

actlibrary.tc.faa.gov in Adobe Acrobat portable document format (PDF). 
 Technical Report Documentation Page 

1. Report No.

DOT/FAA/AR-04/40,P3

2. Government Accession 

No.

3. Recipient's Catalog No.

5. Report Date

November 2004

4. Title and Subtitle 

EXPLICIT FINITE ELEMENT MODELING OF MULTILAYER COMPOSITE 

FABRIC FOR GAS TURBINE ENGINE CONTAINMENT SYSTEMS 

PART 3: MODEL DEVELOPMENT AND SIMULATION OF EXPERIMENTS 

6. Performing Organization Code

7. Author(s)

Jeffrey Simmons, David Erlich, and Donald Shockey

8. Performing Organization Report No. 

9. Performing Organization Name and Address 10. Work Unit No. (TRAIS) 

SRI International 

33 Ravenswood Avenue 

Menlo Park, CA 94025 

11. Contract or Grant No. 

01-C-AW-ASU 

13. Type of Report and Period Covered

Final Report 

8/2001 – 5/2003

12. Sponsoring Agency Name and Address

U.S. Department of Transportation 

Federal Aviation Administration 

Office of Aviation Research 

Washington, DC 20591 

14. Sponsoring Agency Code

 ANE-100, ANM-100

15. Supplementary Notes

The FAA William J. Hughes Technical Center COTR was Donald Altobelli.

16. Abstract

Under the Federal Aviation Administration Airworthiness Assurance Center of Excellence and with support from the Aircraft 

Catastrophic Failure Prevention Program, SRI International collaborated with Arizona State University; Honeywell Engines, Systems 

and Services; and the National Aeronautics and Space Administration Glenn Research Center to develop a computational model for 

designing and evaluating ballistic fabric containment structures. This report describes the model and compares the results of static 

and impact penetration tests with the results of computational simulations. 

17. Key Words 

Fan containment, Jet engine, Composite fabric, 

Ballistic impact, Zylon, Kevlar

18. Distribution Statement 

This document is available to the public through the National Technical 

Information Service (NTIS) Springfield, Virginia 22161.

19. Security Classif. (of this report)

 Unclassified 

20. Security Classif. (of this page)

Unclassified

21. No. of Pages

 59

22. Price

Form DOT F1700.7 (8-72) Reproduction of completed page authorized
TABLE OF CONTENTS 

Page 

EXECUTIVE SUMMARY vii 

1. INTRODUCTION 1 

1.1 Purpose 1 

1.2 Background 1 

2. FINITE ELEMENT ANALYSIS 2 

2.1 Background 2 

2.2 The Finite Element Model for Ballistic Fabric 2 

2.2.1 Fabric Response in Uniaxial Tension 2 

2.2.2 Fabric Response Under Cyclic Loading 3 

2.2.3 Model Basics 3 

2.2.4 Damage Mechanics 4 

2.2.5 Rate Dependence for Failure 5 

2.2.6 Model Response 6 

2.2.7 Off-Diagonal and Shear Response 6 

2.2.8 Multiple Plies 7 

2.2.9 Fabric Slack 8 

2.3 Analysis Considerations 8 

2.3.1 Mesh Size 8 

2.3.2 Contact Parameters 8 

2.4 Approach for Model Development 8 

2.5 Single-Ply, Quasi-Static Uniaxial Tests 8 

2.6 Analysis of ASU Static Ring Tests 10 

2.6.1 Kevlar Ring 10 

2.6.2 Zylon Ring 12 

2.7 NASA Tests 13 

2.7.1 Ballistic Test Simulation Results 14 

2.7.2 Computational Results: Comparing With Data 15 

2.7.3 Kevlar Tests 16 

2.7.4 Zylon Tests 17 

2.8 Discussion of Results 18 

2.9 Other Items Worthy of Consideration and Further Investigation 18 

iii
3. SWATH PUSH TESTS 19 

3.1 Swath Push Test Results 20 

3.2 Multiple-Ply Swath Push Test Results 22 

3.3 Unloading Behavior 24 

3.4 Summary 26 

4. REFERENCES 26 

APPENDICES 

A—User’s Guide

B—SRI Summary Report, March 2002 

LIST OF FIGURES

Figure Page 

1 Response of Kevlar Fabric to Uniaxial Tension Loading 3 

2 Cyclic Response of Fabric to Uniaxial Tension 3 

3 Stress-Strain Curve for a Fiber 4 

4 Fiber Damage as a Function of Strain 5 

5 Uniaxial Stress-Strain Curve for Ballistic Fabric Model 6 

6 Static Ring Test Results 7 

7 Fabric Model for Uniaxial Tests 9 

8 Finite Element Model of Static Ring Test 11 

9 Calculated Force Displacement for Static Ring Test With Kevlar 11 

10 Calculated Response of Single-Layer Zylon 12 

11 Calculated Force Displacement for Static Ring Test With Zylon 13 

12 Test Setup for the NASA Ballistic Tests 13 

13 Experiment and Simulation Results for NASA Test LG408, Eight-Ply Target of 

Zylon 15 

iv
14 Dimensionless Plot of V50 Results of Fabric Armor Systems 16 

15 Simulation Results for Ballistic Tests on Kevlar Fabric 17 

16 Simulation Results for Ballistic Tests on Zylon Targets 17 

17 Load-Stroke Response of Kevlar Swaths 21 

18 Stress-Strain Response of Kevlar Swaths 21 

19 Load-Stroke Response of Zylon Swaths 22 

20 Stress-Strain Response of Zylon Swaths 22 

21 Averaging Two Kevlar Plies Displaced by a Strain of 0.01 23 

22 Strain Offset Model Results for Two-, Four-, and Eight-Ply Kevlar 24 

23 Cyclic Unloading Response of Kevlar Swaths 25 

24 Cyclic Unloading Response of Zylon Swaths 25 

LIST OF TABLES 

Table Page 

1 Properties of Zylon and Kevlar Fabrics 2 

2 Constants for Kevlar and Zylon 10 

3 Ring Test Force and Ply 12 

4 Ballistic Test Results for Kevlar 14 

5 Ballistic Test Results for Zylon 14 

6 Summary of Swath Test Data on 1.5-in.-Wide Baseline Fabric Swaths 20 

v
LIST OF ACRONYMS

ACFPP Aircraft Catastrophic Failure Prevention Program 

ASU Arizona State University 

FAA Federal Aviation Administration 

LSTC Livermore Software Technology Company 

NASA GRC National Aeronautics and Space Administration Glenn Research Center 

SRI SRI International 

vi
EXECUTIVE SUMMARY

Modeling a multilayer fabric composite for engine containment systems during a fan blade-out 

event has been a challenging task. Nonlinear transient (explicit) finite element (FE) analysis has 

the greatest potential of any numerical approach available to industry for analysis of these 

events. Significant research is still required to overcome difficulties with numerical stability,

material modeling (pre- and postfailure), and standardizing model methods to achieve accurate 

simulation of the complex interactions between individual components during these high-speed 

events. The primary focus of this research was to develop the methodology for testing, 

modeling, and analyzing a typical fan blade-out event in a multilayer fiber fabric composite 

containment system. ABAQUS FE code was used to verify the basic material model (prefailure

state) developed through laboratory testing. LS-DYNA was the primary modeling tool used in 

the explicit FE analysis of ballistic events. 

During the Fourth Federal Aviation Administration (FAA) Uncontained Engine Debris 

Characterization Modeling and Mitigation Workshop (held in May 2000 at SRI International,

Menlo Park, CA), a representative of Honeywell Engines, Systems & Services presented the 

capability of modeling complicated engine hub-burst and fan blade-out events. Predicting most 

of the event with high confidence was shown. At the same time, SRI presented their efforts on 

modeling the material characteristics within LS-DYNA and developing a new composite fiber 

material called Zylon that appeared to be stronger, lighter, and more temperature-resistant than

Kevlar. Both parties showed interest in each other’s work, and both agreed they could benefit 

from each other if collaborative mechanisms could be arranged. After the workshop, Honeywell 

and SRI contacted each other and began talks of a joint project. The FAA, the National 

Aeronautics and Space Administration (NASA) Glenn Research Center (GRC), and Arizona 

State University (ASU) were later invited into the discussion, resulting in this FAA-funded 

research under the Aircraft Catastrophic Prevention Program and the Airworthiness Assurance 

Center of Excellence Program.

The goal of this research, was to use the technical strengths of Honeywell, SRI, NASA GRC, and

ASU for developing a robust explicit FE analysis modeling methodology for the purposes 

mentioned above. Since the development of an experimental set of data to support the 

calibration of the FE models is essential, various experimental methods to measure material and 

structural response of the fabrics were conducted. NASA GRC, under the NASA Aviation 

Safety Program, conducted a series of fabric engine containment ring tests that were used for 

modeling in this program.

Each member of the team took a leadership role and developed a comprehensive report 

describing the details of the research task and the findings. The complete FAA report is 

comprised of the following four separate reports (parts 1 through 4). 

• 

• 

Part 1: Static Tests and Modeling by Arizona State University Department of Civil 

Engineering 

Part 2: Ballistic Testing by NASA Glenn Research Center 

vii
• 

• 

Part 3: Material Model Development and Simulation of Experiments by SRI

International 

Part 4: Model Simulation for Ballistic Tests, Engine Fan Blade-Out, and Generic Engine 

by Honeywell Engines, Systems & Services 

SRI’s role was to produce a FE model for multilayer fabric, implement the model into LSDYNA, determine material properties for Zylon and Kevlar from static tests, and evaluate the

model by simulating fragment impact experiments. This report describes the model development 

and the results of the computational simulations. 

The form of the fabric model and the material parameters were based on observations and 

measurements from uniaxial and cyclic load experiments on single plies of fabric performed at

ASU and SRI. Multiple fabric plies were modeled with one shell element, whose thickness was

proportional to the number of plies. A single set of material constants was determined for each 

fabric. 

