February 2011 

NASA/TM–2011-217049 

Evaluation of Long Composite Struts 

Dawn C. Jegley and K. Chauncey Wu 

Langley Research Center, Hampton, Virginia 

James E. Phelps and Martin J. McKenney 

ATK Space Systems, Inc., Hampton, Virginia 

Leonard Oremont 

Lockheed Martin, Hampton, Virginia 

Ansley Barnard 

University of Washington, Seattle, Washington 
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February 2011 

NASA/TM–2011-217049 

Evaluation of Long Composite Struts 

Dawn C. Jegley and K. Chauncey Wu 

Langley Research Center, Hampton, Virginia 

James E. Phelps and Martin J. McKenney 

ATK Space Systems, Inc., Hampton, Virginia 

Leonard Oremont 

Lockheed Martin, Hampton, Virginia 

Ansley Barnard 

University of Washington, Seattle, Washington 
Available from: 

NASA Center for AeroSpace Information 

7115 Standard Drive 

Hanover, MD 21076-1320 

443-757-5802 

The use of t rademarks or na mes of m anufacturers in thi s re port is for accurate reportin g and does not 

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National Aeronautics and Space Administration. 
1

Evaluation of Long Composite Struts 

Dawn Jegley, K. Chauncey Wu

NASA Langley Research Center, Hampton VA 

James Phelps, Martin McKenney

ATK Space Systems, Inc., Hampton VA

Leonard Oremont 

 Lockheed Martin, Hampton VA

Ansley Barnard

University of Washington, Seattle WA

Carbon-epoxy tapered struts are structurally efficient and offer opportunities for 

weight savings on aircraft and spacecraft structures. Seven composite struts were 

designed, fabricated and experimentally evaluated through uniaxial loading. The 

design requirements, analytical predictions and experimental results are presented. 

Struts with a tapered composite body and corrugated titanium end fittings 

successfully supported their design ultimate loads with no evidence of failure.

I. Introduction

NASA evaluates the efficiency of structural concepts for aircraft and spacecraft components with the 

goal of designing lightweight parts. Currently NASA is interested in structurally efficient truss structures 

for application to spacecraft and aircraft components such as supports for antennae or solar panels, lunar 

lander struts, components for use on the lunar surface, strut-braced wing components, aircraft spars or 

unmanned aerial vehicles. 

Structural efficiency is defined here as the ratio of a structure’s load-carrying ability to its mass. 

Studies involving the structural efficiency of shells date back to the NACA days and are continued into 

aircraft design today.1-7 Composite materials and non-traditional designs offer opportunities for reduced 

mass compared to conventional aluminum tubes with a uniform circular cross section. An example of a 

structure dependent upon struts in the truss arrangement is the Altair lunar lander8

 as shown in Fig. 1. The 

Altair project is one element of NASA’s Constellation program, which is intended to return human’s to the 

Moon within approximately the next 10-15 years. Because of the need to reduce mass, the lunar lander 

truss elements must be designed to operate at maximum efficiency. The use of composite struts to directly 

carry propellant and payload induced launch loads is generally speaking unique. Long struts have been 

evaluated in the context of deployable parts of a space station, telescope elements, and solar panels. In 

general, these applications require precise positioning but carry little load and tend to be driven by stiffness 

rather then strength requirements. In contrast, some of the struts for Altair must carry loads greater than 

100,000 lb.8

 Therefore, the primary focus of this study is heavily loaded struts, although several lightlyloaded strut designs are discussed as well. 

The struts in the current study involve predicted maximum loads ranging from 1,260 lb to 129,998 lb. 

Struts were designed to a specified Design Ultimate Load (DUL) with a 

1.4 safety margin with respect to strength, Euler buckling, and local 

buckling so 1.4 was used as the margin to determine Design Limit Load 

(DLL). Detailed discussions of the design, optimization and fabrication 

of the struts discussed herein is presented in references 9-14. The focus 

of the current paper is the experimental behavior and a comparison to the 

finite element analysis of each strut. Two struts fabricated by Boeing and 

five struts fabricated by Northrop Grumman were evaluated 

experimentally while subjected to uniaxial loading. Results of these 

experiments are described in the following sections. Figure 1. Altair lunar lander.
2

II. Test Article Description

All struts were fabricated using IM7-977-3 or IM7-8552 carbon fibers and epoxy resin and titanium 

end fittings. Each strut consists of a composite tube body and a titanium end fitting on each end. Two 

types of struts were considered. The Northrop Grumman struts contained a composite body-to-fitting joint 

in which the fibers were wrapped around the corrugations in the titanium and the other Boeing struts 

contained a step-joint fitting where the titanium was bonded to the composite. Struts with the corrugated 

end fitting had a composite body with a wider diameter in the center and small diameter on the end. Struts 

with the step-lap joint bonded fitting had a uniform diameter body and a tapered titanium fitting.

