Static and Dynamic Strength of Scarf-Repaired 

Thick-Section Composite Plates 

by Bazle A. Gama, Stephane Mahdi, Curt Cichanowski, 

Shridhar Yarlagadda, and John W. Gillespie, Jr. 

ARL-CR-549 September 2004 

prepared by 

Center for Composite Materials 

University of Delaware 

Newark, DE 19716 

under contract 

DAAD19-01-2-0005 

Approved for public release; distribution is unlimited.
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Destroy this report when it is no longer needed. Do not return it to the originator. 
Army Research Laboratory 

Aberdeen Proving Ground, MD 21005-5069 

ARL-CR-549 September 2004 

Static and Dynamic Strength of Scarf-Repaired 

Thick-Section Composite Plates 

Bazle A. Gama, Stephane Mahdi, Curt Cichanowski, 

Shridhar Yarlagadda, and John W. Gillespie, Jr. 

Center for Composite Materials 

University of Delaware, Newark, DE 

prepared by 

Center for Composite Materials 

University of Delaware 

Newark, DE 19716 

under contract 

DAAD19-01-2-0005 

Approved for public release; distribution is unlimited. 
ii

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4. TITLE AND SUBTITLE 

Static and Dynamic Strength of Scarf-Repaired Thick-Section Composite Plates 

5c. PROGRAM ELEMENT NUMBER 

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622618.AH80 

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6. AUTHOR(S) 

Bazle A. Gama, Stephane Mahdi, Curt Cichanowski, Shridhar Yarlagadda, and 

John W. Gillespie, Jr. 

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Center for Composite Materials 

University of Delaware 

Newark, DE 17916 

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13. SUPPLEMENTARY NOTES 

14. ABSTRACT 

Composite structural armor is typically a sandwich construction consisting of thick-section composite, rubber, and ceramic 

layers, which are combined to provide an optimal balance of structural and ballistic performance at minimum weight. Our focus 

was on the scarf repair of the thick-section plain weave S-2 Glass/SC15 composite subjected to static and dynamic loading. 

Deliberate damage to backing plates was repaired at elevated temperatures using induction heating and at room temperature. 

The static response of control and repaired plates was compared via four-point bend testing. The effect of three scarf angles and 

four repair adhesives was quantified. Using these repair techniques, renewal of stiffness was achievable, except for the case of 

the highly ductile, low-stiffness adhesive. The renewal of moment capacity of the repair beams was highly dependent on the 

scarf angle for various adhesives, and a maximum renewal of strength was 60%. The dynamic strength of scarf patch-repaired 

composite specimens was investigated through axial compression strength testing using a split-Hopkinson bar. Under dynamic 

loading, the axial strength was found to be dependent on the scarf angle and rate of loading. Loci of failure are reported for the 

various materials, scarf angles, and loading conditions. 

15. SUBJECT TERMS

repair, induction cure, thick-section composites, axial compression, Hopkinson bar, failure analysis, punch shear test, energy 

absorption, S-2 Glass/SC15, hybrid 

16. SECURITY CLASSIFICATION OF: 

19a. NAME OF RESPONSIBLE PERSON

Bazle A. Gama 

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UNCLASSIFIED 

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UNCLASSIFIED

c. THIS PAGE 

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OF ABSTRACT 

UL 

18. NUMBER 

OF PAGES 

 38 19b. TELEPHONE NUMBER (Include area code) 

302-831-8352 

Standard Form 298 (Rev. 8/98) 

Prescribed by ANSI Std. Z39.18
iii

Contents 

List of Figures iv

List of Tables iv

Acknowledgments v

1. Introduction 1

2. Test Methods 3

3. Materials and Repair Procedure 4

3.1 Materials..........................................................................................................................4 

3.2 Repair Procedure .............................................................................................................5 

4. Induction Repair 5

5. Static Performance of Scarf-Repaired Composite Beams 7

6. Dynamic Axial Compression of Scarf-Repaired Composite Plates 9

7. Conclusions 12

8. References 14

Distribution List 16 
iv

List of Figures 

Figure 1. Cross section of a model four-layer structural armor (5). ...............................................1 

Figure 2. Ballistic damage of a composite structural armor: (a) the ceramic tiles after the 

cover layer have been removed and (b) cross section (8)..........................................................2 

Figure 3. Four-point bending testing of scarf-repaired composite backing plate...........................3 

Figure 4. Hopkinson bar procedure for dynamic axial compression of scarf repair composite 

specimens...................................................................................................................................4 

Figure 5. Induction-heating of the elevated-temperature cure adhesives. ......................................6 

Figure 6. The renewal in static strength of the repaired backing plates. ........................................8 

Figure 7. Hopkinson bar responses of Hysol EA9359.3 induction-cured scarf-repaired 

composite specimens under axial compression. ......................................................................10 

Figure 8. Stress equilibrium in Hysol EA9359.3 scarf-repaired composite specimens................10 

Figure 9. Stress analysis of scarf-repaired composite specimens under dynamic axial 

compressive load......................................................................................................................11 

Figure 10. Axial compressive strength of scarf-repaired composite specimens under dynamic 

compressive load......................................................................................................................11 

Figure 11. Fracture surfaces of scarf-repaired composite specimens under dynamic axial 

compression. ............................................................................................................................11 

Figure 12. Stresses at failure in the rotated coordinate system of the scarf-repaired composite 

specimens.................................................................................................................................12 

List of Tables 

Table 1. Bending stiffness (per unit width) of the repaired backing plates. ...................................7 

Table 2. Bending moment at failure of the repaired baking plates.................................................8 
v

Acknowledgments 

The authors wish to acknowledge the financial support provided by the Strategic Environmental 

Research and Development Program (grant no. DAAL01-98-K-0058) and the Composite 

Materials Technology Collaborative Program sponsored by the U.S. Army Research Laboratory 

under Cooperative Agreement DAAD19-01-2-0005.
vi

INTENTIONALLY LEFT BLANK. 
1

1. Introduction 

Composite structural armor typically consists of different material layers stacked together to 

provide unique structural and ballistic properties, as well as satisfying other multifunctional 

requirements (e.g., fire, smoke, and toxicity [FST] resistance, electromagnetic shielding, etc.) 