Additional experiments were performed in which a fragment simulator penetrated multiple layers

of fabric wrapped around a steel ring, a configuration representative of an engine containment 

ring. Quasi-static tests were performed at ASU, and ballistic tests were performed at NASA 

GRC. The model was evaluated by simulating these tests. 

Overall, the simulation results agreed well with test results. For the quasi-static tests on Kevlar,

the calculated values for peak load and strain to failure were close to the measured average 

values. For Zylon, the calculated peak load was about 15% below the measured average for the 

quasi-static tests. 

For the ballistic tests on Kevlar, the model closely predicted the energy absorbed for the 4-, 8-,

16- and 24-ply tests, but underpredicted the energy absorbed for the 1- and 2-ply tests. For 

Zylon, the model predicted the energy-absorbed values within the test data scatter for 4, 8, and 

16 layers, but underpredicted the energy absorbed for 24 plies. 

The results showed that the choice of analysis parameters and solution algorithms (particularly 

the choice of slideline parameters) has a big effect on the energy absorbed that was calculated in

ballistic tests. Future work will need to determine the sensitivity of the model to these 

parameters. An improved treatment of rate effects is also needed. In addition, a procedure for

using multiple layers of shell elements to model many fabric plies will facilitate design 

computations. 

viii
1. INTRODUCTION.

1.1 PURPOSE. 

This research effort was undertaken as a direct result of discussions from the Fourth Federal 

Aviation Administration (FAA) Uncontained Debris Characterization Modeling and Mitigation

Workshop (held in May 2000 at SRI International). A teaming effort between government, 

academia, and industry was seen as an excellent opportunity to transition fabric modeling and 

testing research, that was being sponsored by the FAA Aircraft Catastrophic Failure Prevention

Program (ACFPP), into commercial aircraft. 

1.2 BACKGROUND.

Over the last 7 years, under the sponsorship of the ACFPP of the FAA, SRI conducted a research 

program aimed at enhancing the safety of civilian air travel by reducing the likelihood of a

catastrophic accident in the event of an uncontained engine failure. 

SRI began by reviewing advanced military armor technology to identify materials, armor

structures, and projectile defeat concepts that could be used in designing barriers to turbine

engine fragments for commercial aircraft. Experiments and computations were then performed 

to evaluate these findings. The work identified an advanced material with significantly better

resistance to fragment penetration, demonstrated barrier structures that can protect against small 

to medium turbine engine blade fragments at a tenth of the weight of aluminum barriers, and 

developed a computational capability for designing fragment barriers. Full-scale ballistic 

experiments on fuselage sections with advanced polymer fabric barriers confirmed the ballistic 

effectiveness of lightweight fabric barriers. 

The results attracted the attention of the commercial aircraft industry. Under continued 

sponsorship by the ACFPP and new grants from the Airworthiness Assurance Center of

Excellence (AACE), SRI worked with an aircraft engine manufacturer and an airframe

manufacturer to transfer this technology to industry and to address specific needs of the aircraft 

industry. This project with Arizona State University (ASU), Honeywell Engines, Systems &

Services, and National Aeronautics and Space Administration (NASA) Glenn Research Center

(GRC) aimed at adapting the SRI fabric model to address fragment containment issues, while

another project with the University of California at Berkeley and The Boeing Company 

evaluated the effectiveness of multilayer Zylon fabric in protecting fuselage structures from 

uncontained debris. 

SRI’s effort in the ASU, Honeywell, and NASA GRC program is presented here. The objective 

of this research was to develop and demonstrate a computational model of ballistic response of 

fabrics under engine containment conditions. The desired end product was a computational 

capability for designing multilayer fabric engine containment structures for commercial transport 

aircraft. SRI’s part was (1) to perform static laboratory experiments (swath tests) to measure 

response of multilayer fabrics to slow penetration, (2) to use the data and observations along with

ring test data from ASU and ballistic data from NASA GRC to develop a computational model 

of fabric response, and (3) to evaluate the model by simulating the experiments and comparing 

1
computed and observed behavior. The procedures and results of this effort are presented in the

following sections. 

2. FINITE ELEMENT ANALYSIS.

2.1 BACKGROUND. 

A computational model is being developed that can be used as a design tool for choosing or 

evaluating parameters for fragment barriers containing ballistic fabric. The design model, 

implemented into the LS-DYNA finite element code, uses shell elements with an orthotropic 

continuum formulation to model the fabric. Because calculations run relatively quickly (about 

15 minutes using a four-processor Pentium III Linux cluster for a 20,000 element simulation of a 

gas gun experiment), the model allows design parameters to be varied and evaluated. The

response of barrier characteristics, such as size of fabric, gripping conditions, number of fabric 

plies, and yarn pitch, can be simulated as well as the size, orientation, and velocity of a fragment. 

Thus, the model is intended to be useful in designing and certifying fragment barriers, i.e., in 

determining how many plies of woven fabric are needed to stop a given fragment threat and in 

computing the loads applied to the supporting structure. The physical properties of the ballistic 

fabrics used in this program, Zylon AS and Kevlar 49, are listed in table 1. 

TABLE 1. PROPERTIES OF ZYLON AND KEVLAR FABRICS 

Trade Name Zylon AS Kevlar 49 

Material 

Polybenzobisoxazole P-Aramid

Volume Density (from manufacturer) (g/cm3) 1.54 1.44

Yarn Denier—As-Ordered (g/9 km) 500 1420 

Yarn Denier—Measured (g/9 km) 500 1490 

Yarn Linear Density—Measured (mg/cm) 0.556 1.656 

Yarn Cross-Sectional Area (cm2 x 10-4) 

(in2

 x 10-5) 

3.61 

5.59 

11.50 

17.82 

Yarn Count (yarns/cm) 35 x 35 17 x 17 

Fabric Ply thickness (approx.) (in.) 

(mm)

0.008 

0.21 

0.011 

0.28 

Fabric Areal Density—Measured (g/cm2) 

(lb/ft2) 

0.01575 

0.0323 

0.02275 

0.0466 

2.2 THE FINITE ELEMENT MODEL FOR BALLISTIC FABRIC.

2.2.1 Fabric Response in Uniaxial Tension. 

The response of a single ply of Kevlar fabric tested in uniaxial tension, shown in figure 1, is 

representative of ballistic fabric, and has the following features: (1) a low initial modulus due to 

straightening of yarn crimp, followed by (2) a linear response up to initial damage (breaking of 

2
fibers), which results in (3) a reduced and nonlinear modulus up to a peak loading, followed by 

(4) a postpeak drop in strength with increased strain. 

FIGURE 1. RESPONSE OF KEVLAR FABRIC TO UNIAXIAL TENSION LOADING 

2.2.2 Fabric Response Under Cyclic Loading.

The cyclic response of ballistic fabric in tension is shown in figure 2. Zylon (figure 2(a)) and 

Kevlar (figure 2(b)) show a similar response. Unloading and subsequent reloading of the fabric 

follow a single curve that is nearly linear with a modulus that is stiffer than the original loading 

modulus for unloading before the peak stress is reached. As the fabric acquires damage, the

unloading modulus decreases. 

(a) Zylon (b) Kevlar

FIGURE 2. CYCLIC RESPONSE OF FABRIC TO UNIAXIAL TENSION 

2.2.3 Model Basics.

The ballistic fabric model being developed is an orthotropic model with a stress-strain response 

for each of the two yarn (local x and y) directions. The stress-strain responses in the two 

directions are assumed to be uncoupled, that is, the interaction between the orthogonal yarns due 

to the weave pattern is ignored. However, the failure response is coupled. 

3
2.2.4 Damage Mechanics.

The approach to modeling the response of ballistic fabric was to combine a mechanistic 

constitutive law for the fibers in the yarns in each direction with a damage law that governs fiber 

breakage. 

Figure 3 shows the stress-strain curve for a fiber in a yarn for cyclic loading. Cyclic loading is 

important because it gives data in the amount of fabric damage that has been experienced in the 

loading process. Initially, the fiber has a low modulus corresponding to the stress required to 

straighten the crimp in the yarn. Once the fiber is straightened, the response in loading is linear 

elastic up to a peak stress at which failure occurs, i.e., the fibers begin to break. On unloading

and subsequent reloading, the fiber is linear elastic with a modulus twice as stiff as the original 

elastic modulus. It is believed that the change in modulus on unloading and subsequent 

hysteresis corresponds to the rearrangement of yarns in the fabric. These effects may depend on 

fabric weave or yarn twist, but this dependence was not investigated. In compression, the model

fiber has a very small (but nonzero) modulus. 

FIGURE 3. STRESS-STRAIN CURVE FOR A FIBER 

To model damage and failure for the fabric, the mechanism of tensile failure of the fibers in the 

yarn was considered. Each yarn is typically made up of 200 to 250 fibers. As shown in figure 3, 

the fibers are assumed to be elastic until they break in tension. For the continuum model, the 

axial stress in the fabric, σ a , is assumed to be the stress from the strain in the unbroken yarns. 

Under monotonic loading, the axial stress is given by 

( σ σ Eε d ) a = a 1− ~ (1)

4
where, E is the fiber axial modulus, d is the fraction of broken yarns (which varies from zero to 

one), and 

σ

a ε

~ is the effective axial strain defined as the total axial strain, a ε , minus the crimp 

strain, cr ε . For a dynamic analysis, the rate of axial stress, σ a & , is given by 

( ) a a a σ E d ε Ed ε σ σ

~ ~ 1 & & & = − − (2)

It is assumed that the fibers do not all break at the same time, but because of imperfections in the 

fibers and differences in fiber stresses due to misalignment or different prestrain, the fibers break 

over some range of stress and strain. The fiber breakage is assumed to be a function of fiber

stress, , as shown in figure 4, namely, that at some minimum value of stress, , fibers start 

to break. Damage is assumed linear with fiber stress up to , the point of peak axial stress, 

f min f

peak f

σ max , and then damage accumulates exponentially, as the fiber stress approaches the stress at 

which all the fibers are broken, . f f

FIGURE 4. FIBER DAMAGE AS A FUNCTION OF STRAIN 

2.2.5 Rate Dependence for Failure.

The model has a rate dependency in the damage formulation that incorporates a time constant, 

, that corresponds to the time it takes fibers to break. Incremental damage to the fibers during 

a time step is calculated as 

c t

( ) 















∆ = − −

∆−

ct

t

d d d 1 e σ 0 (3)

5
where is the level of damage at the beginning of the time step, and d is the damage level 

based on the current stress state (given in equation 3). This form of rate dependence helps to 

prevent a sharp short-duration load spike (such as those experienced in initial impact for the

slideline formulation) from instantaneously failing the fabric. 

d0 σ

In incremental form, equation 4 is given by 

( ) a a a σ E d ε t E dε ~ ~ ∆ = 1− ∆ − ∆ & (4)

where ∆t is the computational time step. 

