J.A. Salem

Glenn Research Center, Cleveland, Ohio

J.L. Bail

Ohio Aerospace Institute, Brook Park, Ohio

N.G. Wilmoth

Vantage Partners, LLC, Brook Park, Ohio

L.J. Ghosn, L.W. Kohlman, G.D. Roberts, and R.E. Martin

Glenn Research Center, Cleveland, Ohio

Burst Testing of Triaxial Braided Composite Tubes

NASA/TM—2014-216615

January 2014
NASA STI Program . . . in Profi le

Since its founding, NASA has been dedicated to the 

advancement of aeronautics and space science. The 

NASA Scientifi c and Technical Information (STI) 

program plays a key part in helping NASA maintain 

this important role.

The NASA STI Program operates under the auspices 

of the Agency Chief Information Offi cer. It collects, 

organizes, provides for archiving, and disseminates 

NASA’s STI. The NASA STI program provides access 

to the NASA Aeronautics and Space Database and 

its public interface, the NASA Technical Reports 

Server, thus providing one of the largest collections 

of aeronautical and space science STI in the world. 

Results are published in both non-NASA channels 

and by NASA in the NASA STI Report Series, which 

includes the following report types:

• TECHNICAL PUBLICATION. Reports of 

completed research or a major signifi cant phase 

of research that present the results of NASA 

programs and include extensive data or theoretical 

analysis. Includes compilations of signifi cant 

scientifi c and technical data and information 

deemed to be of continuing reference value. 

NASA counterpart of peer-reviewed formal 

professional papers but has less stringent 

limitations on manuscript length and extent of 

graphic presentations.

• TECHNICAL MEMORANDUM. Scientifi c 

and technical fi ndings that are preliminary or 

of specialized interest, e.g., quick release 

reports, working papers, and bibliographies that 

contain minimal annotation. Does not contain 

extensive analysis.

• CONTRACTOR REPORT. Scientifi c and 

technical fi ndings by NASA-sponsored 

contractors and grantees.

• CONFERENCE PUBLICATION. Collected 

papers from scientifi c and technical 

conferences, symposia, seminars, or other 

meetings sponsored or cosponsored by NASA.

• SPECIAL PUBLICATION. Scientifi c, 

technical, or historical information from 

NASA programs, projects, and missions, often 

concerned with subjects having substantial 

public interest.

• TECHNICAL TRANSLATION. Englishlanguage translations of foreign scientifi c and 

technical material pertinent to NASA’s mission.

Specialized services also include creating custom 

thesauri, building customized databases, organizing 

and publishing research results.

For more information about the NASA STI 

program, see the following:

• Access the NASA STI program home page at 

http://www.sti.nasa.gov

• E-mail your question to help@sti.nasa.gov

• Fax your question to the NASA STI 

Information Desk at 443–757–5803

• Phone the NASA STI Information Desk at

 443–757–5802

• Write to:

 STI Information Desk

NASA Center for AeroSpace Information

 7115 Standard Drive

 Hanover, MD 21076–1320
J.A. Salem

Glenn Research Center, Cleveland, Ohio

J.L. Bail

Ohio Aerospace Institute, Brook Park, Ohio

N.G. Wilmoth

Vantage Partners, LLC, Brook Park, Ohio

L.J. Ghosn, L.W. Kohlman, G.D. Roberts, and R.E. Martin

Glenn Research Center, Cleveland, Ohio

Burst Testing of Triaxial Braided Composite Tubes

NASA/TM—2014-216615

January 2014

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135
Acknowledgments

Thanks to Chris Burke for running tube tests.

Available from

NASA Center for Aerospace Information

7115 Standard Drive

Hanover, MD 21076–1320

National Technical Information Service

5301 Shawnee Road

Alexandria, VA 22312

Available electronically at http://www.sti.nasa.gov

Trade names and trademarks are used in this report for identifi cation 

only. Their usage does not constitute an offi cial endorsement, 

either expressed or implied, by the National Aeronautics and 

Space Administration.