(1–4). This program used a simplified four-layer configuration that consists of a composite 

cover layer for durability, a layer of ceramic tile for ballistic protection, a rubber layer, and a 

thick composite backing plate for structural and ballistic performance (figure 1). Fabrication of 

the composites used the vacuum assisted resin transfer molding (VARTM) process, which has 

been shown to provide superior mechanical properties in a single-step operation as compared to 

bonding individual layers in a multistep process (5). Additional details on structural behavior 

can be found in (5–7). 

Figure 1. Cross section of a model four-layer structural armor (5). 

Under ballistic impact, both the cover layer and the backing plates could be completely 

perforated, and the extent of damage in the ceramic tiles varied from a single tile to all of the 

tiles surrounding the impact site (see figure 2). Multiple interfaces in the backing plate also had 

delaminations that were larger than the extent of damage to the ceramic tiles (i.e., three tiles in 

length). Extensive fiber damage also occurred at the impact site. Experiments have shown that 

such damage degrades the ballistic performance of the structural armor (8). The compression 

strength after ballistic impact was also shown to drop to levels approaching 25% of the virgin 

strength (9). Repair that is capable of renewing the structural and ballistic performance after a 

ballistic impact is a key issue for the use of composites for Future Combat Systems (FCS). 

The extent of damage in different layers determines the level of repair to be performed. 

Different repair strategies and repair methods that use conventional repair techniques or 

induction curing techniques have been documented in previous studies (8, 10). Three levels of 

repair have been identified according to the level of damage through the thickness of the armor. 

Level I is concerned with the repair of the cover layer only, as in the case of damage due to a 

low-velocity impact. Level II represents the case of the repair of both the cover layer and the 

ceramic strike face. Finally, Level III addresses the repair of all the layers that compose the 

Resin-filled gap between

the ceramic tiles:

0.5mm thick rubber pad

13.2mm
2

structure, including the repair of the composite backing plate, as in the case of the high-velocity 

ballistic impact damage shown in figure 2. A repair by resin infusion was attempted (8), but it 

was shown to provide only moderate improvement in the ballistic performance of the repaired 

panels. Because resin infusion does not effectively repair fiber damage, the present study 

evaluates a scarf repair that enables all damaged materials to be replaced for Level III repairs. 

Figure 2. Ballistic damage of a composite structural armor: (a) the ceramic tiles after the cover layer have 

been removed and (b) cross section (8). 

The repair should renew stiffness, strength, and ballistic performance to meet design 

requirements. The present effort on scarf repair builds on the procedures developed by Gama 

et al. (10) who demonstrated a Level III repair of structural armor. The repair involved the 

removal and replacement of all the layers shown in figure 1, including that of the composite 

backing plate. Following the removal of the damaged cover layer and the damaged ceramic tiles, 

the damaged region of the backing plate was removed using a wet grinding tool to form a 

circular hole. The edge of the hole was machined to produce a 45° scarf angle. The backing 

plate was then repaired with an adhesive-bonded flush scarf plug of the same material, and the 

ceramic and cover layer were replaced. The repair to the thin cover layer used a flush plug repair 

(i.e., butt joint). The repair of the entire cross-section was done in a single-step operation from 

one-side that used a vacuum bag for consolidation and induction heating for rapid heating and 

cure. Various susceptors were used at the various interfaces to locally heat the adhesive bond 

lines to cure temperatures without overheating any single interface. A repaired demonstrator 

panel was manufactured to prove process viability, but was not tested. These repair techniques 

were used in the present study to fabricate test specimens to characterize the structural 

performance of the backing plate with scarf repair subjected to static and dynamic loading. 

Composite structural armor is subjected to bending moments and shear forces due to terraininduced loads and lateral impacts. Some understanding of the complex interaction between 

layers of structural armor is needed to establish simple test methods that evaluate Level III repair 

approaches and maintain some relevance to the application loads. In previous work (11), the 

finite element method was used to model the deformation behavior of the armor in bending. The 

through-thickness strain distribution deviates greatly from that of the linear classical analysis due 

to the compliant rubber layer that decouples the ceramic from the composite backing plate. 

These results show that it is possible to idealize the behavior of the backing plate in the structural 305 mm 40 mm

(a) (b)
3

armor as a beam subjected to bending loads. In addition, in-plane axial compression loads can 

also be used to evaluate the strength of bonded joints. The present study focused on developing 

simple test methods to quantify the performance of scarf-repaired thick-section composite 

backing plates of composite structural armor. 

2. Test Methods 

The four-point bending test was selected to characterize the static behavior of the virgin and 

repaired beams. The nominal dimension of the test beam was 889 × 30 × 13.2 mm. The support 

span was 762 mm, which provided a span-to-thickness ratio of 57.7. The span of the loading 

noses was sufficiently large (381 mm) to include the scarf repair. The specimen and test 

configurations are shown in figure 3. An Instron 8562 (servo-hydraulic) machine equipped with 

a custom built four-point bend test fixture (support and loading nose diameter of 25.4 mm) was 

used to conduct the experiment at a constant crosshead displacement rate of 2.5 mm/min. Load 

and crosshead displacement data were collected using the Instron Series IX data acquisition 

software. 

101.6 mm

190.5 mm 63.5 mm

381 mm

(a) 3-D view of the scarf-repaired beam (b) Schematic of four-point bend test 

Figure 3. Four-point bending testing of scarf-repaired composite backing plate. 

Dynamic axial compression tests of scarf-repaired composite joints were performed using the 

compression split Hopkinson pressure bar (SHPB) technique (12) (figure 4). The results 

generated were representative of the response of the scarf repair joints loaded in axial 

compression. The striker bar, the incident bar, and the transmitter bar were all made from 

Inconel 718 alloy (Young’s modulus–200 GPa, bar velocity–4920 m/s, Poisson’s ratio–0.29, and 

diameter–19.05 mm). Specimens of a nominal cross section (13.5 × 13.5 mm) were machined 

from scarf-repaired composite plates. The lengths of the specimens were 37.5, 70, and 90 mm 

for 45, 18.4, and 11.3° scarf angles, respectively. In order to load such long specimens for 

sufficient duration, a long striker bar (711 mm) was used to produce a long incident pulse. A 

rubber disk was used between the striker bar and the incident bar to shape the pulse and to 

uniformly load the specimen. The impact velocity of the striker bar was varied between 5 and 

10 m/s, which produced a displacement rate of 2.5–5 m/s. Maximum force at failure was 

calculated using a “3-wave” analysis from the incident, reflected, and transmitted stress pulses as 

obtained from each test. A minimum of two specimens was tested for each scarf ratio and 

adhesive used. 
4

Figure 4. Hopkinson bar procedure for dynamic axial 

compression of scarf repair composite specimens. 