For numerical stability, the increase in damage at any computational step is limited to a

small number, e.g., 0.002. When all the fibers are broken (i.e., the damage equals one), the 

element is removed from the calculation (eroded). 

∆d

2.2.6 Model Response.

Combining the stress-strain response for a fiber (figure 3) with the damage law shown in figure 4 

gives a calculated stress-strain response for fabric under cyclic load, as shown in figure 5. The

behavior predicted by the model agrees very well with the test results over the entire range of

strain. The peak stress, the softening portion of the curve, and the unloading and reloading 

cycles are consistent. 

FIGURE 5. UNIAXIAL STRESS-STRAIN CURVE FOR BALLISTIC FABRIC MODEL 

2.2.7 Off-Diagonal and Shear Response.

Values for other moduli such as shear modulus, G, compression modulus, or crimp modulus are 

assumed to be linear elastic with values that should be very small compared with the modulus in 

tension. However, if these moduli are chosen as zero, the simulation typically crashes due to 

6
numerical problems. Therefore, the values for these moduli are set to 0.5% of the tensile 

modulus. Thus, the constitutive relations for the fabric material in the shell elements are given 

by 

σxx = f(εxx)d (5)

 σyy = f(εyy)d

 σzz = 0 

 τxy = G εxy

 τyz = G εyz

τzx = G εzx

2.2.8 Multiple Plies.

Static ring tests were performed at ASU in which an impactor was pushed through multiple

layers of fabric wrapped around a 32-inch-diameter steel ring. Multiple layers of fabric ranging

from 1 to 24 plies were tested. The impactor force and displacement were measured. The results 

are shown in figure 6(a) for Kevlar and in figure 6(b) Zylon in terms of force and ply as a

function of impactor displacement. The curves were realigned slightly to lineup at

approximately 800 pounds force to allow for the differences in initial slack in the test fixture. 

The bumps in the curves at low displacement are due to clamping of the fabric in the test setup.

Except for the two 24-ply tests, the ASU ring test results show very little difference in forcedisplacement for multiple plies of fabric, i.e., the force to displace eight layers of fabric is simply 

eight times the force for a single layer. The characteristic of the curve for 24 layers of Kevlar, 

namely, a lower stiffness and peak stress and larger strain to peak and strain to failure, is

characteristic of the layers having nonuniform pre-tension. Likewise, 24 layers of Zylon had 

lower force and ply than the tests with fewer plies. 

(a) Kevlar (b) Zylon 

FIGURE 6. STATIC RING TEST RESULTS 

Based on these results, multiple plies of ballistic fabric were modeled by simply increasing the 

thickness of the shell element by a factor equal to the number of plies. Because the fabric model 

uses shell elements with single-point integration, the element responds like a membrane, i.e., 

7
bending stiffness is negligible, and thus, the nonlinear increase in the bending stiffness with 

increased element thickness is not a problem. This model assumes that, for a multiple-ply target,

the fabric yarns are all aligned in the same direction (e.g., 0° and 90°). 

2.2.9 Fabric Slack.

The effect of fabric slack, which is due to lack of pre-tension in attaching the fabric (as opposed 

to yarn crimp, which is the waviness of the yarn due to the weaving process), can be modeled by 

adding an equivalent strain due to slack to the crimp strain. The effect of slack is to allow an 

initial strain to occur in the fabric without significant stress developing. It should be noted that 

fabric slack was not measured or modeled in this program. 

2.3 ANALYSIS CONSIDERATIONS.

2.3.1 Mesh Size. 

The finite element analyses were performed using LS-DYNA (beta Version MPP970 revision 

1405). For this investigation, the mesh size of the fabric elements, 1/4-inch square, was dictated 

by Honeywell. This is an element size that is small enough to give reasonable resolution of 

stresses, but large enough that full engine simulations can be completed in a reasonable time. 

2.3.2 Contact Parameters.

Simulations performed on this project have shown that the penetration response of the fabric can 

be strongly affected by choices of contact formulation and contact parameters. Details on these 

differences can be found in Honeywell’s project report [1]. It was found that the best choice of 

contact algorithm was to use, in keyword format 

*CONTACT_AUTOMATIC_SURFACE_TO_SURFACE 

with segment-checking for interpenetration (the parameter soft set equal to 2). All other contact 

parameters were set equal to the default values. 

2.4 APPROACH FOR MODEL DEVELOPMENT.

The approach for material model development for Zylon and Kevlar was to derive material 

parameters from simple single-ply uniaxial stress tests, and then verify the model by simulating 

the ASU quasi-static ring tests and the NASA ballistic impact tests. 

2.5 SINGLE-PLY, QUASI-STATIC UNIAXIAL TESTS.

Single-ply uniaxial stress tests were performed at SRI and at ASU on both Zylon and Kevlar. 

The tests at SRI, which are described in this report, were performed by pushing an impactor

transversely against a 10-inch-long swath of fabric that was held at both ends. The tests at ASU, 

described in reference 2, were performed by gripping and pulling in-plane 12-inch-long, 2.5-

inch-wide swaths of fabric. 

8
The results of these tests are shown in figure 7, along with the stress-strain response of a singleelement model. The lower peak stresses for the Kevlar tests performed at SRI may be due to 

failure of the yarns at the grips. Material properties for the Kevlar and Zylon model are listed in 

table 1. The meaning of the parameters is explained in appendix A. Note that the material 

constants listed in table 2 are appropriate for this size mesh. 

(a) Kevlar 

(b) Zylon 

FIGURE 7. FABRIC MODEL FOR UNIAXIAL TESTS 

9
TABLE 2. CONSTANTS FOR KEVLAR AND ZYLON 

Name Symbol Kevlar 49 Zylon 

Tensile 

Modulus 

E 10.2 Msi

(70.3 GPa) 

13.3 Msi 

(91.7 GPa) 

Peak stress σmax 0.305 Msi 

(2.10 GPa) 

0.421 Msi 

(2.90 GPa) 

Initial damage strain εmin 0.0235 0.025

Strain at peak stress εmax 0.0262 0.036

Crimp strain

x direction 

εcrx 0.01 0.037

Crimp strain

y direction 

εcry 0.01 0.006

Failure strain εfail 0.010 0.010

Crimp Modulus cr 0.091 0.047 

Compression 

Modulus 

co 0.005 0.005

Time constant tc 8 μs 2 μs 

Density ρ 2.69e-4 lbf s

2

/in3 

2.88 g/cm3

2.88e-4 lbf s

2

/in3 

3.08 g/cm3

Shell thickness t 0.0031 in./ply 

(0.008 cm/ply) 

0.0020 in./ply 

(0.005 cm/ply) 

2.6 ANALYSIS OF ASU STATIC RING TESTS.

SRI developed a finite element mesh, shown in figure 8, to simulate the static ring tests

performed by ASU. The static test apparatus, as shown in figure 8(a), was a 32-in.-diameter, 

6-in.-wide steel ring, assumed fixed in space. A 2-in.-wide by 3/8-in.-thick impactor with

rounded edges, shown in figure 8(b), was pushed through a machined 3.0-in.-square hole in the 

ring. For this configuration, the fabric model was a single layer of elements 4 in. wide. The

mesh for the fabric, as shown in figure 8(c), was uniform, 1/4-inch-square elements. The mesh 

for the fragment is also shown in figure 8(c). 

2.6.1 Kevlar Ring.

Figure 9 shows the calculated force-displacement relation for the ring test with Kevlar. As the 

stroke increases, the fabric response stiffens considerably due to both geometric effects and slack 

in the fabric. For these calculations, the impactor was given a large mass, an initial velocity of

20 m/s, and was constrained to move in the radial direction. In this way, the force on the 

impactor was calculated by multiplying the deceleration of the impactor by its mass. The 

simulation time was 80 ms. 

10
 

(a) Overall test configuration (b) Impactor and hole in steel ring 

(c) Fabric and impactor mesh 

FIGURE 8. FINITE ELEMENT MODEL OF STATIC RING TEST 

FIGURE 9. CALCULATED FORCE DISPLACEMENT FOR STATIC RING 

TEST WITH KEVLAR 

To account for variations in initial slack and zero settings between the simulation and the 

experiment, the calculated displacement was shifted (by 0.1 in.) to match the measured 

displacement at 200 lb. As shown in table 3, the average peak load for the experiments on one, 

two, four, and eight plies of Kevlar was 1639 lb with a standard deviation of 49 lb. The 

calculated peak load was 1637 lb. Based on normalized force-displacement curves (all curves 

set to equal displacements at 200 lb), the average displacement at failure in the experiments was

3.44 in. with a standard deviation of 0.10 in. The calculated value was 3.36 in. 

11
TABLE 3. RING TEST FORCE AND PLY 

Plies Kevlar (lb) Zylon (lb) 

1 1568 1739

2 1680 1704

4 1656 1840

8 1653 1644

Avg. 1639 1729

Std. Dev. 49 82 

Model 1637 1493

Figure 10 shows the calculated deformed shape of the fabric along with effective stresses in the 

fabric at four times intervals, corresponding to impactor displacement values of 1.0, 1.5, 2.0, and 

2.1 in. Ripples in the fabric are clearly seen. As shown in figure 10(d), penetration of the fabric 

initiated near the center of the fabric swath. 