Seven full-scale, optimumdesign, test articles were fabricated based on designs developed 

in references 12 and 13. Photographs of representative test 

articles are shown in Fig. 2. Struts 

designed to support a compressive 

DUL of 107,845 lb are shown in 

Figs. 2a and 2b. A strut designed 

to support a compressive DUL of 

1,182 lb is shown in Fig 2c. Struts 

shown in Figs. 2a and 2c have 

corrugated end fittings and the 

strut shown in Fig. 2b has a 

bonded end fitting. Photographs of 

typical fittings (prior to strut fabrication) are shown in 

Fig. 3. Descriptions of the composite struts are shown in 

Table 1. A sketch of a strut with the definitions of the 

dimensions used in Table 1 is shown in Fig. 4. All 

fittings contained internal threads to match the test 

fixtures.

Strut types are identified herein by their designer (B 

for Boeing and N for Northrop Grumman) and their load 

level (L for lightly loaded struts and H for heavily loaded 

struts). Additionally, a replicate number of 1,2, or 3 is 

added to differentiate individual struts within a type. 

Table 1. Strut Description

Strut 

type

Qt

y

Strut Length

(in.),

L

Central

Inner 

Diameter

(in.),

D

Central Stacking 

Sequence

%0/%45/%90

Composite 

Untapered

Length

(in.),

b

Taper 

Location, 

Angle,

α

Fitting 

Length 

(in.),

f

Fitting 

Concept*

B-H 2 77.7 5.8 63/25/12 61.0 Fitting, 22o 8.35 Bond

N-H 3 77.7 6.0 58/24/18 50.7 Body, 10o 2.94 Cor.

N-L 2 100.3 2.0 33/44/22 92.7 Body, 10o 1.25 Cor.

* Cor.=corrugated; Bond=bonded 

 The composite strut bodies of the heavily loaded struts contained 16 and 18 plies with stacking 

sequences [+45/03/90/02]s for struts B-H and [90/02/+45/03/90]s for struts N-H. B-H and N-H struts were 

designed to support an axial compressive DUL of 107,585 lb and an axial tensile DUL of 60,182 lb. The 

composite strut bodies for the lightly loaded struts contained nine plies with a stacking sequence of 

[90/0/+45/0]s and were designed to support 1,183 lb in compression and 1,443 lb in tension. Additionally,

 a) Heavily loaded Northrop Grumman strut.

b) Heavily loaded Boeing strut.

c) Lightly loaded Northrop Grumman strut.

Figure 2. Strut test articles.

Figure 3. Titanium fittings.
3

several extra 90o plies were added to a small region 

of the N-L and N-H struts around the outermost 

end of the taper, as shown in Fig. 4. These 

additional wraps were applied to reduce the stress 

concentration in that region without adding another 

ply along the whole length of the strut. 

The struts with the tapered body and the 

corrugated end fittings, N-H and N-L, had the endmost fibers wound around the corrugations in the 

fitting to create a mechanical joint as well as curing the part with the fitting in place to create an adhesive 

bond. The corrugated fitting concept has a series of grooves in the outer surface of the fitting which allow 

fibers to be placed into the groves and overwrapped with additional fibers which lock the layers in place. 

After the strut is cured, the fibers are locked into the grooves and the fitting cannot be removed solely by 

breaking the adhesive bond. The only way to fail this joint is to break the fibers themselves. The geometry 

of the fitting is dependent upon the design load levels. Details of the manufacturing approach are presented 

in reference 12. 

The composite portion of the struts of type B-H, which have a uniform diameter body and tapered end 

fittings, were fabricated with dropped plies on each end to create steps of reduced numbers of plies. The 

machined titanium fittings were bonded to the dropped-ply region to allow an adhesive bond to hold the 

fittings in place. The bonded fitting concept uses a lap-stepped joint and adhesive to attach the titanium 

fitting to the carbon-epoxy during cure. When the bonded fitting was used, the taper was in the fitting 

rather than in the composite so a much steeper taper angle was achieved than could be fabricated in a 

composite body taper. However, a longer fitting was required to gain adequate bond area. Details of 

manufacturing approach are presented in reference 13.