Level of Review: This material has been technically reviewed by technical management. 
NASA/TM—2014-216615 1 

Burst Testing of Triaxial Braided Composite Tubes 

J.A. Salem 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

J.L. Bail 

Ohio Aerospace Institute 

Brook Park, Ohio 44142 

N.G. Wilmoth 

Vantage Partners, LLC 

Brook Park, Ohio 44142 

L.J. Ghosn, L.W. Kohlman, G.D. Roberts, and R.E. Martin 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Abstract 

Applications using triaxial braided composites are limited by the materials transverse strength which 

is determined by the delamination capacity of unconstrained, free-edge tows. However, structural 

applications such as cylindrical tubes can be designed to minimize free edge effects and thus the strength 

with and without edge stresses is relevant to the design process. The transverse strength of triaxial braided 

composites without edge effects was determined by internally pressurizing tubes. In the absence of edge 

effects, the axial and transverse strength were comparable. In addition, notched specimens, which 

minimize the effect of unconstrained tow ends, were tested in a variety of geometries. Although the 

commonly tested notch geometries exhibited similar axial and transverse net section failure strength, 

significant dependence on notch configuration was observed. In the absence of unconstrained tows, 

failure ensues as a result of bias tow rotation, splitting, and fracture at cross-over regions. 

Introduction 

Triaxial braided composite employing large-tow-size yarns offer a good combination of cost and 

performance for many aerospace applications. One factor limiting the performance of such composites is 

the effect of free tow-ends: sections with terminated tow-ends subjected to tensile loads exhibit low 

strength relative to sections with constrained tow ends. The failure mechanisms have been studied by 

Littell et al. (Ref. 1), and imply tow splitting and delamination at edges subjected to load. 

Because tow-end stresses and related edge effects can be minimized in structures such as tubes and 

spheres, it is of interest to measure the transverse strength both with and without free-end effects. 

Standard straight-sided uniaxial test specimens (e.g., ASTM D3039) exhibit these edge effects and do not 

represent the potential of the material under certain design condition because the load resolved on the 

unconstrained edge tows causes damage to propagate along the tow. On the other hand, it will be shown 

in this work that failure progression in pressurized tubes, as observed by interruption of pressure tests, 

occurs by delamination at bias tow edges via a scissoring action. Delamination is followed by 

longitudinal splitting and then transverse breakage of split tows where the bias tows cross the axial tows. 

The low strength of transverse tensile specimens is thus a structural effect. 
NASA/TM—2014-216615 2 

The strength without edge effects can be measured by employing partially pressurized tubes because 

the overhung regions are subjected to substantially lower hoop stresses than the highly loaded central 

region. The length of overhang can be used to interrogate the sensitivity to edge stress and the transition 

from internal failure to edge failure. Another advantage of a tubular geometry is its geometric similarity 

to applications such as aircraft engine cases, which are critical to efficiency and safety, and are locally 

loaded during rotating component failure. 

Besides tubes, uniaxial specimens can be designed to either restrain free-end-tows or shield restrained 

tows such that strength reduction is mitigated. Examples of such specimens are the bowtie and edgenotched sheet specimens that restrain the tows passing through the ligament. Transverse tests with these 

specimens indicate similar net section ultimate strengths to D3039 axial values for T700S/E-862 carbonepoxy composite (Refs. 2 and 3). FEA of notched tensile specimens subjected to fixed grip ends implies 

that the free edges, where unconstrained tows exist, are shielded from significant load via the notch. 

However, the notch does create a stress concentration and audible damage occurs at relatively low load 

levels. Thus the effect of notch geometry on resultant tensile strength was examined in this work as a 

function of notch angle and depth. Although tubes and notched uniaxial specimens can be compared, 

complications occur because of the different defects caused by the more complex manufacturing process 

required for tubes as compared to that used for flat plates. These effects were partially mitigated by visual 

inspection of the tubes. 

The purpose of this work was to apply tube testing methods generally employed for testing monolithic 

ceramics, ceramic matrix composites and graphite to polymer composite materials and thereby measure the 

strength in the absence of terminated tows. For comparison, notched specimens were also tested. 

Tubes can be pressurized by bladders or with fluids such as oil (Refs. 4 to 7), however, fluids are 

messy and hard to contain under large displacements and the high pressures and displacements required 

for polymer composites make the use of bladders more difficult. Using a polymer insert as a 

pressurization medium has been used previously with relatively stiff materials such as ceramics (Ref. 8), 

however, application to relatively compliant strong materials such as polymer composites present the 

additional complication of large displacements, high pressures, and extrusion of the insert. 

Mosley has done substantial investigation into tube testing of ceramics and put forth detailed 

recommendations (Ref. 8) for infinite length tubes. As composite materials are expensive to manufacture, 

a short tube that results in sufficiently low edge stress relative to the maximum stress is desirable. Reuter 

and Guess (Ref. 7) derived an elegant criterion for the minimum overhang to eliminate edge effects in 

short tubes and demonstrated strength improvements in the absence of edge effects. Besides being short, a 

uniform wall stress with low second and third principal stresses is needed as the transverse uniaxial 

tensile strength is of interest. Further, simple strength of materials solutions should be reasonably 

descriptive. 