3. Materials and Repair Procedure 

3.1 Materials 

The backing plate considered in the present work consisted of 22 layers of Knytex plain weave 

S-2 Glass* fabric of areal density 0.81 kg/m2

. The lay-up orientation used was (0°/90°) (i.e., the 

fabric warp direction is at 0°, and the weft direction is at 90°). The backing plates were 

fabricated by VARTM (5), using an Applied Poleramic SC15 epoxy resin system, specifically 

developed for the VARTM process. SC15 has a tensile modulus of 2.7 GPa and strain to failure 

of 6%. The SC15 system gels at room temperature (8–12 hr). Additionally, a 4-hr post-cure at 

149 °C was used. The total thickness of the backing plate was nominally 13.2 mm with a 

volume fraction of ~50% and <1% void content. 

The choice of a proper adhesive system for repair of structures is always crucial. The selected 

adhesive must offer good mechanical properties, high toughness, and meet service temperature 

requirements. It has to be polyvalent and adhere well to different materials. Also, a moderate 

viscosity is needed to fill the gaps in between the ceramic tiles and the adhesive should be easily 

spread. Finally, cure at a relatively low temperature in a short period of time is desirable. 

Four adhesives were chosen for this study that offered a broad range of properties. The Dexter 

Hysol†

 EA9359.3NA adhesive is a two-part system that cures at 82 °C for 1 hr. According to the 

manufacturer’s datasheet (13), it has a bulk tensile modulus of 2.2 GPa, a tensile strength of 

31 MPa, and an elongation at failure of 10%. The Dexter Hysol EA9394 adhesive is also a 

two-part system that cures at 66 °C for 1 hr. According to the manufacturer’s datasheet (13), it 

has a bulk tensile modulus of 4 GPa, a tensile strength of 46 MPa, and a much lower elongation 

at failure of 1.7%. Both systems have been successfully cured using induction heating of 

structural armor. The use of a room-temperature cure adhesive was also considered. Plexus‡

 *

 S-2 Glass is a registered trademark of Owens Corning. 

†

 Hysol is a registered trademark of the Dexter Corporation. 

‡

 Plexus is a registered trademark of ITW Plexus, a business unit of Illinois Tool Works. 

Incident

Bar

Transmitter

Bar

Specimen

Incident

Pulse

Reflected

Pulse

Transmitted

Pulse

Scarf Repair

Incident

Bar

Transmitter

Bar

Specimen

Incident

Pulse

Reflected

Pulse

Transmitted

Pulse

Scarf Repair
5

MA425 is a two-part methacrylate adhesive that cures in 30 min at room temperature. This 

adhesive provides good gap filling capability as well as a much higher elongation to failure 

(120%). This adhesive has a lower tensile modulus (345 MPa) and strength (17 MPa) compared 

to the elevated temperature cure adhesives. The fourth adhesive was also a methacrylate (Plexus 

AO420), which has similar mechanical properties, but a much lower curing time of 6 min. 

3.2 Repair Procedure 

The scarf patch repair concept shown in figure 3 is a very efficient way of repairing highly 

loaded composite structures. Care must normally be taken to ensure that the scarf angle is low 

enough to allow for a smooth stress transfer between the two adherends. Scarf repairs of thin 

aerospace structures are commonly limited to angles ranging between 2 and 6°. This is not 

practical in thick-section laminates. In the present work on the repair of thick sections, the scarf 

patch was limited to a diameter of 350 mm (~3 tile diameters, where extensive fiber damage 

occurs). Therefore, the maximum allowable scarf angle was ~11° (i.e., 1/5 scarf ratio for a 

13.2-mm-thick adherend). Three scarf angles were investigated using our test methods. The 

angles included 45 (1/1 scarf ratio), 18.4 (1/3 scarf ratio), and 11.3° (1/5 scarf ratio). 

Twenty backing plates, 889 mm long and 30 mm large, were fabricated by the VARTM 

process (5) and used for the characterization of the static and ballistic performance of the repairs. 

Two backing plates were kept virgin and tested as control beams. Eighteen beams were 

machined to receive the scarf patch repairs, two beams for each repair adhesive and scarf angle 

combination. For the repair, a distance equal to 101.6 mm separates the lower tips of the scarf, 

as shown in figure 3b. The placement of the repair patch only required that the surfaces in 

contact be completely wetted with the repair adhesive. The repair stack was then placed in a 

vacuum bag for hardening of the repair adhesive. 

The backing plates repaired using the elevated-temperature cure Hysol adhesives were cured by 

induction-heating technique as described in the next section. The Plexus repaired beams were 

cured at room temperature. The quality of the control and repaired beams was visually observed 

to be very good. 

4. Induction Repair 

Induction heating is a noncontact method by which electrically conductive materials (susceptors) 

are heated in an electromagnetic field. This technique has successfully heated multiple interfaces 

as shown in figure 1 in a single-step operation, with the use of appropriately placed 

susceptors (14). A stainless steel mesh susceptor, with a wire density of 5 × 5/cm2

 and a wire 

diameter of 0.165 mm, was used in the present study. The stainless steel mesh was cut to the 

shape of the area to be bonded, cleaned with acetone, impregnated with adhesive, and then 

placed in the bond area between the parent and patch laminate (which was also covered with 
6

adhesive). The stack was then vacuum bagged and placed horizontally under the induction coil 

at an optimal stand-off distance. The power setting of the induction generator, the coil shape 

(see figure 5), and the coil stand-off distance were selected for uniform heating of the bond line 

to the adhesive cure temperature. This relationship was established experimentally by using an 

actual repair backing plate that incorporated the susceptor mesh wrapped in a Kapton*

 film to 

allow multiple heating cycles. A two-turn rectangular coil was used for the fabrication of the 

45 and 18.4° scarf repair as shown in figure 5a. A four-turn spiral coil (shown in figure 5b) was 

used for the fabrication of the 11.3° scarf repairs. The 11.3° scarf repairs were bonded in two 

passes, in order to ensure a complete coverage of the bond area with the induction-heating field. 