(a) displacement = 1.0 in. (b) displacement = 1.5 in. (c) displacement = 2.0 in. (d) displacement = 2.1 in. 

FIGURE 10. CALCULATED RESPONSE OF SINGLE-LAYER ZYLON 

This calculation took about 30 minutes when run on four processors of a 16-processor LINUX 

cluster (eight nodes and two Intel Pentium III-900 MHz processors with 1 GB memory on the 

master node and 256 MB memory on each slave node). 

2.6.2 Zylon Ring.

Figure 11 shows the calculated force-displacement relation for the ring test with Zylon along

with the experimental results for eight plies. The response is similar to that for Kevlar, although 

the displacements are greater because the Zylon fabric has more crimp in the warp direction. To 

account for variations in the initial slack and zero settings between the experiment and 

simulation, the calculated displacement was shifted to match the measured displacement at 200

lb. As shown in table 3, the average peak load for the experiments on one, two, four, and eight 

plies of Zylon was 1729 lb/ply with a standard deviation of 82 lb. The calculated peak load was 

1493 lb (about 15% below the average of the tests). The average value of displacement at failure 

was 4.5 in. with a standard deviation of 0.10 in. The calculated value was 4.4 in. One theory for 

12
this lower prediction is that the failure is so abrupt that the numerical treatment of the contact can 

lead to early failure prediction. 

FIGURE 11. CALCULATED FORCE DISPLACEMENT FOR STATIC 

RING TEST WITH ZYLON 

2.7 NASA TESTS.

Two sets of ballistic impact tests were performed at NASA GRC, one set on Zylon and one set 

on Kevlar. The test setup, shown in figure 12, is similar to the ASU ring tests. A 10-inch-wide 

fabric swath was wrapped around a 40-inch outside diameter steel ring. The ring was positioned 

at a 15° angle to allow the impactor to clear the steel ring and impact the center of the fabric 

swath from inside the steel ring. The steel ring had a 10-inch section missing around the point of

impact. More information about the NASA Test Facility is included in reference 3. 

FIGURE 12. TEST SETUP FOR THE NASA BALLISTIC TESTS 

The impactor was a 4- by 2- by 5/16-inch-rectangular steel block with rounded edges (5/16-inch 

radius), as shown in figure 8(b), and a nominal weight of 317 g. The impactor was launched 

from a gas gun and oriented to hit the fabric, as shown in figure 12, with the narrow side facing a 

viewer looking at the side of the ring. Because the ring was positioned at a 15° angle relative to 

the trajectory of the impactor, and the fabric was pulled across the missing section of the ring, 

one corner of the impactor hit the fabric first. 

13
The velocity of the impactor varied from 345 to 915 ft/s. Tests were performed on Kevlar with 

1, 2, 4, 8, 16, and 24 plies of fabric and for Zylon 4, 8, 16, and 24 plies. A list of the test 

numbers with measured impact velocities and residual velocities are shown in tables 4 and 5. 

TABLE 4. BALLISTIC TEST RESULTS 

FOR KEVLAR 

Test 

No. 

No. 

Plies

Impact 

Velocity

(ft/s) 

Residual

Velocity 

(ft/s) 

LG434 1 384 361

LG433 1 389 367

Model 1 385 367

LG449 2 345 279

LG444 2 349 278

Model 2 347 310

LG403 4 900 845

LG410 4 912 865

Model 4 900 855

LG424 8 834 744

LG409 8 888 808

LG404 8 897 819

Model 8 900 810

LG432 16 896 650

LG429 16 914 720

Model 16 900 665

LG411 24 886 413

LG405 24 898 496

LG427 24 915 608

Model 24 900 519

TABLE 5. BALLISTIC TEST RESULTS 

FOR ZYLON 

Test No.

No. 

Plies 

Impact 

Velocity

(ft/s) 

Residual

Velocity 

(ft/s) 

LG412 4 798 733

LG406 4 895 835

Model 4 900 845

LG417 8 892 791

LG413 8 901 776

LG408 8 904 792

LG425 8 908 804

Model 8 900 778

LG426 16 911 630

Model 16 900 614

LG414 24 830 0

LG407 24 904 0

Model 24 900 480

2.7.1 Ballistic Test Simulation Results.

The NASA ballistic tests were simulated with the model. One simulation was performed for

each fabric (Kevlar or Zylon) and target thickness (number of plies). As listed in tables 4 and 

5, the impact velocity for the simulation was generally chosen as 900 ft/s, except for the 

cases with one or two plies of fabric. In those cases, a velocity near the average of the test 

values was chosen. For the simulations, the impactor was assumed to impact the fabric with

no yaw or pitch. In the tests, the yaw of the impactor estimated from the videos was typically 

about 4°. 

Figure 13 shows a direct comparison of an experiment and simulation for NASA Test 

LG408, an eight-ply target of Zylon. The measured impact velocity was 904 ft/s, and the 

measured residual velocity was 792 ft/s. For the simulation, the impact velocity was 900 ft/s, 

and the residual velocity was 778 ft/s. Figure 13 shows the deformed fabric at three

14
different times during the impact. Visually, the deflected shape of the fabric in the simulation is

very close to that of the experiment for the three times shown. The fabric is pulled out by the 

impactor in a diamond-shaped pyramid, and the calculated extent of the deformation matches the 

experiments well. The impactor starts to become visible through the fabric at 0.35 ms in both the 

experiment and the simulation. At 0.53 ms, the front of the impactor is clearly visible both in the 

simulation and the experiment, and it appears that in the model, the fabric has pulled back from

the front of the impactor more in the simulation than in the experiment. 

 Experiment Simulation Experiment Simulation Experiment Simulation 

time = 0.17 ms 0.35 ms 0.53 ms

FIGURE 13. EXPERIMENT AND SIMULATION RESULTS FOR NASA TEST LG408, 

EIGHT-PLY TARGET OF ZYLON 

2.7.2 Computational Results: Comparing With Data.

To compare the results of the experiments and simulations of the NASA ballistic tests,

normalized energy absorbed is plotted as a function of normalized impact energy. The impact 

energy is the kinetic energy of the impactor (e.g., kJ), and the absorbed energy is the difference 

between the initial and residual kinetic energy of the impactor. The energies are normalized by 

the cross-sectional area of the impactor (in this case, 4.03 cm2

), and then, to allow comparisons 

between different materials and different numbers of plies, the energies are normalized by the 

areal density of the fabric (e.g., 0.01575 g/cm2 for Zylon). The figure of merit, i.e., the 

normalized energy absorbed, is then given in units of (energy/area)/(g/area), corresponding to the 

energy absorbed for a given weight of fabric. 

This choice of normalization quantities is consistent with analyses performed by P. Cunniff of a 

large collection of ballistic performance data on different types of fabrics [4]. Figure 14 shows 

V50 (velocity at which half the projectiles are stopped) normalized by a material constant 3 ~ 1

U as 

a function of the areal density of the fabric, , the area of the projectile, , and the mass of 

the projectile, m . Figure 14 shows that a square root function is a good fit to the data, i.e., 

Ad Ap

p

2

1

3

1

50 100

~ 



 



 = p

d p

m

A A

c

U

V (6)

15
where c is a constant. Squaring both sides of the equation and rearranging gives 

3

2 2

2

50 ~ 50 2

1

c U

A A

m V

d p

p

= (7)

Thus, for these test data, the maximum kinetic energy that the target is able to stop normalized

by the areal density of the fabric, and the area of the projectile is a constant for each material

over the range of velocities and target thicknesses tested. 

FIGURE 14. DIMENSIONLESS PLOT OF V50 RESULTS OF FABRIC ARMOR SYSTEMS 

(Illustrating curve fit of experimental data taken from reference 4.) 

2.7.3 Kevlar Tests.

Figure 15 shows the results for the NASA tests on Kevlar in terms of normalized energy 

absorbed as a function of normalized impact energy. In this graph, if each ply of fabric absorbed 

the same amount of energy, the data would all lie on a horizontal line. The diagonal line

corresponds to absorbing 100% of the impact energy (i.e., stopping the impactor). The points to 

the right of the line absorb less than 100% of the impactor energy. Considering the scatter in the 

experimental data, the simulations look reasonable for targets with four plies or more. In the 

one- and two-ply tests, the model absorbs significantly less energy than in the experiments and, 

also, significantly less energy than in the other simulations. Because the energy absorbed is 

calculated based on the difference between two large numbers (impact and residual velocity), 

especially for tests away from the diagonal line of 100% energy absorbed, uncertainties in 

measuring the velocities can make for a large error in the absorbed energy. 

Model results are very sensitive to details of the impact, including stresses developed at the 

slidelines due to interpenetration of the impactor with the fabric. These slideline forces are 

typically of high stress and short duration and also appear to be somewhat random. Thus, the 

variations in the simulations do not always show clear trends. For example, from 8 to 16 plies, 

16
the model absorbs more energy, but for 24 plies, the model absorbs slightly less energy than for 

16 plies. 

FIGURE 15. SIMULATION RESULTS FOR BALLISTIC TESTS ON KEVLAR FABRIC 

2.7.4 Zylon Tests.

Figure 16 shows the results for the NASA tests on Zylon in terms of normalized energy absorbed 

as a function of normalized impact energy. For these tests, the model does a very good job of 

matching the data for 4, 8, and 16 plies, but underpredicts the energy absorbed by 24 plies, which 

stops the impactor in the experiments. In the 4, 8, and 16 plies, the normalized energy absorbed

is relatively constant both in the experiments and in the simulations. In comparison to Kevlar, 

the Zylon absorbed significantly more energy for a given weight of fabric. 

FIGURE 16. SIMULATION RESULTS FOR BALLISTIC TESTS ON ZYLON TARGETS 

17
2.8 DISCUSSION OF RESULTS. 