Geometric imperfection evaluations for all struts were conducted either by the contractor who did the 

fabrication or by NASA. The only imperfection shape considered herein is a single half-wave deformation 

as shown in the sketch in Fig. 5. The measured imperfection amplitudes for each strut are shown in Table 

2. 

III. Analysis

All fabricated struts were analyzed using the 

finite element code ABAQUS16 to determine the 

buckling load, strain distribution and failure load. 

Linear analyses indicated that none of the struts 

would fail in the composite body in tension due to 

excessive strains at loads less than DUL. 

However, additional factors beyond strains had to 

be considered for compression. In each 

compression analysis an imperfection shape of one 

axial half-wave was assumed, as shown in Fig. 5. 

The largest measured imperfection was used as the 

wave’s amplitude in the nonlinear analysis.

A typical finite element model is shown in Fig. 6. Between 10,000 and 70,000 shell elements were 

used in each model. The properties of carbon-epoxy and titanium, as shown in Table 3, were used for the 

analysis. Pinned boundary conditions were implemented by adding multi-point constraints to attach each 

node at the end of the strut to a single load application point. These constraints are indicated by the pink 

connectors in Fig. 6. This load application point was located three to six inches from 

Figure 5. Shape of initial geometric imperfection.

Table 2. Measured Geometric Imperfections

Strut ID Measured imperfection 

amplitude (in.)

NASA 

measurement

Contractor 

measurement

N-L-1 0.020 0.029

N-L-2 0.031 0.012

B-H-1 0.020 0.004

B-H-2 0.020 0.004

N-H-1 0.006 0.0043

N-H-2 0.034 0.225

N-H-3 0.025 0.023

Figure 4. Reinforcement at taper edge.
4

the end of the strut. Displacements and rotations 

were constrained at this point to avoid 

convergence difficulties in the analysis. This 

pinned condition is assumed to be conservative 

since the actual hardware is more rigid than a true 

pinned condition but not as rigid as a clamped 

condition. Load-displacement and load-strain 

relationships were predicted for all struts to guide 

testing. Detailed analyses of the fittings and the 

fitting-composite interface were conducted during 

the design process and are presented in references 

12 and 13, but are not presented herein. 

Predictions of maximum loads for each strut type 

are shown in Table 4. Bifurcation buckling loads 

were calculated with no assumed imperfections 

while maximum predicted loads were calculated based on the assumed imperfection.

Table 3. Assumed Material Properties

Carbon-epoxy

Compression

Carbon-epoxy

Tension

Titanium

Tension and 

Compression

E11, psi 21,400,000 24,500,000 16,000,000

E22, psi 1,460,000 1,260,000 16,000,000

G12, psi 690,000 690,000 6,060,610

U12 0.3 0.3 0.32

Table 4. Analytical Predictions

Strut 

Type

Required 

Compressive 

DUL (lb)

Required 

Tensile 

DUL (lb)

Predicted 

Bifurcation 

Buckling Load 

(lb)

Predicted 

Maximum 

Load (lb)

Assumed 

Imperfection 

Amplitude

(in.)

B-H 107,585 60,182 121,600 113,804 0.020

N-H 107,585 60,182 141,700 129,998 0.043

N-L 1,183 1,443 1,290 1,260 0.029

IV. Specimen Preparation and Testing

Between twenty and forty strain gages were attached to the outer surface of each test article prior to 

testing. Strain gages were applied at select locations to monitor stress concentrations and buckling 

behavior. Typical strain gage patterns are shown in Fig. 7. In addition, one section of each specimen was 

painted with a black and white speckle pattern so that a Vision Image Correlation (VIC)17 system could be 

used to monitor full-field displacements and strains over the taper and fitting region on one end of each 

strut, as shown for specimen N-H-1 in Fig. 8.

Strut N-H-3 was fabricated with Teflon inserts to simulate manufacturing-induced delaminations that 

may not be detected. These inserts were placed at selected locations through the thickness and along the 

length. The location of these inserts is shown in Fig. 9 where the color of the squares indicating the Teflon 

inserts indicates the location through the thickness, with green being the closest to the outer surface, blue 

being the closest to the inner surface and red being in the center of the stacking sequence. Strut N-H-3 was 

also subjected to impact damage prior to loading. The impact was inflicted using a drop weight impactor 

with a 25 lb weight with a 1-inch diameter tup. Trial impacts of 40, 80 and 100 ft-lb were inflicted to strut 

N-H-1 after it was failed to determine the impact energy to use on strut N-H-3. A photograph of the results

of these trials is shown in Fig. 10. Since the 100 ft-lb impact resulted in minimal visible damage, this level 

was selected for impact to strut N-H-3. A photograph of the pre-test damage to strut N-H-3 is shown in Fig. 