The stack was placed underneath the induction coil at a selected distance, and the power was 

increased from 50% to 100%, in 10% increments. Process temperatures were recorded with a 

thermal camera and an embedded thermocouple. 

 (a) two-turn coil used for 18.4° scarf repair and temperature profile 

(b) four-turn coil used for 11.3° scarf repair and temperature profile 

Figure 5. Induction-heating of the elevated-temperature cure adhesives.

An AGEMA Thermo-vision† 900 thermal camera was positioned in front of the setup to capture 

the full-field surface temperature of the bond area in real time. An E-type thermocouple was 

placed at the susceptor/adherend interface to monitor the internal bond-line temperature. The 

bondline temperature was slightly higher than the surface temperature due to heat losses. 

Steady-state was reached in a few minutes for each increment in power. Once a susceptor 

steady-state temperature of ~82 (i.e., the cure temperature of Hysol EA9359.3) and 66 °C 

(i.e., the cure temperature of Hysol EA 9394) was achievable within the range of the power 

 *

 Kapton is a registered trademark of DuPont. 

†

 Thermo-vision is a registered trademark of AGEMA Company. 

°F

°F 4-Turn Coil

Left Pass Left Pass

°F

2-Turn Coil °F
7

settings, the stand-off distance was recorded, and this combination was used for further bonding 

trials. The steady-state surface temperatures, recorded from the thermal camera, are presented in 

figure 5 for the 18.4° scarf and the 11.3° scarf repairs. 

5. Static Performance of Scarf-Repaired Composite Beams 

The load-deflection behavior of all the specimens tested was observed to be linear elastic until 

failure. Because the span-to-thickness ratio of the beams was large (i.e., 57.7), classic beam 

theory was used to experimentally determine the bending stiffness per unit width of the beams 

under four-point bending. 

E I b P D L b f / ( / ) ( /96)/ 3

1 = ⋅ , (1) 

where Ef is the flexural modulus, I is the moment of inertia of the cross section, P and D1 is the 

instantaneous load and cross-head displacement in the linear-elastic region, L is the length of the 

support span, and b is the width of the beam. A comparison of the bending stiffness of the 

control beams with that of the repaired beams is presented in table 1. Results showed a 100% 

renewal of structural stiffness was attainable with the elevated-temperature cure repair adhesives. 

However, the room-temperature cure repaired backing plates were more compliant by ~15% due 

to the lower modulus of the Plexus adhesive (i.e., 0.345 vs. 2.2 GPa for EA9359.3NA). Because 

stiffness is typically the critical design parameter for composite armor structures, the lower 

stiffness of the Plexus repaired beams may prevent the use of the Plexus adhesive for repair of 

composite armor. However, it should be pointed out that the stiffness loss in a 3-D scarf repair is 

likely to be less than that measured in the 2-D beam specimen. 

Table 1. Bending stiffness (per unit width) of the repaired backing plates.

Repair Type Bending Stiffness 

(kN-m) 

Undamaged 5.75 

Elevated-temperature cure repaired 5.75 

Room-temperature cure repaired 5.00 

The strength renewal of repaired beams was evaluated using the ratio of moment capacity of a 

repaired beam to that of the control beam. The two baseline specimens (i.e., no repair) failed at 

an average bending moment of 8.79 kN-m/m (per unit width). The failure mode was observed to 

be a compressive failure of the fibers on the specimen surface layer. This failure mode did not 

cause catastrophic failure of the beam, and only a minor change in compliance resulted. 

However, the moment capacity at the onset of damage was used as the baseline for quantifying 

the effectiveness of the repairs. The bending moment at failure of all the repaired backing plates 

tested is summarized in table 2. The values given are an average value from two beams tested 

with minimal variation. In general, a major reduction in the moment capacity was observed and 
8

Table 2. Bending moment at failure of the repaired baking plates. 

Bending moment at failure / unit width (kN-m/m)a

Hysol 

EA9359.3NA 

Hysol 

EA9394 

Plexus 

MA425 

Plexus 

AO420 

45° scarf (1/1) 1.97 1.65 1.13 0.95 

18.4° scarf (1/3) 4.60 3.34 3.36 4.05 

11.3° scarf (1/5) 5.49 3.83 5.91 5.44 

a To be compared with that of control beams, i.e., 8.79 kN-m/m. 

was strongly dependent on scarf angle and adhesive type. Furthermore, the repaired beams 

failed suddenly at the adhesive bond line (in contrast to the progressive failure of the baseline). 

The extent of strength renewal (i.e., renewal of bending moment capacity), with respect to scarf 

angles and repair adhesives, is shown in figure 6a and 6b and ranges from 10% to 60%. The 

strength renewal of the induction-cured Hysol adhesives is shown in figure 6a. The repair 

efficiency of the 45° scarf angle was low (~20%) and was almost independent of the adhesive 

system used. The structural performance of the 18.4 and 11.3° scarf angles improved 

significantly as the scarf angle was reduced. It is observed in figure 6a that the Hysol EA9359.3 

repaired beams restored as much as 62% of the control strength of the backing plates, compared 

with only 43% for the beams repaired with the EA9394 adhesive. The improved performance of 

the Hysol EA9359.3 may be attributed to the higher elongation and toughness of this adhesive 

system. In a recent study (15), induction-cured EA9359.3 single-lap shear joints tested in 

tension and four-point bending were also shown to be stronger and tougher than similar EA9394 

joints. 

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25 30 35 40 45 50

EA 9359.3

EA 9394

11.3° - Scarf: Interfacial

(adhesive/parent)

18.4° - Scarf: Interfacial

(adhesive/parent)

45° - Scarf:

Interfacial

(adhesive/mesh)

Scarf Angle, deg.

Strength Renewal, %.

0

10

20

30

40

50

60

70

80

0 5 10 15 20 25 30 35 40 45 50

AO 420

MA 425

11.3° - Scarf: Interfacial

(adhesive/parent)

18.4° - Scarf: Interfacial

(adhesive/parent)

45° - Scarf:

Adhesive failure

Scarf Angle, deg.

Strength Renewal, %.

(a) Elevated-temperature cure adhesive, Hysol. (b) Room-temperature cure adhesive, Plexus. 