The model predicted absorbed energies that were in general agreement with measurements, 

although in certain instances, energies were underpredicted in the model. The peak load and 

strain to failure computed in the quasi-static tests on Kevlar agreed well with measured values; 

however, the peak load computed for Zylon was about 15% lower than measured. For the 

ballistic tests, the model predicted the energy absorbed by Kevlar in the 24-layer test, but 

underpredicted the absorbed energy in the 1-, 2-, and 4-ply tests. For ballistic tests on Zylon, the 

absorbed energy predicted by the model was within the test data scatter for 4, 8, and 16 layers; 

however, the model underpredicted the energy absorbed by 24 layers. 

The choice of analysis parameters and solution algorithms (particularly the choice of slideline 

parameters) had a big effect on energy absorbed calculated in the ballistic tests. The sensitivity 

may be related to the high-amplitude, high-frequency stresses caused by the numerical 

algorithms used to treat interfaces. The fabric probably exhibited significant damping that 

minimized these stresses. Future work should investigate material damping. 

An improved treatment of rate effects is also needed. Currently, rate effects are included as a 

simple constant that limits how fast damage can occur. Some data are available for Kevlar that 

indicate the rate sensitivity of strength and stiffness. 

During the project, Honeywell and SRI compared results obtained on different computer 

platforms and with different versions of LS-DYNA. Significant discrepancies were found, and 

Livermore Software Technology Company (LSTC) was consulted. LSTC agreed to consider 

benchmarking LS-DYNA for problems involving containment and uncontainment of engine 

debris, in the manner they do for automobile crashworthiness problems, to ensure uniform results

across platforms and LS-DYNA versions. 

A procedure for using multiple layers of shell elements to model many fabric plies will facilitate 

design computations. Honeywell began to investigate such a procedure for calculating 

penetration resistance of thick fabric layers, which is addressed in reference 1. Such an effort

needs to involve both material model development and treatment of slidelines. 

Additional information is available that would be useful in checking and further developing the 

model. In particular, the high-speed videos of the deforming fabric provided velocity history 

data that could be compared with computed histories to determine if the forces on the impacting

and decelerating fragment are correct. 

Finally, for robustness and to ensure applicability to engine burst scenarios, additional 

experimental data should be acquired. Ballistic experiments should be performed in which 

sharper projectiles more representative of fan blade fragments impinge at oblique angles on 

multiple-layer fabric targets. 

2.9 OTHER ITEMS WORTHY OF CONSIDERATION AND FURTHER INVESTIGATION.

• Sensitivity of model to slideline parameters. The cause of slideline parameter sensitivity

should be investigated. The sensitivity may be related to the nonphysical high-amplitude, 

18
high-frequency forces at yarn intersections predicted by the numerical algorithms. In 

reality, however, the fabric probably damps these forces. 

• Sensitivity of model to material parameters. Simulations of quasi-static and ballistic ring 

tests should be performed in which material constants, mesh size, and fragment

orientation (pitch and yaw) are varied to determine model sensitivity to these parameters. 

• LS-DYNA benchmarks for containment and uncontainment problems. LSTC should 

establish several well-defined scenarios as standards to allow designers to benchmark 

their computations and ensure that the LS-DYNA code gives consistent results across 

platforms and across versions. 

• Rate effects on fabric deformation and failure. The treatment of rate effects should be 

improved. Rate effects are currently accounted for by a simple time constant that limits 

how fast damage can occur. Guidance can be obtained from data that exist on the 

strength and stiffness of Kevlar fabrics. 

• Use of several layers of elements to model multiple-fabric plies. The effort begun by 

Honeywell to investigate using multiple layers of shells to model thick fabric layers 

should be continued because multilayer shells indicate better how many layers of fabric 

would be penetrated for a given threat. The effort should involve both material model 

development and slideline treatment. 

• Fabric deformation profiles. Frames from the high-speed videos of the ballistic tests 

should be analyzed to extract data for checking and refining the model. The deformation 

profiles of the impacted fabric as a function of time should be compared with velocity 

histories of the impactor to determine if the fabric model is correctly applying forces as 

the fabric deforms. 

• More representative engine burst experiments. Ballistic experiments should be 

performed in which sharper projectiles more representative of fan blade fragments 

impinge at oblique angles on multiple-layer fabric rings. The data and observations will 

help ensure applicability to engine burst scenarios. 

3. SWATH PUSH TESTS.

SRI performed a series of quasi-static push tests on 1.5-in.-wide swaths of the baseline Kevlar 

and Zylon fabrics to (1) help determine material properties for modeling the impact response of 

multiple plies of fabric, and (2) investigate the unloading and reloading response of these fabrics. 

The setup and analysis procedures for the swath push tests are described in appendix B. A 

1.5-in.-wide, 7.2-inch-long fabric swath is gripped at two ends and is quasi-statically loaded in

the center by a round-edged, 2-inch-wide penetrator. The load on the penetrator and deflection 

of the fabric are measured, and the stress and strain in the fabric is calculated using geometric 

relationships. 

19
A matrix of the parameters and results of the multiple-ply tests (tests 22-31) and cyclic unloading

tests (tests 33-35) is shown in table 6, along with those of the previously reported single-ply,

1.5-in.-wide swath push test (tests 6-10). 

TABLE 6. SUMMARY OF SWATH TEST DATA ON 1.5-in.-WIDE BASELINE 

FABRIC SWATHS 

Targetb Max. Load Energy Absorbedc SEAd

Weave 

Areal 

Test Density Fabric Width per Ply per Ply

No.a Material (Yarn/in.) 

No. 

Plies (g/cm2

) (in.) (Yarns) (lb) (lb) (J) (ft-lb) (J) (ft-lb)

( J 

g/cm2

) 

( ft-lb

lb/ft2) Comments 

6 Kevlar 17x17 1 0.02275 1.47 25 530 530 26 19 26 19 1157 417

8 Kevlar 17x17 1 0.02275 1.47 25 570 570 25 19 25 19 1113 401

7 Zylon 35x35 1 0.01575 1.43 50 760 760 40 29 40 29 2517 906

10 Zylon 35x35 1 0.01575 1.43 50 839 839 30 22 30 22 1891 681

Single-Ply 

Tests 

24 Kevlar 17x17 2 0.0455 1.47 50 1093 547 45 33 22 16 980 353

25 Kevlar 17x17 2 0.0455 1.47 50 1049 525 55 41 28 20 1219 439

31 Kevlar 17x17 4 0.091 1.47 100 1635 409 85 63 21 16 936 337

27e Kevlar 17x17 8 0.182 1.47 200 >2256 >282 >244 >180 >31 >23 >1343 >484

28e Kevlar 17x17 8 0.182 1.47 196 >2256 >282 >234 >173 >29 >22 >1286 >463

30 Kevlar 17x17 8 0.182 1.47 197 2462 308 225 166 28 21 1238 446

22 Zylon 35x35 2 0.0315 1.43 100 1229 615 >148 >109 74 54 >4684 >1686

23 Zylon 35x35 2 0.0315 1.43 100 1287 644 >119 >88 >59 >44 >3768 >1357

32 Zylon 35x35 4 0.063 1.43 200 1922 481 ---- ---- ---- ---- ---- ----

MultiplePly Tests 

Major Slip 

at Grips 

35 Kevlar 17x17 1 0.02275 1.47 25 580 580 ---- ---- ---- ---- ---- ---- Cyclic 

Unload 

33 Zylon 35x35 1 0.01575 1.43 50 735 735 ---- ---- ---- ---- ---- ---- Cyclic 

Unload 

34 Zylon 35x35 1 0.01575 1.43 50 744 744 ---- ---- ---- ---- ---- ---- Cyclic Hold

a. All tests involve the baseline fabrics used in this program. All tests used the 2-in.-wide rounded-end penetrator (same as that used by ASU), a stroke rate of 

0.05 in./s. and the yarns tested were all in the warp direction. 

b. Distance between grips was approximately 7.2 in.

c. Equals the area under the load-deflection curve. 

d. SEA = Specific Energy Absorbed = energy absorbed divided by areal density of the target 

e. Load record cut off above 2256 lb for these two tests.

3.1 SWATH PUSH TEST RESULTS.

Tests were performed to investigate the response of two, four, and eight plies of Kevlar, and 

these results, along with those from the previously performed single-ply Kevlar tests, are shown 

in figures 17 and 18. To normalize the results for all the tests, figure 17 shows the load per ply 

as a function of stroke, and figure 18 shows the stress-strain results. 

For the case of Zylon, four- and eight-ply tests could not be performed due to significant slip at 

the grips, so only two-ply tests were reported (and even those had a slight amount of slip at the 

grips). Figures 19 and 20 show the two-ply Zylon results, compared with the previously reported

single-ply results. 

20
FIGURE 17. LOAD-STROKE RESPONSE OF KEVLAR SWATHS 

FIGURE 18. STRESS-STRAIN RESPONSE OF KEVLAR SWATHS 

21
FIGURE 19. LOAD-STROKE RESPONSE OF ZYLON SWATHS 

FIGURE 20. STRESS-STRAIN RESPONSE OF ZYLON SWATHS 

3.2 MULTIPLE-PLY SWATH PUSH TEST RESULTS.

For both fabrics, the overall trend was that, as the number of plies increases, the loading modulus

and the peak load (or peak stress) decrease, while the strain at peak and strain at failure increase 

(greater spread). The area under the curve in all cases remains relatively constant. Repeat tests 

showed some scatter, but the trend was significant and consistent. 

22
This trend was consistent with combining curves that have the same shape, but were offset from 

each other. An example is shown in figure 21. Two curves for one-ply Kevlar offset by 1% 

strain are averaged. The resulting curve has the response characteristics seen in the tests: lower 

initial modulus, lower peak, and greater spread. 