Results for tubes and multiple notch geometries are presented with the goal of understanding failure 

behavior and the limits of simple notched coupon tests. 

Test Method Verification 

Finite Element Analysis 

Because of the large displacements associated with high strength, low stiffness materials such as 

polymer composites (Sf = 120 ksi, E = 7106

 psi), containment of the polymer plug was of concern. Thus 

tube pressurization in combination with and without edge constraint was considered. Finite element 

analysis (FEA) results are shown in Figures 1 and 2. Edge constraint proved difficult due to the sharp 

transition between constrained and unconstrained regions. Thus simple pressurization over a short region 

(~1/3rd) via a low durometer insert was pursued. For this case, a peak hoop stress occurs about the 

centerline and is accompanied by a small axial stress. For practical purposes, all stresses at the free edges 

are zero, as shown in Figure 2. 
NASA/TM—2014-216615 3 

The size of the pressure zone is critical to the stress distribution through the wall and the ratio of axial 

to hoop stress. As shown in Figure 3 for 6 in. isotropic tubes, a zone of 1.875 in. produces uniform stress 

through the wall and separates the axial peaks from the hoop peaks making it a quasi-uniaxial test. Longer 

pressurized zones approach the strength-of-materials solution, with the peak stress on the inner wall. 

However, the peak axial and hoop stresses coincide. Thus a 1.88 to 2 in. zone was used in experiments. 

For notched tensile specimens subjected to fixed grip ends, strain concentrations are exhibited, as shown 

in Figure 4. However, the concentration in a composite is strongly associated with the tows terminated at the 

notch root. All gripped tows exhibit a very uniform strain and the terminated tows at specimen edges are 

practically strain free. Thus the load capacity of the tows crossing the ligament are employed. 

Although the stress concentration is significant, such effects in composites are less governed by SaintVenant’s principle and tend to be transmitted over larger areas. Thus the concentration might be mitigated 

via fiber architecture and end restraints. 

Experimental Stress Analysis 

High strength aluminum, which has a similar elastic modulus to the composite, was used for 

experimental verification of tube testing. Pressurization was achieved by compressing a polymer plug 

between platens. Strain was monitored by using a photogrammetry system and associated software 

(Trilion Quality Systems). The elastic constants of the tube material were determined by compressing the 

tube while monitoring the surface displacements via the camera system. 

Pressurization of the tube resulted in excellent agreement with theory for an infinite tube, as shown in 

Figure 5. The bending induced axial stress at the mid-plane is ~12 percent of the maximum hoop stress. 

However, by the time the hoop stress decays significantly (~5 percent); the axial stress has increased to 

~25 percent of the hoop. Observations of failure locations in composite tubes, discussed in the following 

sections, indicated mid-plane failure, implying that the state is sufficiently uniaxial for engineering 

purposes. 

Test Materials and Experimental Procedures 

The composites were made with high strength, standard modulus TORAYCA T700S fiber (Toray 

Carbon Fibers America, Inc.) and EPIKOTE Resin 862/EPIKURE with Curing Agent W (Resolution 

Performance Products) or 5208 one part resin. Fiber and resin properties reported by the fiber 

manufacturers are shown in Table 1. The 2D triaxial braid preforms were made by A&P Technology 

(A&P Technology), and composite panels and tubes were fabricated by resin transfer molding (RTM) at 

North Coast Composites (North Coast Composites). 

A representative picture of the 0/+60/–60 triaxial braid architecture unit cell is shown in Figure 6. The 

±60° bias fibers are visible on the surface, and portions of the 0° axial fibers that lie below the ±60° bias 

fibers can be seen in open spaces. The 0° axial fibers were 24k flattened tows while the ±60° bias fibers 

were 12k flattened tows. Although larger fiber bundles were used in the axial direction, the fiber bundle 

spacing in the axial and bias directions were adjusted to give the same fiber volume in the axial and bias 

directions. This fiber architecture is quasi-isotropic in-plane, so the in-plane stiffness is expected to be the 

same in all directions when the region of measurement includes several unit cells. 

Composite plates were fabricated by placing 12 layers of the 0/+60/–60 braid preform into the RTM 

mold with the 0° fibers aligned in the same direction. Although the axial (0°) fibers in the various layers 

were parallel, the lateral position in the six layers was not controlled, but typically exhibited shifting often 

referred to as nesting. The braided composite tubes were manufactured by over-braiding six layers of 

triaxial braid on a mandrel, followed by RTM. The fiber volume fraction, braid angle, tow selection, and 

cure conditions were chosen to match, as closely as possible, the parameters of the flat panel material. 