Figure 6. The renewal in static strength of the repaired backing plates.

The loci of failure of the repaired backing plates were, nevertheless, found to be similar for both 

elevated-temperature cure repair systems. The 45° scarf-repaired beams were seen to fail by an 

interfacial failure at the adhesive/metal mesh interface. Decreasing the scarf angle to 18.4° 
9

resulted in a slight change in the failure mode, and the locus of failure was observed to be 

interfacial, but 50% between the metal mesh susceptor and the adhesive (top end of the scarf), 

and 50% between the adhesive and the parent (lower end of the scarf). The 11.3° scarf-repaired 

beams failed by an adhesive/parent interfacial failure. Some adherend failure was also observed 

to have occurred. 

The strength renewal with respect to scarf angle for the Plexus adhesives is shown in figure 6b. 

The structural performance was mainly dependent on the scarf angle because the adhesive 

properties were similar. The repair efficiency was slightly lower (~10%) for the 45° scarf angle 

compared to the Hysol adhesives. However, the repair efficiency increased to 40% and 65% for 

the 18.4 and 11.3° scarf-repaired backing plates, respectively. This level was comparable to that 

of the induction-cured EA9359.3NA repaired backing plates. The locus of failure of the 45°

scarf was in the adhesive. Decreasing the scarf angle to 18.4° promoted a mixed, 

adhesive/interfacial failure. The 11.3° scarf-repaired backing plates failed by interfacial failure. 

In summary, the 11.3° scarf angle enabled renewal of moment capacity approaching 65% of the 

static baseline. This range of scarf angles is considered practical for the repair of thick-section 

backing plates used in structural armor. 

6. Dynamic Axial Compression of Scarf-Repaired Composite Plates 

Dynamic axial compression of scarf-repaired composite specimens was performed using the 

Hopkinson bar technique. The Hopkinson bar responses for Hysol EA9359.3 adhesive repaired 

specimens are presented in figure 7. All of these specimens failed under dynamic loading. At 

lower impact velocity, a rubber pulse shaper generates an incident pulse that is almost triangular 

in shape. The incident pulse becomes trapezoidal in shape at higher impact velocities. The 11.3°

scarf specimens did not fail when the pulse shaper was used, and thus these tests were performed 

without a pulse shaper. 

The incident, reflected, and transmitted pulses were used to calculate the forces at the incident 

bar-specimen (IB-S) and specimen-transmitter bar (S-TB) interfaces, F1 and F2, respectively, 

following the procedure described in reference (12). The condition of stress equilibrium was 

checked using the nonequilibrium parameter, R = 2(F1 – F2) / (F1 + F2) (15). Figure 8 shows the 

bar-specimen interface forces and the nonequilibrium parameter calculated for the 45 and 11.3°

specimens (bar responses presented in figure 7). Stress equilibrium (R = 0.09) was achieved in 

the 45° specimen only at maximum/failure load. However, better stress equilibrium was 

achieved in the 11.3° specimen (R = 0.06 at failure). The average maximum force, Fmax, was 

used to calculate the average axial strength of the specimen. Because the thickness of the 

adhesive bond is small as compared to the length of the incident bar, it is assumed that a volume 

element in the adhesive layer is under stress equilibrium (see figure 9). The axial strength was 

then transformed into normal and shear stresses along the failure plane (figure 9b). 
10

Figure 7. Hopkinson bar responses of Hysol EA9359.3 induction-cured scarf-repaired 

composite specimens under axial compression.

-30

-25

-20

-15

-10

-5

0

5

10

0 50 100 150 200 250 300 350-12

-10

-8

-6

-4

-2

0

2

4

R

Average Force in the Specimen, F

S-TB Interface Forces, F2 IB-S Interface Force, F1

Time, t, μs.

Bar-Specimen Interface Forces, F, kN.

R = (F1 - F2)/(F1 + F2).

-60

-50

-40

-30

-20

-10

0

10

20

0 50 100 150 200 250 300 350-3.0

-2.5

-2.0

-1.5

-1.0

-0.5

0

0.5

1.0

R

Average Force in the Specimen, F

S-TB Interface Force, F2 IB-S Interface Force, F1

Time, t, μs.

Bar-Specimen Interface Forces, F, kN.

R = 2(F1 - F2)/(F1 + F2).

(a) Specimen with 45° - scarf (b) Specimen with 11.3° - scarf 

Figure 8. Stress equilibrium in Hysol EA9359.3 scarf-repaired composite specimens.

Figure 10 shows a plot of axial strength of the specimen as a function of scarf angle for 

quasi-static and dynamic loading cases for the various adhesives. One important observation 

from these tests is that the dynamic axial strengths of the scarf joints were higher than the 

quasi-static axial strength by a factor of 2-3 for each scarf angle. The influence of scarf angle 

was similar to the results shown in figure 5 for the four-point beam tests. The dynamic axial 

strength increased as the scarf angle decreased, for all adhesives under dynamic and quasi-static 

loading conditions (with the 11.3° scarf-repaired Plexus MA425 specimen being the exception). 

-300

-200

-100

0

100

200

0 100 200 300 400 500 600 700 800

11.3° - Scarf

18.4° - Scarf

45° - Scarf

Reflected Pulses

Transmitted

Pulses

Incident Pulses

Time, t, μs.

Stress on the Bars, 

σ, MPa.
11

θ

Fmax Fmax

σ X

σ X

X

Y 1

2

θ

Fmax Fmax

σ X

σ X

X

Y 1

2

σ1

12 τ

12 τ

σ1

σ 2

σ 2

σ1

12 τ

12 τ

σ1

σ 2

σ 2

(a) Notations (b) Volume Element 

Figure 9. Stress analysis of scarf-repaired composite specimens under dynamic axial compressive load. 

0

50

100

150

200

250

300

350

400

450

0 5 10 15 20 25 30 35 40 45 50

o

o

o

o

o

o

EA 9359.3

EA 9394

11.3° - Scarf:

Interfacial (adhesive/

mesh/parent)

11.8° - Scarf: Interfacial

(adhesive/mesh/parent)

45° - Scarf:

Interfacial

(adhesive/

mesh)

Dynamic

Quasi-Static

Scarf Angle, θ, deg.

σ

FX, MPa.