FIGURE 21. AVERAGING TWO KEVLAR PLIES DISPLACED BY A STRAIN OF 0.01 

The response of multiple plies can be offset from each other for several reasons. First, in any 

multiple-ply impact configuration, inner plies are impacted before outer plies. Second, any

differences in slack among the plies will have the effect of offsetting the curves with respect to

each other. For the swath push test configuration, the two ends of the specimen are wrapped 

around a flattened bar and then clamped between two plates. As the clamping is increased, the 

outermost wrapped plies will be stretched more than the inner plies. This effect is, of course,

very dependent upon the test configuration. Presumably, if all the plies are prepared with a 

known prestress, as in the NASA GRC tests, any effect due to different amounts of slack would 

be minimized. 

A model was developed for the response of multiple plies by combining the single-ply response 

for several plies offset by a specified amount of strain. Figure 22 shows the predicted response 

of two, four, and eight plies of Kevlar, using the single-ply response with an assumed strain 

offset of 0.077 between plies. The effect of reducing the peak and spreading the curve is slightly 

overpredicted for two plies, but shows reasonably good agreement for four and eight plies. 

23
FIGURE 22. STRAIN OFFSET MODEL RESULTS FOR TWO-, FOUR-, AND 

EIGHT-PLY KEVLAR

3.3 UNLOADING BEHAVIOR.

Cyclic-loading swath push tests were performed on single plies of Kevlar and Zylon to determine 

the unloading and reloading response. For each test, the fabric was partially unloaded and then

reloaded several times throughout the test. The response of Kevlar is shown in figure 23, and the 

response of Zylon is shown in figure 24. The fabrics show similar response. In all cases, the

loading and reloading response is nearly elastic, i.e., the curves show little hysteresis. However, 

unloading does not follow the original loading curve. When the fabric is unloaded before the

peak load is reached, the unloading modulus is stiffer than the loading response. The unloading

modulus progressively decreases for unloading after the peak. As shown in figure 24, a cyclic 

test with very slight unloading (over the same time as the full unloading in the other tests) was

performed to determine how much of the unloading response was due to relaxation of the fabric 

at constant strain. The effects of relaxation were not significant. 

A possible explanation for the observed response is that during loading, the yarns rearrange and 

on unloading do not re-rearrange to the initial configuration. After peak load, fibers are broken, 

therefore, the unloading modulus decreases. 

24
FIGURE 23. CYCLIC UNLOADING RESPONSE OF KEVLAR SWATHS 

FIGURE 24. CYCLIC UNLOADING RESPONSE OF ZYLON SWATHS 

25
3.4 SUMMARY.

SRI successfully performed several quasi-static push tests on Zylon and Kevlar fabrics, which 

provided the material properties and the unloading are reloading properties. This data was used 

in the analyses performed by SRI, ASU, and Honeywell, which is documented in this

comprehensive FAA report. 

4. REFERENCES.

1. Gomuc, Reha, “Explicit Finite Element Modeling of Multilayer Composite Fabric for 

Gas Turbine Engines Containment System, Part 4: Model Simulations for Ballistic Tests, 

Engine Fan-Out, and Generic Engine,” FAA report DOT/FAA/AR-04/40,P4, 2004. 

2. Rajan S.D., Mobasher B., Sharda J., Yanna V., Deenadayla C., Lau D., and Dhah D.,

“Explicit Finite Element Modeling of Multilayer Composite Fabric for Gas Turbine

Engines Containment System, Part 1: Static Tests and Modeling, FAA report

DOT/FAA/AR-04/40,P1, 2004. 

3. Periera, M. and Revilock, D., “Explicit Finite Element Modeling of Multilayer 

Composite Fabric for Gas Turbine Engine Containment Systems, Part 2: Ballistic Impact 

Testing,” FAA report DOT/FAA/AR-04/40,P2, 2004. 

4. Cunniff, P.M., “Dimensionless Parameters for Optimizations of Textile-Based Body 

Armor Systems,” Proceedings of the 18th International Symposium on Ballistics, San 

Antonio, TX, November 1999. 

26
APPENDIX A—USER’S GUIDE

USER-DEFINED MATERIAL TYPE 47 (BALLISTIC FABRIC) 

LS-DYNA reserves material types 41-50 as user-defined materials. Material type 47 was

arbitrarily chosen for implementation of ballistic fabric. 

 Columns Quantity Format

Card 3 

1-10 E Tensile modulus E10.0 

11-20 εmin Bidirectional failure strainStrain at first nonlinearity E10.0 

21-30 εmax Strain at peak stress E10.0 

31-40 σmax Peak stress E10.0

41-50 εcrx Crimp strain, local x direction E10.0 

51-60 εcry Crimp strain, local y direction E10.0 

61-70 εsoft Softening strain, (default = 0.01) E10.0 

71-80 cr Crimp stiffness, as fraction of E (default = 0.01) E10.0 

Card 4 

 1-10 co Compression stiffness, as fraction of E (default: 0.02) E10.0 

11-20 2G Shear modulus, (default: 0.02*E) E10.0 

21-30 εtcr Transition strain for uncrimping (default = 0.005) E10.0 

31-40 εtul Transition strain for unloading (default = 0.005) E10.0 

41-50 εfail Strain at failure (default = 5.0*εmax) E10.0 

51-60 ddmax Maximum increment in damage per step (default, 0.01) E10.0 

61-70 tc Time constant for damage rate (default = 1.0e-6) E10.0 

The ballistic fabric model is orthotropic with stress-strain response as shown in figure A-1 for 

each of the two yarn (local x and y) directions. The stress-strain responses in the two directions 

are uncoupled, but the failure response is coupled. 

σxx = f(εxx) 

σyy = f(εyy) 

σzz = 0 

τxy = G εxy

τyz = G εyz

τzx = G εzx 

A-1
FIGURE A-1. UNIAXIAL STRESS-STRAIN CURVE FOR BALLISTIC FABRIC MODEL 

As shown in figure A-1, the uniaxial response has the following features: 

• 

• 

• 

• 

• 

• 

Initial crimp before straightening, prestraighted modulus = cr*E 

Linear response after straightening with modulus E up to a strain of first nonlinearity, εmin

Peak stress, σmax, and strain at peak stress, εmax

Linear elastic unloading with modulus of 2E and with reduced modulus with damage 

Postpeak softening 

Failure (element erosion) at full damage 

A.1 FITTING MODEL PARAMETERS FROM STRESS-STRAIN CURVES. 

A procedure for fitting constants to an existing stress-strain curve is described below and 

demonstrated for Kevlar 49 in figure A-2, as tested at SRI International. Note that the values 

listed in table A-1 are taken from a collection of curves; therefore, they are different from those 

given is this example. 

1. Identify the peak stress σmax. 

2. Draw a line tangent to the loading curve at half the peak stress to get the loading 

modulus, E. 

3. Fit a line to the initial portion of the loading curve to get the crimp stiffness factor, cr. 

4. For the descending portion of the curve, draw a tangent at 0.8 σmax. 

5. Identify the amount of crimp strain, εcr, where the crimp stiffness line crosses the loading 

line. 

A-2
6. Starting from εcr, identify εmin, the strain at which the curve departs from linearity, and 

the strain at peak stress, εmax. 

7. Starting from εmax, identify εsoft, where the unloading tangent crosses the x-axis. 

8. Calculate the shell element thickness as one-half the areal density of the fabric divided by 

the material density. For density, ρ, use twice the solid density of the material.

∗

(a) steps 1-4 

(b) steps 5-7

FIGURE A-2. DEMONSTRATION OF FITTING CONSTANTS TO A STRESS-STRAIN 

CURVE FOR KEVLAR 

 

∗

 The effective thickness for the fabric is given by the areal density of the fabric divided by the material density. 

However, for woven fabric, only half the yarns provide stiffness in a given direction. Therefore, in order to use

the measured modulus, only half the effective thickness of the fabric should be used for the shell thickness. To

keep the correct mass properties, the density must be twice the solid density of the material. 

A-3
TABLE A-1. CONSTANTS FOR KEVLAR AND ZYLON 

Name Symbol Kevlar 49 Zylon 

Tensile 

Modulus 

E 10.2 Msi

(70.3 GPa) 

13.3 Msi 

(91.7 GPa) 

Peak stress σmax 0.305 Msi 

(2.10 GPa) 

0.421 Msi 

(2.90 GPa) 

Initial damage strain εmin 0.0235 0.025

Strain at peak stress εmax 0.0262 0.036

Crimp strain

x direction 

εcrx 0.01 0.037

Crimp strain

y direction 

εcry 0.01 0.006

Failure strain εfail 0.010 0.010

Crimp Modulus cr 0.091 0.047 

Compression 

Modulus 

co 0.005 0.005

Time constant tc 8 μs 2 μs 

Density ρ 2.69e-4 lbf s

2

/in3 

2.88 g/cm3

2.88e-4 lbf s

2

/in3 

3.08 g/cm3

Shell thickness t 0.0031 in/ply 

(0.008 cm/ply) 

0.0020 in/ply 

(0.005 cm/ply) 

A.2 PROCEDURE FOR AVERAGING RESULTS FROM MULTIPLE STRESS-STRAIN 

CURVES. 

For multiple stress-strain curves, use the following procedure to average the curves. 

• For each curve, identify a peak stress, σ max . 

• Calculate an average peak stress, σ max . 

• For each curve, identify the strain value at half the average peak stress, 50 ε . 

• Calculate an average value for the strain at half the peak stress, 50 ε . 

• Shift curves so that strain equals 50 ε at 0.5σ max . 

• Calculate an average stress-strain curve from the shifted curves by averaging the stresses

at each strain value. 