The lateral position of the axial tows in the tubes were clocked slightly, creating nesting and improving 

consolidation. Also, the preforming methods and molding procedures were adjusted as successive tubes 
NASA/TM—2014-216615 4 

were fabricated in order to reduce defects and optimize the process. As a result, some of the tubes had 

known defects (mainly local distortion of fiber tows) prior to testing. 

Resin was injected into the closed mold and cured according to processing conditions recommended 

by the resin manufacturer. The E862 was cured for 2 1/2 hr at 177 °C (350 °F) and the 5208 was cured for 

2 hr at 177 °C (350 °F). Cured panel dimensions (after trimming) were 2 ft wide by 2 ft long by 0.25 in. 

thick. Cured tube dimensions were 24 in. long with 4.17 in. outer diameter (OD) and 3.91 in. inner 

diameter (ID). Some of the tubes were observed to have defects (mainly distortion of fiber tows) prior to 

testing. 

Uniaxial tensile specimens were cut from the panels by using an abrasive water jet followed by fine 

grinding. Straight-sided specimens were prepared and tested in accordance with ASTM International 

D3039. The transverse and axial notch configurations are summarized by Figure 7. Relative notch depths 

of 2a/W = 0.5 and 0.8 were used with notch angles between 0 to 120 in order to understand notch 

sensitivity. A relative depth of 0.8 leaves sufficient material for one unit cell. Uniaxial tests were 

performed at a stroke rate of 0.050 in./min which typically gave a strain rate of 1 to 1.510–4/s. Typically, 

five or more specimens were tested per condition. For tubes, a nominal pressure rate 68 psig/s was 

applied to produce a hoop strain rate of 1.510–4/s. 

Photogrammetry (digital image correlation) was used to monitor both global and local strain fields by 

choosing a gage length larger than the local architecture or within a region of interest. 

Stresses in a Finite Tube 

The nominal tensile stress in the central region of an infinite isotopic tube can be estimated from 















	


  
  2

2

2 2

2

1

r

R

R R

R p o

o i

i i

t (1)















	   
  2

2

2 2

2

1

r

R

R R

R p o

o i

i i r (2)















	


  
  2

2

2 2

2

max 1

i

o

o i

i i

R

R

R R

R p (3)

where pi is the applied internal pressure, Ri is inner radius, Ro is the outer radius, and r is the radial 

position of interest. Based on Equations (1) to (3), the maximum principal stress is tangential (hoop) and 

occurs on the ID. This is not the exact case for a partially pressurized tube as the free, overhung ends 

constrain expansion of the central region and thereby cause axial bending stresses. Also, small frictional 

effects may occur. Despite these effects, the full elastic solutions considering the constraint (Ref. 8) were 

within 4 percent of strength-of-materials (SOM) values (Eq. (3)). 

Results 

Uniaxial Strength 

Axial and transverse, macro-level stress-strain curves are shown in Figure 8 for T700S/E862 and / 

5208. For transverse loading, nonlinear extension occurs at ~0.5 percent strain and ~25 percent of the 

ultimate axial strength. A linear regression fit range of 0.5 percent, which was larger than that used 

previously (0.2 percent, (Ref. 1)), provided more consistency in Poisson’s ratio and met reciprocity 

requirements (12/E11 = 21/E22) with E11 = 6.95103

 ksi, 12 = 0.34, E22 = 6.03103 ksi, and 21 = 0.30, as 

shown in Table 2. 
NASA/TM—2014-216615 5 

T700S/5208 exhibited less nonlinear extension prior to failure and lower fracture strength. The effects 

of notch angle and depth are shown in Figure 9, and imply that net section strength decreases with 

increasing notch angle for transverse and axial orientations. For transverse orientation, large angles cut 

bias tows thereby resulting in less length of bias tows through which load can be transmitted, and less 

shielding, via surrounding material, of tows that do carry load through the ligament. Interestingly, shallow 

notches exhibit lower strength than deep notches again implying that additional material to the side of a 

tow shields the loaded tow: unsupported tows running along a free edge (e.g., 60 notch) are weaker. 

Axial specimens are most sensitive to notches, with deep notches exhibiting the greatest net section 

strength, implying significant edge effects even in well machined, unnotched sections. Notching improves 

axial strength by shedding load to bias tows crossing the ligament and mitigating machining damage to 

the load bearing tows. Further, all load bearing tows are supported by surrounding material, thereby 

maximizing load redistribution. The notched axial values converge to somewhat below the smooth section 

value, implying a small stress concentration effect once edge effects are mitigated. 