0

50

100

150

200

250

300

350

400

450

0 5 10 15 20 25 30 35 40 45 50

MA 425

1/3-scarf:

Interfacial

(adhesive)

11.3° - Scarf:

Interfacial

(adhesive)

45° - Scarf:

Interfacial

(adhesive) Dynamic

Quasi-Static

Scarf Angle, θ.

σ

FX, MPa.

(a) Elevated-temperature cure adhesive, Hysol (b) Room-temperature cure adhesive, Plexus

Figure 10. Axial compressive strength of scarf-repaired composite specimens under dynamic compressive load. 

The loci of failure were found to be 50% in the adhesive/metal mesh interface and 50% between 

the adhesive and parent material in the case of Hysol adhesives (figure 11a), except for the 45°

scarf for which the failure was 100% in the adhesive metal mesh interface. In the case of the 

Plexus adhesive joint, the loci of failure were found to be in the adhesive (figure 11b). These 

failure patterns were very similar to those obtained from quasi-static four-point bend tests. 

(a) Hysol EA 9359.3 (b) Plexus MA 425 

Figure 11. Fracture surfaces of scarf-repaired composite specimens under dynamic axial compression.
12

The axial compressive strength for the Hysol 9359.3 adhesive together with the transformed 

stresses in the scarf plane are presented in figure 12a. Results showed that the adhesive bond 

was subjected to normal compressive stress that varied significantly with the scarf angle. 

However, the shear stress was relatively constant in the range of 40–60 MPa. This suggested 

that the failure was governed by the shear stress component. In figure 12b, shear stresses of 

various adhesives are compared in case of static and dynamic loads. The results showed that the 

dynamic shear stresses at failure were significantly higher than those obtained under static 

loading. Furthermore, the Hysol adhesives offered higher dynamic shear strength compared to 

the Plexus adhesive. 

0

50

100

150

200

250

300

350

400

0 5 10 15 20 25 30 35 40 45 50

σX

F

`

σ1

F

`

σ2

F

`

τ

12

F

Hysol EA 9359.3

Scarf Angle, θ, deg.

Stress at Failure, σ

F, MPa.

0

20

40

60

80

100

120

140

160

180

200

0 5 10 15 20 25 30 35 40 45 50

-100

-80

-60

-40

-20

0

20

40

60

80

100

EA 9359.3

EA 9394

MA 425

DYNAMIC

QUASI

STATIC

Scarf Angle, θ, deg.

Quasi-Static Shear Stress at Failure, τ

F12, MPa.

Dynamic Shear Stress at Failure, τ

F12, MPa.

0

20

40

60

80

100

120

140

160

180

200

0 5 10 15 20 25 30 35 40 45 50

-100

-80

-60

-40

-20

0

20

40

60

80

100

EA 9359.3

EA 9394

MA 425

DYNAMIC

QUASI

STATIC

Scarf Angle, θ, deg.

Quasi-Static Shear Stress at Failure, τ

F12, MPa.

Dynamic Shear Stress at Failure, τ

F12, MPa.

(a) Stresses at failure (b) Shear stresses at failure

Figure 12. Stresses at failure in the rotated coordinate system of the scarf-repaired composite specimens. 

7. Conclusions 

The scarf patch repair scheme used in the present study is viable for level III repair of composite 

structural armor. The scheme uses practical scarf geometries for thick-section applications, 

requires only one-sided access for repair and vacuum consolidation, and provides rapid heating 

via induction to cure adhesives at elevated temperatures. The repair procedure was demonstrated 

using four adhesive systems. Simple test methods were proposed to apply realistic static and 

dynamic loads to composite backing plates with scarf repairs. 

The repair efficiency of scarf repairs, having scarf angles much greater than commonly used in 

aerospace applications, was assessed in a four-point bending test. The baseline behavior 

exhibited compression failure on the specimen surface followed by progressive failure. 

Scarf-repaired beams failed catastrophically in the adhesive bond, while the virgin specimens 

exhibited a progressive compressive failure of the fibers on the specimen surface layer. This 
13

difference in failure mode may indicate that the inherent energy-absorbing mechanisms of the 

composite were limited by the repair. 

The results, however, have shown that complete renewal of stiffness is achievable for the 

elevated temperature cure adhesives (a slight reduction of 15% is measured for the low modulus 

adhesive). The degree of strength recovered (based on first damage in the baseline) from the 

repairs increased from ~10% to 20%, to 40%, and 60% for the 45, 18.4, and 11.3° scarf repairs, 

respectively. 

In the dynamic experiments, scarf-repaired composite backing plates were subjected to 

compression loading using the SHPB. The dynamic axial strength for all adhesives was higher 

than the corresponding quasi-static data by a factor of 2-3 for each scarf angle. The dynamic 

axial strength increased as the scarf angle decreased, consistent with the four-point bend tests. 

The results showed that the dynamic shear strength in the scarf plane was also rate dependent 

and significantly greater than the static strengths. Higher rate impact testing is needed to fully 

characterize the strength and energy absorption capabilities of the scarf repair. 

Based on the limited results generated for the adhesives considered in the present study, 

induction curing of Hysol EA9359.3 with a 11.3° scarf offered the best combination of structural 

and rate-dependent properties. 
14

8. References 

1. Gama, B. A.; Bogetti, T. A.; Fink, B. K.; Yu, C. J.; Claar, T. D.; Eifert, H. H.; Gillespie 

J. W., Jr. Aluminum Foam Integral Armor: A New Dimension in Armor Design. 

Composite Structures 2001, 52, 381–395. 

2. Gama, B. A.; Gillespie, J. W., Jr.; Bogetti, T. A.; Fink, B. K. Innovative Design and 

Ballistic Performance of Lightweight Composite Integral Armor. Technology for the U.S. 

Army’s Transformation, SAE World Congress and Exposition; paper no. 2001-01-0888; 

Detroit, MI, 5–8 March 2001. 

3. Gama, B. A.; Gillespie, J. W., Jr. Mahfuz, H.; Bogetti, T. A.; Fink, B. K. Effect of 

Non-Linear Material Behavior on the Through-Thickness Stress Wave Propagation in 

Multi-Layer Hybrid Lightweight Armor. International Conference on Computational 

Engineering and Sciences, Los Angeles, CA, 21–25 August 2000; Atluri, S. N., Burst, F. W., 

Eds.; Tech Science Press: Palmdale, CA, Vol. I, pp 157–162. 