11. 

Figure 6. Typical finite element model.
5

Each test article was mounted vertically between 

fixtures and a load applied through the threaded end fittings. 

Photographs of the test arrangements are shown in Fig. 8. In 

each case, the strut fitting at the top was attached to brackets 

which attached to a rigid backstop. The test fixtures at the 

top and bottom were designed to allow for rotation in any 

direction during the testing as described in reference 15. For 

the heavily loaded struts, the fitting at the bottom of the strut 

attached to a 225 kip load cell, which was attached to a 225 

kip actuator which was attached to the backstop and 

supported by the floor. For the lightly loaded struts, the 

fitting at the bottom was attached to a 10 kip load cell. Load 

was applied through a hand pump hydraulic actuator. More 

details of the testing arrangement and set-up are presented in 

reference 15.

Struts were loaded in tension alone, compression alone 

or tension followed by compression. Heavily loaded struts 

were loaded at a rate of 0.001 in/sec in either tension or 

compression until the desired load or failure was reached. 

Pristine struts N-H-1 and N-H-2 were loaded to DLL in 

tension then to their maximum loading in compression. 

Strut B-H-1 was loaded to DLL in tension and then to failure in compression. Strut B-H-2 was loaded only 

to failure in compression. Strut N-H-3 was loaded to its maximum load in compression. Strut N-L-1 was 

loaded to 2 times its DUL in tension and then to its maximum load in compression. Strut N-L-2 was loaded 

to its maximum load in compression. 

Figure 8. Strut B-H-1 in the test fixture prior to testing. 

Figure 9. Location of inserts and damage for strut N-H-3.

Northrop Grumman Boeing

Figure 7. Typical strain gage pattern.
6

The behavior of each test article was monitored by numerous linearly variable displacement 

transducers (LVDT). Three LVDTs placed at 0, 120, and 240 degrees around the circumference of the strut 

measured axial lengthening or shortening of the entire strut; one LVDT measured the axial displacement of 

the constant-cross section region of the composite body, and two LVDTs at 0 and 90 degrees around the 

circumference measured radial motion at the midlength of the strut. Strain gage data, LVDT data, and VIC 

images of the lower region of each specimen were recorded at the rate of 1 Hz.

V. Results and Discussion

A summary of the displacements, strains and failure modes for the seven test articles is presented 

herein, and these results are compared to analytical predictions. Predicted and experimental maximum

compression loads are shown in Table 5. 

Table 5. Strut loads.

Strut ID Design 

Requirement (lb)

Predicted

Maximum Load (lb)

Experimental

Maximum Load (lb)

B-H-1 107,585 113,804 43,500

B-H-2 107,585 113,804 79,900

N-H-1 107,585 129,998 119,300

N-H-2 107,585 129,998 121,700

N-H-3 107,585 not analyzed 21,100

N-L-1 1,183 1,290 2180

N-L-2 1,183 1,290 1980

A. Loads and Displacements 

1. Heavily loaded struts

The heavily loaded bonded-fitting struts did not support the required loads. The first bonded-fitting 

strut tested, B-H-1, sustained damage in the initial tensile loading to DLL so it only sustained 

approximately 43,000 lb in the subsequent compressive loading. In order to start with a pristine strut for 

compression loading, the second bonded-fitting strut, B-H-2, was not initially loaded in tension. Even 

without the preload, strut B-H-2 failed in compression at approximately 80,000 lb, or approximately DLL. 

Little warning of impending failure was evident in the global measure of displacement. The end-shortening 

versus load for the compressive loadings are shown in Fig. 12. Strut B-H-1 measurements are shown as 

blue dashed lines for the three measurement locations around the circumference. Strut B-H-2 

measurements are shown as dashed red lines while the analytical prediction is shown as the solid black line. 

Figure 11. Impact damage on strut N-H-3.

 Figure 10. Trial impacts.
7

The three measurements for each strut 

agree well with each other and with 

analysis, indicating no buckling behavior is 

occurring in either strut and the struts have 

similar global stiffnesses.