The notched specimen is simple to fabricate and forces fiber tensile failure for transverse specimens 

by constraining free-end tows subjected to load. Although notched specimens have a non-uniform gage 

region and resultant stress concentrations, notched coupons produced higher strength values compared to 

the straight-sided coupon. This is due to the structural nature of the composite and resultant notch 

insensitivity. 

Noteworthy for notched specimens was early, audible popping during loading and nonlinearity at 

about 60 ksi, implying that the notch does influence damage initiation. However, propagation and 

ultimate failures were limited by tow restraint: all bias fibers passing through the ligament are restrained 

within the grip and carry load. For axial specimens, failure begins with separation of the ears from the 

ligament by transverse splitting of bias tows along the interface between notched and un-notched sections, 

as shown in Figure 10(a). As the notch angle is increased, the maximum axial strains shift from the interface 

region to the bias tows intersecting the edges, weakening the section. For transverse specimens, failure 

begins with high strains in bias tow intersecting the notch root. The strains gradually migrate inward along 

the ligament centerline until distinct jumps in bias tow strain levels (~2 percent) occur due to delamination. 

Large, horizontal bands of high strain can be seen in some axial tows, as shown in Figure 10(c). 

For large notch angle (120), bias tows along the notch edges also exhibit large strains, weakening the 

specimens, as shown in Figure 10(d). For deep notches, highly strained bias tows can be seen to extent to 

the grips. 

Noteworthy for unnotched transverse specimens is the strain banding shown in Figure 11. These local 

strain concentrations and the applied stress were used to estimate local, effective elastic Young’s modulus 

and Poisson’s ratio. These in situ measurements gave E = 7.74103

 ksi and = 0.41 for the low strain 

regions between 0 tows and E = 5.11103

 ksi with = 0.25 in the high strain regions corresponding to 

tows. This is ~±10 percent from the macro-level modulus values. Non-linear strains are exhibited in high 

strain regions at about two-thirds the applied stress required to create non-linearity in the low strain 

regions, and the equality of transverse strains implies good compatibility between regions, Figure 11(c). 

For the axial tests, banding is less apparent and the macro values are representative. 

The low strength in transverse loading has been described (Refs. 1 and 3) as the result of tow 

splitting, outward delamination, and pull-out of a terminated tow. The delamination and splitting occurs 

because the resolved tow loads are only supported by surrounding epoxy. Besides notching, which gives 

acreage for load shedding, continuity of the tows, as occurs in tubes, can be used to mitigate this 

structural affect. 

Tubular Strength 

Tube burst strength is compared to straight sided, bowtie, and notched section tensile strength in 

Table 3 for T700S/E862. Partial pressurization of tubes and notching of tensile specimens results in 

comparable transverse strength estimates (118 vs. 119 ksi). Some difficulty was encountered in failing 

several tubes. The loads were sufficient to partially extrude the plug between the platen and tube wall, 
NASA/TM—2014-216615 6 

implying that the plug was too soft. These tests were not included in reported averages, but the loads 

implied potential strength greater than 140 ksi. 

The axially woven tubes failed from the center where maximum hoop stress occurred, as shown in 

Figure 12, with an explosive report. The transverse wrapped tubes failed by delamination at the braid end 

rather than by fiber failure. This resulted in low and more variable strengths. Stress-strain curves for 

tubular and transverse uniaxial specimens are compared in Figure 13. The tubes exhibit linear behavior 

to near maximum stress and give little indication of damage until stresses are comparable to the uniaxial 

strength in the axial direction. Ultimate strains are ~1.8 percent which is near that of the yarn 

(~2.1 percent). Most interesting is the tendency of the tubes to stiffen just after the failure stress exhibited 

by straight-sided uniaxial transverse tests (dotted horizontal lines in Fig. 13). This presumably occurs 

from locking of the braid and engagement of the bias fibers. 

Significant strength variation between specimens taken from different tubes (90 to 125 ksi) was 

exhibited and most failures occurred within one tow repeat of a mold line. The variation was traced 

predominantly to variation the quality of the braid architecture and section thickness near the mold-lines. 

Two sources of low strength were identified: fiber distortion or curved fibers at the mold lines and 

damaged or cut fibers. Figure 14(a) shows the stress—strain curve for a specimen with cut fibers. The 

tube exhibited low compliance and strain relaxation at the central gage region as the specimen 

delaminated between the cut and uncut lamina shown in Figure 14(b). Failure occurred at only 90 ksi. 