A-4
APPENDIX B—SRI SUMMARY REPORT, MARCH 2002

B-1/B-2
Summary Report • March 2002 

Explicit Finite Element Analysis 

Modeling of Multi-Layer Composite 

Fabric for Gas Turbine Engines 

Containment Systems 

Donald A. Shockey, Associate Director

Jeffrey W. Simons, Senior Engineer 

David C. Erlich, Physicist III 

Center for Fracture Physics

SRI Project P11546 

Covering the Period from Sept 2001 through March 2002 

Prepared for: 

Department of Civil and Environmental Engineering 

Arizona State University 

Tempe, Arizona 85287-5306 

ATTN: Dr. S. D. Rajan 

 Ms. Deborah Shaver 

Approved: 

Angel Sanjuro, Director 

Materials Research Laboratory 
FIGURES........................................................................................................................... iv 

TABLES ............................................................................................................................. v 

INTRODUCTION .............................................................................................................. 1 

ANALYSIS......................................................................................................................... 1 

Finite element analysis of ASU static test with Zylon.................................................... 1 

Support for Advancing Honeywell’s Analytic Capabilities........................................... 2 

EXPERIMENTS: Push Tests on Fabric Swaths................................................................. 3 

Test Configuration .......................................................................................................... 3 

1.5-in.- wide Swaths ....................................................................................................... 6 

Effect of Mesh Density ................................................................................................... 7 

4-in.-wide Swath Tests.................................................................................................... 8 

Analysis of Swath Tests.................................................................................................. 9 

Determining stress and strain for swath tests............................................................ 10 

Determining cross-sectional area of yarns................................................................ 10 

Fabric reference web sites......................................................................................... 11 

Gluing of Zylon......................................................................................................... 11 

MEETINGS ...................................................................................................................... 11 

Appendix 1. User’s Manual for SRI Zylon Fabric ........................................................... 12 

iii
FIGURES 

Figure 1. Finite element model of ASU static test.............................................................. 1 

Figure 2. Calculated force displacement for ASU static test with Zylon. ......................... 2 

Figure 3. Experimental configuration for push tests of fabric swaths................................ 4 

Figure 4. Penetrator geometry............................................................................................. 5 

Figure 5. Kevlar push tests on 1.5 in. swaths...................................................................... 6 

Figure 6. Zylon push tests on 1.5 in. swath ........................................................................ 6 

Figure 7. Effect of mesh density on push test load-stroke results ...................................... 7 

Figure 8. Nominal Zylon stress-stress curves for different mesh density swaths............... 8 

Figure 9. Kevlar push tests on 4 in. swaths ....................................................................... 9 

Figure 10. Zylon push tests on 4 in. swaths....................................................................... 9 

Figure 11. Analysis of push test data................................................................................ 10 

Figure 12. Uniaxial stress-strain curve for ballistic fabric model..................................... 13 

iv
TABLES 

Table 1. Properties of Zylon and Kevlar Fabrics tested by SRI ......................................... 3 

Table 2. Parameters and results for PUSH tests ................................................................. 5 

v
INTRODUCTION 

As part of FAA’s AACE program entitled “Explicit Finite Element Analysis Modeling of

Multi-Layer Composite Fabric for Gas Turbine Engines Containment Systems,” SRI

International is providing experimental and computational support to the team led by 

Arizona State University (ASU) and including Honeywell Engines & Systems Phoenix 

and NASA Glenn Research Center (GRC). The overall goals of the program are to 

investigate the use of Kevlar and Zylon in containment barriers against aircraft engine 

debris and to develop finite element capabilities for barrier design and analysis. 

The program is organized in two phases. Phase I includes experiments and analyses of 

static tests performed at ASU investigating the response of a penetrator pushed through 

wrapped plies of Kevlar and Zylon. Phase 2 includes experiments and analyses of 

ballistic tests of the same configuration to be performed at NASA. This report 

summarizes SRI’s contributions to Phase 1. 

ANALYSIS 

Finite element analysis of ASU static test with Zylon 

SRI developed a finite element mesh, shown in Figure 1, for the static tests planned by 

ASU. The static test apparatus, as shown in Figure 1a, was a steel ring 32 in. in diameter 

and 6 in. wide. A 2-in.-wide by 3/8-in.-thick-penetrator with rounded edges, shown in 

Figure 1b, was pushed through a machined 3.0-in.-square hole in the ring. For this 

configuration, the fabric was a single 4-in.-wide ply of Zylon AS 35x35. 

 

a) Overall test configuration b) Penetrator and hole in steel ring 

Figure 1. Finite element model of ASU static test 

The finite element analysis was performed using LSDYNA. The finite element model 

contained 4688 solid elements for the penetrator and the steel ring and 4192 thin shell 

elements for the fabric. Figure 2 show the calculated force-displacement relation for the 

1
penetrator. As the stroke increases, the fabric response stiffens considerably due to 

geometric effects. The steps in the curve occur because the analysis was run

dynamically; the penetrator impacts the fabric at 20 m/s, and the simulation time is 3.0 

ms. The dynamic component of the response could be reduced by lowering the impact 

velocity, but at the cost of increased CPU time. This calculation took about 20 minutes 

when run on a single processor of our 16-processor LINUX cluster. 

0

500

1000

1500

2000

0.0 1.0 2.0 3

DISPLACEMENT (in.)

F

ORCE (lb)

.0

Figure 2. Calculated force displacement for ASU static test with Zylon. 

For these simulations we did not include slack in the fabric, which depends on how the 

test article is fabricated (e.g., how much the fabric is pretensioned). The amount of slack 

will change both the load and displacement response of the fabric on the penetrator. 

Without considering slack we estimated a peak load of about 1600 lb. for a single ply of 

Zylon at a displacement of about 2 in. 

Support for Advancing Honeywell’s Analytic Capabilities 

To assist Honeywell enhance analytic capabilities, SRI provided Honeywell with the 

source code for SRI Zylon fabric model and step-by-step instructions on how to 

implement user-defined materials into LSDYNA. As part of this effort, SRI provided 

several LSDYNA input files for simulating the ASU static tests, including files with the 

fabric modeled as a: (1) standard elastic material, (2) user-defined elastic material model 

41 and (3) SRI’s Zylon fabric model as a user-defined material number 47. With each 

input file SRI also provided Honeywell with Microsoft Excel spreadsheet files containing 

the calculated force-displacement results of the calculations and estimates of 

computational times for completion. SRI provided Honeywell with an input file for

SRI’s Zylon material model written in keyword format so that Honeywell could more 

easily modify the input. In addition, SRI also provided Honeywell with LSDYNA input 

files for simulations of SRI push tests P25 and P47. 

In order to implement the user-defined materials, SRI modified Honeywell’s version of 

the LSDYNA subroutine dyn21.F and included in that file the FORTRAN source code 

for SRI’s Zylon fabric model. To aid Honeywell’s understanding of the source code and 

to help simplify its use, SRI inserted explanatory comment cards for many of the source 

2
code lines and added default values for many of the required material parameters for the 

Zylon fabric model. 

SRI also provided Honeywell with a user’s guide for the Zylon material model. The 

User’s guide was written in a style consistent with the LSDYNA Structured User’s

manual. It is included in this document as Appendix 1. 

EXPERIMENTS: Push Tests on Fabric Swaths 

Test Configuration 

SRI performed a series of push tests on the Zylon AS 35x35 and Kevlar 29 17x17 fabric. 

The properties are listed in Table 1. 

The test set-up is shown in Figure 3. The test specimen was a 12-in.-long swath of fabric, 

either 4 or 1.5 in. wide, gripped at 7.25 in. and loaded quasistatically by a steel penetrator

at the center of the fabric swath. The penetrator, shown in Figure 4, is a 2.0-in.-wide 

5/16-in.-thick steel penetrator had rounded ends with a 5/32 in. radius of curvature in two 

directions. During the test force and displacement on the penetrator were measured and 

the response of the fabric in two directions was recorded on videotape. 

For the 1.5-in. swaths, the width of the fabric was less than the 2.0 in. width of the 

penetrator. Based on our previous experience with push tests of single yarns by roundedend penetrators, these fabric swath push tests should yield stress-strain curves that are 

similar to uniaxial strain tensile test results. 

Table 1. Properties of Zylon and Kevlar Fabrics tested by SRI 

WOVEN FABRICS USED IN ASU / HONEYWELL / SRI / FAA PROGRAM

Trade Name Zylon-AS Kevlar-29

Material Poly-benzobisoxazole (PBO) P-Aramid

Volume Density (from manufacturer) (g/cm3

) 1.54 1.44

Yarn Denier - Nominal (g/9 km) 500 1420

Yarn Denier - Measured (g/9 km) 500 1490

Measured Linear Density (mg/cm) 0.556 1.656

Yarn Cross-sectional Area (cm x 10-4) 3.61 11.50

(in. x 10-5) 5.59 17.82

Mesh Density (yarns/in.) 35x35 17x17

Fabric Thickness (approx.) (in.) 0.008 0.011

(mm) 0.21 0.28

Fabric Areal Density - measured (g/cm2

) 0.01575 0.02275

(lb/ft2

) 0.0323 0.0466

(oz/yd2

) 4.65 6.71

Degree of Crimp* - warp yarns (%) 3.1 1.1

* Elongation when straightened after removal from fabric

Table 2 shows a matrix of the test parameters and results, including tests on both the 1.5-

and 4-in. swaths for Kevlar and Zylon. In addition to testing different materials, tests 

were performed to compare the response of different mesh densities for Zylon (30x30, 

35x35, and 40x40) and the difference for fabrics gripped on the fill and warp yarns. 