4. Monib, A. M.; Gillespie, J. W., Jr. Damage Tolerance of Composite Laminates Subjected to 

Ballistic Impact. Proceedings of ANTEC 98, Society of Plastics Engineers, Brookfield, CT, 

1998; pp 1463–1467. 

5. Mahdi, S.; Gama, B. A.; Yarlagadda, S.; Gillespie, J. W., Jr. Effect of the Manufacturing 

Process on the Interfacial Properties and Structural Performance of Multi-Functional 

Composite Structures. Composites Part A 2003, 34, 635–647. 

6. Huang, X. G.; Gillespie, J. W., Jr.; Kumar, V.; Gavin, L. Mechanics of Integral Armor: 

Discontinuous Ceramic-Cored Sandwich Structure Under Tension and Shear. Composite 

Structures 1996, 36, 81–90. 

7. Davila, C. G.; Chen, T-K. Advanced Modeling Strategies for the Analysis of 

Tile-Reinforced Composite Armor. Applied Composite Materials 2000, 7 (1), 51–68. 

8. Gama, B. A.; Bogetti, T. A.; Fink, B. K.; Gillespie, J. W., Jr. Processing, Ballistic Testing 

and Repair of Composite Integral Armor, Proceedings of the 32nd International SAMPE 

Technical Conference, Boston, MA, 5–9 November 2000. 

9. Gillespie, J. W., Jr.; Monib, A. M.; Carlsson, L. A. Damage Tolerance of Thick-Section S-2 

Glass Fabric Composites Subjected to Ballistic Impact Loading. Journal of Composite 

Materials January 2003. 
15

10. Gama, B. A.; Yarlagadda, S.; Bogetti, T. A.; Fink, B. K.; Gillespie, J. W., Jr. Repair of 

Composite Integral Armor. Proceedings of the SAMPE 2001 Symposium and Exhibition, 

Materials and Processes, Long Beach, CA, 6–10 May 2001. 

11. Mahdi, S.; Gillespie, J. W., Jr. Finite Element Analysis of Tile-Reinforced Composite 

Structural Armor Subjected to Bending Loads. Composites B: Engineering May 2003. 

12. ASM Handbook, Volume 8, Mechanical Testing and Evaluation; ASM International: 

Materials Park, OH, 2000. 

13. Loctite Aerospace. Dexter Hysol Paste Adhesive and Specialty Resins Selector Guide. 

www.loctiteaero.com (accessed 2001). 

14. Mahdi, S.; Kim, H.-J.; Gama, B. A.; Yarlagadda, S.; Gillespie, J. W., Jr. A Comparison of 

Oven-Cured and Induction-Cured Adhesively Bonded Composite Joints. Journal of 

Composite Materials 2003, 37 (6), 519–542. 

15. Ravichandran, G.; Subhash, G. Critical Appraisal of Limiting Strain Rates for Compression 

Testing of Ceramics in a Split-Hopkinson Pressure Bar. J. Am. Ceram. Soc. 1994, 77, 

263–267. 
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 E MUNLEY 

 6300 GEORGETOWN PIKE 

 MCLEAN VA 22101 

 1 USDOT FEDERAL RAILROAD 

 M FATEH RDV 31 

 WASHINGTON DC 20590 

 3 CYTEC FIBERITE 

 R DUNNE 

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 R MAYHEW 

 1300 REVOLUTION ST 

 HAVRE DE GRACE MD 21078 

 1 DIRECTOR 

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 IANG TMT 

 2055 BOULDERS RD 

 CHARLOTTESVILLE VA 

 22911-8318 

 1 SIOUX MFG 

 B KRIEL 

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 FT TOTTEN ND 58335 

 2 3TEX CORP 

 A BOGDANOVICH 

 J SINGLETARY 

 109 MACKENAN DR 

 CARY NC 27511 

 1 3M CORP 

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 3M CENTER BLDG 60 IN 01 

 ST PAUL MN 55144-1000 
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 1 DIRECTOR 

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 TA 5 

 K CRELLING 

 WASHINGTON DC 20310 

 1 ADVANCED GLASS FIBER YARNS 

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 281 SPRING RUN LANE STE A 

 DOWNINGTON PA 19335 

 1 COMPOSITE MATERIALS INC 

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 19105 63 AVE NE 

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 1 JPS GLASS 

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 1 COMPOSITE MATERIALS INC 

 R HOLLAND 

 11 JEWEL CT 

 ORINDA CA 94563 

 1 COMPOSITE MATERIALS INC 

 C RILEY 

 14530 S ANSON AVE 

 SANTA FE SPRINGS CA 90670 

 2 SIMULA 

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 R HUYETT 

 10016 S 51ST ST 

 PHOENIX AZ 85044 

 2 PROTECTION MATERIALS INC 

 M MILLER 

 F CRILLEY 

 14000 NW 58 CT 

 MIAMI LAKES FL 33014 

 2 FOSTER MILLER 

 M ROYLANCE 

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 195 BEAR HILL RD 

 WALTHAM MA 02354-1196 

 1 ROM DEVELOPMENT CORP 

 R O MEARA 

 136 SWINEBURNE ROW 

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 NEWPORT RI 02840 

 2 TEXTRON SYSTEMS 

 T FOLTZ 

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 1449 MIDDLESEX ST 

 LOWELL MA 01851 

 1 O GARA HESS & EISENHARDT 

 M GILLESPIE 

 9113 LESAINT DR 

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 2 MILLIKEN RESEARCH CORP 

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 1 CONNEAUGHT INDUSTRIES INC 

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 1 ARMTEC DEFENSE PRODUCTS 

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 COACHELLA CA 92236 

 1 NATL COMPOSITE CTR 

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 2000 COMPOSITE DR 

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 3 PACIFIC NORTHWEST LAB 

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 1 SAIC 

 M PALMER 

 1410 SPRING HILL RD STE 400 

 MS SH4 5 

 MCLEAN VA 22102 
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 1 ALLIANT TECHSYSTEMS INC 