Figure 14(c) shows a specimen with distorted fibers on one side of a mold-line. Although the initial 

compliance is nearly that of well molded specimens, large strains were exhibited at the mold-line as the 

fibers straightened. The large mold line strains resulted in strain relaxation between the mold lines where 

the optical strain gage was placed. Failure occurred just below the mold line where the fibers were distorted, 

Figure 14(c). The strain enhancement associated with distorted fibers can be seen in detail for another tube 

in Figure 15. The values in Tables 3 and 4 represent measurements from tubes of good visual quality. The 

strongest tube (~135 ksi) was comparable to the axial strength measured in the presence of a deep notch. 

When axial load is applied parallel to a curved tow, it rotates and aligns to the applied load direction 

causing local deformation and weakness. Although properties might be improved by minimizing braid 

distortions and gaps between tows, regions exhibiting the most apparent distortions did not always exhibit 

failure. Failure may be controlled by distortion occurring through multiple layers. 

Due to the coarse architecture and associated stiffness variations, the axial overbraided tubes exhibit 

substantial strain variation in the hoop direction for any axial position, as shown in Figure 16, with high 

strains corresponding to axial tows, and overall agreement with Figure 3. As pressure increases, specific 

axial tows exhibit large surface strains (~5 percent) that persist until failure. At relatively high pressures, 

regions initially exhibiting low hoop strain in axial tows develop significant hoop strain (~4 percent) in 

the pattern of the bias tows. Thus the choice of location and size of the strain gage can affect 

interpretation of results. In order to estimate the stress and strain distributions and variations during the 

failure process, strain measurements were averaged for multiple “hot” and “cold” regions. 

As shown in Figure 17, substantial nonlinear extension occurs prior to failure in some regions, and 

the strains at axial tows are more than double those between tows, Figure 17(a). Once the axial tows 

exhibit non-linear strain, which occurs simultaneously in hoop and longitudinal directions at ~2 percent 

hoop strain, the bias tows shed hoop strain while increasing longitudinal strain, Figure 17(b). Thus the 

axial tows shed load to the bias tows as failure ensues and the system globally stiffens. The bias tows in a 

tube can take extra load because they are contiguous. Figures 8 and 17 imply that the ultimate strength 

and the onset of damage are significantly improved relative to transverse uniaxial loading, with the onset 

of nonlinear behavior occurring at ~75 percent of ultimate for T700S/E862. The global stiffening might 

be used as an indicator of localized damage. 
NASA/TM—2014-216615 7 

Tube data for T700S/5208 is summarized in Table 4. Figure 9 implies that deep notches (2a/W = 0.8 

with  = 0 to 60) give the most comparable results to tubes for the current architecture. Clearly, notched 

specimens show systematic dependence on notch depth and angle for both transverse and axial 

orientations. It is noteworthy that both the strongest notched specimens and the tubes that presented 

testing difficulty due to plug extrusion exhibited nominal stresses of 140 ksi, implying that tube strength 

can be improved. More study of notched geometry effects is needed with various epoxies to give specific 

recommendations. 

Failure Mode 

Observations of partially failed tubes indicated that as axial tows subject to transverse loads relax, the 

tube stiffens by engagement of the bias tows, Figure 18. Failure of tubes ensues as the bias tows rotate 

parallel to the hoop stress and delaminate along tow edges, Figure 18(a). Delamination and further 

loading leads to impingement of the fibers at cross-over regions and a scissoring action that splits the 

tows, Figure 18(b). Eventually, tow breakage occurs where bias tows cross over axial tows, Figure 18(c), 

resulting in expulsion of a single axial tow as shown in Figure 12(a). 

Thus the increased strength exhibited by notched specimens is a structural effect of more load being 

transmitted along bias tows, elimination of edge damage, and sheltering of tows carrying load. Even a 

well prepared axial specimen with free edges will be weaker because of the unrestrained tows and low 

delamination strength. Making the tows contiguous and retaining free material to the side of bias tows, as 

in a tube or deeply notched specimen, improves structural integrity by employing the load carrying 

capability of all tows. 