3
Figure 3. Experimental configuration for push tests of fabric swaths 

4
Figure 4. Penetrator geometry 

Table 2. Parameters and results for push tests 

Test Target

a Areal Fabric Penetrator Stroke Peak Energy SEAd

No. Material Weave Direc- No. Density Width Typeb Rate Load Absorbedc ( J ( ft-lb

(H-P:) (Yarn/in.) tion Plies (g/cm2

) (in.) (Yarns) (in./s) (lb) (J) (ft-lb) g/cm2 ) lb/ft2 )

1 Kevlar 17x17 Warp 1 0.02275 4.0 68 RE (2" wide) 0.05 912 62.0 45.7 2723 980

4 Kevlar 17x17 Warp 1 0.02275 4.0 68 RE (2" wide) 0.05 1001 69.0 50.9 3032 1092

2 Zylon 35X35 Warp 1 0.01575 4.0 140 RE (2" wide) 0.05 1417 115.4 85.1 7325 2637

3 Zylon 35X35 Warp 1 0.01575 4.0 140 RE (2" wide) 0.05 1345 121.7 89.8 7727 2782

6 Kevlar 17x17 Warp 1 0.02275 1.47 25 RE (2" wide) 0.05 530 26.3 19.4 1157 417

8 Kevlar 17x17 Warp 1 0.02275 1.47 25 RE (2" wide) 0.05 570 25.3 18.7 1113 401

7 Zylon 35X35 Warp 1 0.01575 1.43 50 RE (2" wide) 0.05 760 39.6 29.2 2517 906

10 Zylon 35X35 Warp 1 0.01575 1.43 50 RE (2" wide) 0.05 839 29.8 22.0 1891 681

12 Zylon 30x30 Fill 1 0.0130 1.50 45 H (2" wide) 0.05 566 37.3 27.5 2869 1033

13 Zylon 30x30 Fill 1 0.0130 1.50 45 H (2" wide) 0.05 551 39.9 29.4 3067 1104

16 Zylon 40x40 Warp 1 0.0185 1.50 60 H (2" wide) 0.05 933 38.8 28.6 2095 754

17 Zylon 40x40 Fill 1 0.0185 1.50 60 H (2" wide) 0.05 643 47.2 34.8 2552 919

PARAMETERS AND RESULTS FOR QUASI-STATIC PENETRATION (PUSH) TESTS

a

 Distance between grips is 7.2 in.

b

 Rounded-end penetrator (2 in. wide, 4/16 in. thick, 5/32-in. radius of curvature in both directions) - same as one used by ASU. 

c

 Equals the area under the load-deflection curve.

d

 SEA = Specific Energy Absorbed = energy absorbed divided by areal density of the target.

5
1.5-in.- wide Swaths 

Figures 5 and 6 show the fabric load displacement curves for the 1.5-in.-wide swath push 

tests on Kevlar and Zylon. The reproducibility of the test results was excellent. A 

comparison of the results for the two materials shows that the ultimate tensile strength of

Kevlar 17x17 is only about 60% of that of Zylon 35x35, and the strain to failure is also 

significantly less. 

Figure 5. Kevlar push tests on 1.5 in. swaths 

Figure 6. Zylon push tests on 1.5 in. swath 

6
Effect of Mesh Density 

To increase understanding of the fabric's deformation and failure phenomenology, we 

performed tests on Zylon fabrics of different mesh densities with 500 denier yarns (Zylon 

30x30, 35x35 and 40x40), including fabric gripped in the fill and warp directions. Results 

of these tests confirmed the importance that crimp in the gripped yarns have on the 

resulting stress-strain curves. 

The results of these tests are shown in Figures 7 and 8. Figure 7 shows the results of ram 

load-deflection and Figure 8 shows the calculated stress-strain curves. Note the large 

difference in the stroke at the first rise in load and in the stroke at the peak load for the 

different tests. The 40x40 warp yarns have a much larger degree of crimp (5.2% increase 

in length when straightened) than the 35x35 warp yarns (2.0%) or the fill yarns of either 

mesh (0.6% crimp). The straightening of the crimp likely accounts for the longer flat 

region prior to the rise. Also notice the large effect that crimp has on the load-stroke 

curve. Although the peak stress on the 40x40 warp yarn is significantly less than for the 

40x40 fill yarns (Figure 8), because of the large amount of crimp the measured load 

(Figure 7) for the warp yarn is significantly greater than for the fill yarn. 

Figure 7. Effect of mesh density on push test load-stroke results 

7
Figure 8. Nominal Zylon stress-stress curves for different mesh density swaths. 

4-in.-wide Swath Tests 

Results of the tests on 4-in.-swaths of Kevlar and Zylon are shown in Figures 9 and 10. 

Peak loads for the 4-in.-wide fabrics are much higher than for the 1.5-in.-wide swaths. 

More significantly, very large energies are absorbed (areas under the load-deflection 

curves) after the initial peak, particularly for Zylon. As listed in Table 2, note the 

difference in the energy absorbed for Zylon as compared to Kevlar, particularly for the 4-

in-wide tests, and the even larger difference in the Specific Energy Absorbed (SEA), 

which is the energy absorbed divided by the fabric's areal density. (The areal density of 

the Kevlar fabric is about 50% higher than that of Zylon). 

8
Figure 9. Kevlar push tests on 4 in. swaths 

Figure 10. Zylon push tests on 4 in. swaths 

Analysis of Swath Tests 

Preliminary comparisons of SRI’s static test results on 1.5-in. swaths with results from 

ASU’s direct tension tests of fabric swaths results showed marked differences in the

strength and stiffness of the fabrics tested. The main source of the difference was

determined to be the procedure for calculating the cross-sectional area of the yarns. SRI

provided ASU with an explanation of the process for converting load-deflection data to 

9
stress-strain data for transverse loading tests and also the procedure to determine the

cross-sectional area of the yarns for calculating stress from force. 

Determining stress and strain for swath tests 

The method for calculating stresses and strains in the fabric from load and deflection on 

the ram is simply a function of geometry as shown in Figure 11. This procedure is only 

applicable to the 1.5-in.-wide swath tests, in which all of the gripped yarns are in direct 

contact with the penetrator and undergo the same loading geometry. 

Figure 11. Analysis of push test data. 

Determining cross-sectional area of yarns 

The method SRI used to determine the cross-sectional area of a yarn removed from the

fabric was to: (1) measure the yarn length and mass to determine linear density. This 

measurement is easy to make with an accurate analytic balance. (2) Divide the linear 

density by the bulk material density (as obtained from the manufacturers) to obtain a

cross-sectional area. The cross-sectional areas for the various yarns using this method are 

shown in Table 1. 

10
The Zylon had a measured linear density of 0.556 mg/cm (exactly equal to the nominal 

500 denier), which when divided by the 1.54 g/cm3 bulk density yields a cross-sectional 

area of 3.61x10-4 sq. cm, or 5.59x10-5 sq. in. per yarn. 

The Kevlar had a measured a linear density of 1.656 mg/cm (equal to 1590 denier, 5% 

higher than the nominal 1420 denier), which when divided by the 1.44 g/cm3

 bulk 

density, yields a cross-sectional area of 1.15x10-3 cm

2

, or 1.78x10-4 in2. per yarn. 

Fabric reference web sites 

As a result of discussions on determining cross-sectional areas for yarns, SRI located an

informative set of web sites that give definitions and descriptions of many terms used in 

discussing fabrics. The first defines cotton counts and other yarn sizing parameters: 

http://www.sonic.net/~garyh/yarns/sizing.html. A related site gives some useful basic 

yarn definitions: http://www.sonic.net/~garyh/yarns/propert.html. 

Cotton count is a measure of length per unit mass. This is inverse to Denier or decitex 

which are measures of mass per unit length or linear density. However, neither of these 

measures alone can be used to determine cross-sectional area (CSA), because a 1-to-1

correspondence between CSA and either cotton count or denier does not exist. 

Gluing of Zylon 

In planning for the static tests at ASU, we addressed the question of how to secure the 

ends for fabric rings, particularly for only a few plies of fabric. Both clamping and gluing 

were considered as alternative strategies. 

SRI inquired of Toyobo, the Zylon manufacturer, whether they were aware of some

adhesive for attaching Zylon to itself. We explained that we are considering a splice joint 

in which two ends of fabric are overlapped. We also asked if it was necessary to remove 

any coatings or sizings from the yarns. 

Toyobo advised that gluing Zylon was difficult. The adhesion of Zylon fiber is lower 

than carbon fiber. Toyobo had no applicable experience in gluing Zylon, but they 

suggested increasing the overlap length of the fabric ends and recommended trying 

various glues and advised that we may get better adhesion after cleaning the surface using 

organic solvents such as alcohol and hexane to promote adhesion. 

MEETINGSSRI personnel participated in the three project meetings held at ASU on 

this program including the kick-off meeting on September 5, 2001, the first review

meeting held October 23, 2001, and the second review meeting held January 28, 2002. 

11
Appendix 1. User’s Manual for SRI Zylon Fabric 

User-Defined Material Type 47 (Ballistic Fabric)

 Columns Quantity 

Format

1-10 Card 3 Tensile modulus, E E10.0 

11-20 Bidirectional failure strain, εmin E10.0 

21-30 Unidirectional failure strain, εmax E10.0 

31-40 Yield stress, σy E10.0 

41-50 Slack strain, εsl E10.0 

1-10 Card 4 Compression factor, fco (default 0.02) 

 E10.0 

11-20 Shear modulus, (default 0.02*E) E10.0 

21-30 Hardening factor , h (default = 0.02) E10.0 

31-40 Residual strength factor, fco (default 0.20) E10.0 

41-50 Failure strain interval, εf (default 0.10) E10.0 

61-70 Max. failure increment, dmax (default 0.01) E10.0 

12
Figure 12. Uniaxial stress-strain curve for ballistic fabric model. 

The ballistic fabric model is orthotropic with stress-strain response as shown in Figure 1 

for each of the two yarn (local X and Y) directions. The stress-strain responses in the two 

directions are uncoupled, but the failure response is coupled. 

σxx = f(εxx) 

σyy = f(εyy) 

σzz = 0 

τxy = G εxy

τyz = G εyz

τzx = G εzx

As shown in Figure 1, the uniaxial response has the following features: 

• Elastic response with modulus E up to yield stress, σy

• Linear hardening with modulus =hE 

• Elastic unloading with reduced compression modulus = fco E 

• Slack before straightening, prestraighted modulus = compression modulus 

13
• Default shear modulus is small = 0.02*E 

• Failure: 

If plastic strain reaches εmin in one direction, the strength drops to a fraction of the

stress at εmin

If plastic strain in both directions reaches εmin, the element is failed 

If plastic strain in a single direction reaches εmax the element is failed 

14