 The radial deflections measured at two 

points at the midlength of each strut can be 

used to evaluate out-of-plane motion. 

These measurements for struts B-H-1 and 

B-H-2 are shown in Fig. 13 as blue and red 

lines, respectively. Since both of these 

measurements remain less than 0.02 inches, 

there was probably no buckle prior to 

failure. 

Both heavily loaded pristine struts with 

corrugated fittings withstood their tensile 

DLL prior to compressive loading. Under 

compressive loading, one heavily loaded 

pristine corrugated-fitting strut withstood 

110% of DUL while the other withstood 

112% of DUL. All evidence of failure was in 

the fitting region. The end-shortening versus 

load for the compressive loadings are shown 

in Fig. 14. Strut N-H-1 measurements are 

shown as blue dashed lines for the three 

measurement locations around the 

circumference. Strut N-H-2 measurements 

are shown as dashed red lines while the 

analytical prediction is shown as the solid 

black line. The three measurements for each 

strut agree well with each other and with 

analysis for load less than 108,000 lb 

indicating no buckling behavior is occurring 

in either strut prior to DUL and the struts 

have similar global stiffnesses. 

The average of the three axial 

displacement measurements for all three N-H 

struts are presented in Fig. 15. The impactdamaged strut, N-H-3, failed through the impact 

site at a load of 21,100 lb. The initial stiffness of 

this strut was the same as N-H-1 and N-H-2 but 

this damage, though not readily visible without 

the right lighting, caused severe enough damage 

to reduce the failure load by 80 percent. 

 The radial deflection measured on two points 

at the midlength of each strut can be used to 

evaluate the out-of-plane motion of each strut. 

These measurements for struts N-H-1 and N-H-2 

are shown in Fig. 16 as blue and red lines, 

respectively. Since both of these measurements 

are very small until immediately prior to failure, 

these struts did not buckle prior to failure. 

Figure 13. Radial displacement for struts of type B-H.

Figure 14. Axial displacement for pristine struts 

of type N-H.

Figure 12. Axial displacement for struts of type B-H.
8

2. Lightly loaded struts

The lightly loaded struts withstood the 

applied tensile and compressive loading with no 

visible damage but buckled at their maximum 

loading. Strut N-L-1 was loaded to two times its 

DUL in tension then loaded in compression to its 

maximum (buckling) load. Its maximum load was 

approximately 2.0 times the predicted maximum 

compressive load. Strut N-L-2 was not preloaded 

in tension but only loaded in compression and 

supported approximately 1.6 times its predicted 

maximum load in compression. Both of the 

lightly loaded struts supported more than DUL in 

compression. 

The average of the axial displacement 

measurements for specimens N-L-1 and N-L-2 

are shown as the green lines in Fig. 17. 

Analytical predictions are shown as back lines. 

The strut motion is not uniform around the 

circumference, however, the measurements are 

repeatable from one strut to the next. In both 

cases, good agreement between an average of 

the three test measurements (shown as the 

solid green lines) and the analysis can be 

seen for the initial loading of the struts. The 

struts buckled at a greater load than predicted 

by the analysis, possibly indicating that the 

influence of the initial geometric 

imperfection may not have been as 

significant as anticipated or that there may 

have been friction or other behavior at the 

ends in the fixtures to invalidate the “pinned” 

assumption at the ends. The full-field axial 

motion obtained from the vision image 

correlation system for a load level just prior 

to failure of 2,187 lb for strut N-L-1 is also 

shown in Fig. 17 for the bottom 10 inches of 

the strut. A relatively uniform axial 

deformation is seen. Note that the white 

regions in the figure represent areas for 

which the strain could not be calculated, 

typically because of surface obstructions or shadow.

The radial deflection measured on two points at the midlength of each strut can be used to evaluate 

the out-of-plane motion of each strut. These measurements for struts N-L-1 and N-L-2 are shown in Fig. 

18 as blue and red lines, respectively. Radial motion indicates buckling occurred for both struts at a similar 

load of approximately 2000 lb and in a similar direction. Applied axial displacement was continued 

beyond the initial buckling point so the radial deflection curves can be seen to flatten out where the 

increased axial displacement supported no additional load. The full field measurements at a load just prior 

to failure for strut N-L-1 in the directions parallel to and perpendicular to the backstop are also shown in 

Fig. 18. As the strut buckles, this lower 10-inch region of the strut moves less than 0.1 inches parallel to 

the test wall (top fringe plot) and up to 0.5 inches toward the test wall (bottom fringe plot) as it buckles.