 4700 NATHAN LN N 

 PLYMOUTH MN 55442-2512 

 1 APPLIED COMPOSITES 

 W GRISCH 

 333 NORTH SIXTH ST 

 ST CHARLES IL 60174 

 1 CUSTOM ANALYTICAL 

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 A ALEXANDER 

 13000 TENSOR LANE NE 

 FLINTSTONE MD 21530 

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 1745 JEFFERSON DAVIS HWY 

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 3 ALLIANT TECHSYSTEMS INC 

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 PO BOX 210 

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 26726-0210 

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 1 PRATT & WHITNEY 

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 400 MAIN ST MS 114 37 

 EAST HARTFORD CT 06108 

 5 NORTHROP GRUMMAN 

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 1 BRIGS COMPANY 

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 2668 PETERBOROUGH ST 

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 L ZERNOW 

 425 W BONITA AVE STE 208 

 SAN DIMAS CA 91773 

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 2 GENERAL DYNAMICS OTS 

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 1 GKN WESTLAND AEROSPACE 

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 5 SIKORSKY AIRCRAFT 

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 1 AEROSPACE CORP 

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 1 UDLP WARREN OFC 

 A LEE 

 31201 CHICAGO RD SOUTH 

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 P JANKE MS 170 

 4800 EAST RIVER RD 

 MINNEAPOLIS MN 55421-1498 

 1 LOCKHEED MARTIN 

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 D FORTNEY 

 1011 LOCKHEED WAY 

 PALMDALE CA 93599-2502 

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 R FIELDS 

 5537 PGA BLVD 

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 & SYSTEMS DIV 

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 1745A W NURSERY RD 

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 991 ANNAPOLIS WAY 

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 6 INST FOR ADVANCED 

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 I MCNAB 

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 C PERSAD 

 3925 W BRAKER LN STE 400 

 AUSTIN TX 78759-5316 

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 1233 SHELBURNE RD STE D8 

 SOUTH BURLINGTON VT 

 05403-7700 

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 409 W CATHERINE ST 

 BEL AIR MD 21014-3613 
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 1 SAIC 

 G CHRYSSOMALLIS 

 8500 NORMANDALE LAKE BLVD 

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 BLOOMINGTON MN 55437-3828 

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 H T HAHN 

 LOS ANGELES CA 90024-1597 

 2 UNIV OF DAYTON 

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 300 COLLEGE PARK AVE 

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 N SCHOTT 

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 LOWELL MA 01854 

 1 IIT RESEARCH CTR 

 D ROSE 

 201 MILL ST 

 ROME NY 13440-6916 

 1 GA TECH RESEARCH INST 

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 P FRIEDERICH 

 ATLANTA GA 30392 

 1 MICHIGAN ST UNIV 

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 212 EARTH ENGR 

 SCIENCES BLDG 

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 1 STANFORD UNIV 

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 & AEROBALLISTICS 

 S TSAI 

 DURANT BLDG 

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 1 UNIV OF MAINE 

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 R LOPEZ ANIDO 

 5793 AEWC BLDG 

 ORONO ME 04469-5793 

 1 JOHNS HOPKINS UNIV 

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 P WIENHOLD 

 11100 JOHNS HOPKINS RD 

 LAUREL MD 20723-6099 

 1 UNIV OF DAYTON 

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 5 UNIV OF DELAWARE 

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 J GILLESPIE 

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 D HEIDER 

 201 SPENCER LAB 

 NEWARK DE 19716 
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 1 DEPT OF MTRLS 

 SCIENCE & ENGRG 

 UNIV OF ILLINOIS 

 AT URBANA CHAMPAIGN 

 J ECONOMY 

 1304 WEST GREEN ST 115B 

 URBANA IL 61801 

 1 UNIV OF MARYLAND 

 DEPT OF AEROSPACE ENGRG 

 A J VIZZINI 

 COLLEGE PARK MD 20742 

 1 DREXEL UNIV 

 A S D WANG 

 3141 CHESTNUT ST 

 PHILADELPHIA PA 19104 

 3 UNIV OF TEXAS AT AUSTIN 

 CTR FOR ELECTROMECHANICS 

 J PRICE 

 A WALLS 

 J KITZMILLER 

 10100 BURNET RD 

 AUSTIN TX 78758-4497 

 3 VA POLYTECHNICAL 

 INST & STATE UNIV 

 DEPT OF ESM 

 M W HYER 

 K REIFSNIDER 

 R JONES 

 BLACKSBURG VA 24061-0219 

 1 SOUTHWEST RESEARCH INST 

 ENGR & MATL SCIENCES DIV 

 J RIEGEL 

 6220 CULEBRA RD 

 PO DRAWER 28510 

 SAN ANTONIO TX 78228-0510 

 1 BATELLE NATICK OPERS 

 B HALPIN 

 313 SPEEN ST 

 NATICK MA 01760 

 3 DIRECTOR 

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 A FRYDMAN 

 2800 POWDER MILL RD 

 ADELPHI MD 20783-1197 

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 W C FRAZER 

 400 COLLERAN RD 

 APG MD 21005-5059 

 91 DIR USARL 

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 AMSRD ARL WM BA 

 D LYON 

 AMSRD ARL WM BC 

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 P PLOSTINS 

 A ZIELINSKI 

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 B FORCH 

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 R LIEB 

 R PESCE RODRIGUEZ 

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 AMSRD ARL WM BF 

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 AMSRD ARL WM MB 

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 T BOGETTI 

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 R DOWDING 

 W DRYSDALE 
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 R EMERSON 

 D HENRY 

 D HOPKINS 

 R KASTE 

 L KECSKES 

 M MINNICINO 

 B POWERS 

 D SNOHA 

 J SOUTH 

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 AMSRD ARL WM MC 

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 R BOSSOLI 

 E CHIN 

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 D GRANVILLE 

 B HART 

 J LASALVIA 

 J MONTGOMERY 

 F PIERCE 

 E RIGAS 

 W SPURGEON 

 AMSRD ARL WM MD 

 B CHEESEMAN 

 P DEHMER 

 R DOOLEY 

 G GAZONAS 

 S GHIORSE 

 C HOPPEL 

 M KLUSEWITZ 

 W ROY 

 J SANDS 

 D SPAGNUOLO 

 S WALSH 

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 AMSRD ARL WM RP 

 J BORNSTEIN 

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 A NIILER 

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 1 LTD 

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 UK 

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 316 ROUTE DE BAYONNE 

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