Conclusions 

Triaxially braided composites exhibited substantial anisotropy and heterogeneity on the scale of tows, 

despite the repeating pattern. This created large local strain variation and high delamination stresses at 

free-edge tows. The use of overhung tubes eliminated the influence of unrestrained free-end tows, 

allowing estimation of the areal strength as opposed to the edge strength. The hoop strength of axially 

overbraided tubes was comparable to uniaxial measurements of the axial strength and to the net-section 

transverse strength when free-end tows were restrained and sheltered via notching. Although tube 

pressurization results in a mildly biaxial stress state, the hoop and axial stress maximums can be separated 

to some extent by pressurizing a short region. This also results in a uniform hoop stress through the tube 

wall. Distortions in the fiber architecture and section thickness, which usually occurred near mold lines, 

occasionally caused strain concentrations and low tube strength. Axially wrapped tubes, in which the ply 

end was not restrained, exhibited low strength due to ply delamination. Failure of hoop loaded axial 

braided tubes ensued by relaxation of axial tows, rotation of bias tows that caused delamination, tow 

impingement and system stiffening, longitudinal tow splitting, and eventual failure where the bias tows 

cross axial tows. 

Notched uniaxial transverse specimens exhibited similar strength to tubes and axial specimens, 

however, some dependence on notch angle and depth exists, and nonlinear strain occurs earlier. In the 

case of notched axial specimens, the dependence is the result of load shedding to bias tows. In the case of 

transverse specimens, the dependence is the result of shielding of the bias tows from a free edge: the 

restrained tows crossing the ligament act as structural elements or “tie rods” that connect the load grips. 

Unconstrained tows only transmit load thru the surrounding epoxy. Depending on the notch design and 

orientation, the areal strength can be under or overestimated. For transverse orientation, underestimates 

tend to occur. For deeper notches, which have more surrounding material, higher strength is exhibited. 

Effective use of all tows by restraint is a key aspect for strength optimization. 
NASA/TM—2014-216615 8 

References 

1. J.D. Littell, W.K. Binienda, G.D. Roberts, and R.K. Goldberg, “Characterization of Damage in 

Triaxial Braid Composites Under Tensile Loading,” NASA/TM—2009-215645. 

2. C.L. Bowman, G.D. Roberts, M.S. Braley, M. Xie, and M.J. Booker, “Mechanical Properties of 

Triaxial Braided Carbon/Epoxy Composites,” in 35th International SAMPE Technical Conference—

vol. 35, Sept. 29-Oct. 2, 2003, Dayton OH (SAMPE Covina, CA 2003). 

3. L.W. Kohlman, J.L. Bail, G.D. Roberts, J.A. Salem, R.E. Martin, W.K. Binienda, “A notched coupon 

approach for tensile testing of braided composites,” Composites Part A: Applied Science and 

Manufacturing, Volume 43, Issue 10, October 2012, Pages 1680–168. 

4. S. Schwentes and W. Elber, “Slow Crack Growth in Spinel in Water,” NASA Technical 

Memorandum 85644, May 1983. 

5. J.A. Salem and J.Z. Gyekenyesi, “Burst Pressure Testing of Ceramic Rings,” NASA TM—2007-

214695, 2007. 

6. M.J. Verrilli, J.A. DiCarlo, A. Calomino, H.M. Yun, and T.R. Barnett, “Hoop Tensile Properties of 

Ceramic Matrix Composite Cylinders,” Journal of Testing and Evaluation, vol. 33, no. Sept. 2005. 

7. R.C. Reuter and T.R. Guess, “Analysis, Testing, and Design of Filament Wound, Carbon-Carbon 

Burst Tubes,” Composite Materials: Testing and Design, ASTM STP 546, pp. 67–83, 1974. 

8. K. Mosley “The Stressing for Test Purposes of Materials in Tubular Form Using Elastomeric Inserts 

– Experimental and Theoretical Development,” Proc. Inst. Mech. Eng., vol. 196, pp. 124–139, 1982. 
NASA/TM—2014-216615 9 

TABLE 1.—FIBER AND RESIN PROPERTIES 

Fiber Tensile strength, 

ksi (MPa) 

Young’s modulus, 

msi (GPa) 

Failure strain,

percent 

Density, 

lb/in.3

 (g/cm3) 

T700S 711 (4900) 33.4 (230) 2.1 0.065 (1.80) 

E862 8.8 (61) 0.39 (2.7) ------ 0.043 (1.270) 

5208 7.3 (50) 0.56 (3.8) ------ 0.046 (1.20) 

TABLE 2.—ELASTIC CONSTANTS FOR TRIAXIAL 

BRAID COMPOSITE AT 0.5 PERCENT STRAIN 

Material E11, 

(msi) 

E22, 

(msi) 

12 21

E862/T700 6.950.23 6.030.12 0.34 0.30 

5208/T700 6.690.24 6.230.25 0.29 0.29 

TABLE 3.—TENSILE STRENGTH OF T700S/E862 MEASURED 

WITH VARIOUS CONFIGURATIONS (ksi) 

Orientation Tube 

(Central pressurization) 

D3039 tensile 60

Notched tensile 

0

Notched tensile 

Axial 90.37.8 (3) 1184 (5) 1171 (2) 11213 (2) 

Transverse 1196 (3) 67.35.2 (5) 1254 (4) 11817 (2) 

TABLE 4.—TENSILE STRENGTH OF T700S/5208 MEASURED 

WITH VARIOUS CONFIGURATIONS (ksi) 

Orientation Tube 

(Central pressurization) 

D3039 tensile 60

Notched tensile 

0

Notched tensile 

Axial ------------ 1068 (6) 1099 (5) 1184 (5) 

Transverse 1134 (4) 442 (6) ------------ 10012 (5) 
NASA/TM—2014-216615 10 

Figure 1.—Maximum principal stresses for (a) constrained 

edge and (b) free edge conditions. 
NASA/TM—2014-216615 11 

Figure 2.—Stress distributions in a partially pressurized tube: (a) axial, (b) hoop, 

(c) radial, and (d) shear. 
NASA/TM—2014-216615 12 

Figure 3.—Stress distribution in a 6 in. partially pressurized isotropic tube: (a) 1.5 in. zone, (b) 1.875 in. zone, 

(c) 2 in. zone, and (d) 5 in. zone. 

Figure 4.—Strain distribution in an (a) aluminum and (b) triaxial braided composite notch tensile specimen 

subjected to fixed-end loading. 
NASA/TM—2014-216615 13 

Figure 5.—(a) Tube schematic and (b) experimentally 

measured and theoretical stresses for a partially 

pressurized infinite aluminum tube (half-symmetry). 
NASA/TM—2014-216615 14 

Figure 6.—Two-dimensional triaxial braid carbon fiber preform with highlighted 

unit cell. The blue arrow indicates the axial fiber direction, and the red arrows 

indicate bias fiber directions. 

Figure 7.—Notched tensile configuration. 
NASA/TM—2014-216615 15 

Figure 8.—ASTM D3039 axial and transverse global stress-strain curves for (a) T700S/E862 and 

(b) T700S/5208. 
NASA/TM—2014-216615 16 

Figure 9.—Uniaxial net section strength as function of notch angle and depth for T700S/5208. 

Figure 10.—Examples of notched axial tensile tests: (a) splitting between notched and 

unnotched regions; (b) axial strain distribution about a 120 notch. Examples of notched 

transverse tests: (c) strain concentrations at bias tows and strain banding at tow 

intersections; (d) strain concentration along a 120 free edge. 
NASA/TM—2014-216615 17 

Figure 11.—Strain banding at 32 ksi (a) and correspond axial (b) and transverse (c) stress-strain curves for the high 

and low strain regions of a T700/5208 transverse tensile test. 
NASA/TM—2014-216615 18 

Figure 12.—Burst tubes with (a) axial tows in axial direction, and (b) axial tows in hoop direction. 
NASA/TM—2014-216615 19 

Figure 13.—Strength-of-materials stress for (a) T700S/E862 and (b) T700S/5208 as a function 

of strain for pressurized axially-woven tubes and ASTM D3039 transverse tensile specimens. 
NASA/TM—2014-216615 20 

Figure 14.—(a) Stress-strain curves for T700S/E862 tubes with (b) damaged (TR-5 at 34 ksi) and (c) distorted 

fibers at mold lines (TR-9 at 89 ksi). 
NASA/TM—2014-216615 21 

Figure 15.—Hoop strain enhancement at distorted fibers and axial tows in an axially woven T700s/E862 

tube (TR-7) at 45 percent of failure (44 ksi). 

Figure 16.—Strength-of-materials stress variation in a T700S/E862 pressurized tube (TR-3) along 

(a) tube circumference, (b) tube longitudinal axis at a SOM stress of 52 ksi, (c) strain variation at 

104 ksi, and (d) burst tube. 

(c) MoldLine

MoldLineAxial

tow

Section0NASA/TM—2014-216615 22 

Figure 17.—Stress as a function of strain for high and low strain regions of a T700S/E862 pressurized tube 

(TR-2): (a) local strain gages and high strain region near mold line at 72 ksi, (b) failed tube, (c) hoop (x) 

strain, and (d) axial (y ) strain). 

Figure 18.—Examples of tow delamination, splitting, and breakage occurring just before failure in a 

pressurized T700S/5208 tube (1035B).