Figure 15. Average axial displacement for struts 

of type N-H.

Figure 16. Radial displacement for struts of type N-H.
9

B. Strain

1. Heavily loaded struts

 

Midlength axial strains for specimens B-H-1 and B-H-2 are shown in red and blue, respectively, in Fig.

19. Midlength axial strains for specimens N-H-1 and N-H-2 are shown in red and blue, respectively, in 

Fig. 20. Measured axial strains at three locations around the circumference are shown as three dashed lines 

of different dash lengths in Figs. 19 and 20. Midlength strains for struts B-H are linear to failure and 

consistent between the two struts and around the circumference. Midlength strains for struts N-H are linear 

until immediately prior to failure and are also consistent between the two struts and around the 

circumference.

Figure 17. Axial displacement for struts of type N-L with full field displacement at a load of 

2,187 lb.

Figure 18. Radial displacement for struts of type N-L with full field displacement at 

a load of 2,187 lb.
10

Measured axial strains at the fitting-to-composite interface region for specimens B-H-1 are shown in 

Fig. 21 for the tensile loading and Fig. 22 for the subsequent compressive loading. The full-field axial 

strain for the lower end of strut B-H-1 at approximately 43,835 lb tensile load and 43,615 lb compressive 

load are shown in Figs. 21 and 22, respectively. Axial gages were located on the exterior surface of the 

composite body and on the exterior surface of the titanium fitting at three locations around the 

circumference. Strains on the composite are shown in red and strains on the titanium surface are shown in 

blue. The discontinuities in the measured strain under tensile loading are believed to indicate damage to 

the bond between the composite steps and the titanium fitting. Damage to the bond began to initiate at a 

load significantly less than DLL in the tensile loading. 

The strains at the interface between the composite and the titanium fitting are shown in Fig. 23 for strut 

B-H-2 loaded in compression. The full field strain for strut B-H-II-2 at approximately 79,800 lb 

compressive load is shown in figure 23. The strains in the titanium are not the same as in the composite 

body, however, both are linear and consistent around the circumference.

The strains on the composite surface in the taper region of strut N-H-2 are shown in Fig. 24. The red 

lines represent strains in the uniform diameter region adjacent to the taper and the blue lines represent 

strains in the taper itself. The strains are not uniform around the circumference, but all remain linear for 

load less than DUL. Nonlinearities can be seen in the strains in the taper region for load greater than DUL 

which could represent a failure in the adhesive bond between the carbon-epoxy and the titanium fitting.

Figure 20. Axial strain midlength for struts of type N-H.

Figure 19. Axial strain midlength for struts of type B-H.
11

Figure 22. Axial strain at interface location for strut B-H-1 loaded in compression with full field 

strain at a load of 43,615 lb.

Figure 23. Axial strain at interface location for strut B-H-2 loaded in compression with a full 

field strain of 79,980 lb.

Figure 21. Axial strain for interface location for strut B-H-1 loaded in tension with full field 

strain at a load of 43,835 lb.
12

The bending strain can be determined by averaging the strains measured at the midlength location and then 

subtracting this value from each measured value. This calculation removes the axial component of the 

strain, leaving only the bending component. The bending component at the three circumferential locations 

for strut N-H-2 is shown in Fig. 25. The bending strain is less than 10 percent of the total strain, also 

shown in the figure. Struts B-H did not display bending behavior.

Figure 24. Axial strain at taper for struts of type N-H-2.

-6,000 -4,000 -2,000 0 2,000

-140,000

-120,000

-100,000

-80,000

-60,000

-40,000

-20,000

0

Load, lb

Measured strain

Measured strain-average strain=bending strain

Average axial strain

Test data at three 

points for each strut

0 degrees

120 degrees

240 degrees

average 

-0.006 -0.004 -0.002 0 0.002

Strain, inches/inch

Figure 25. Bending strain for strut N-H-2.
13

2. Lightly loaded struts

Midlength axial strains for specimens N-L-1 and N-L-2 are shown in blue and red, respectively, in Fig.

26. Measured axial strains at three locations around the circumference are shown as three dashed lines of 

different dash lengths. Midlength strains for struts N-L show evidence of buckling well before failure, 

consistent with the displacement results. 

The strains on the composite surface in the uniform diameter region near the taper region and in the 

taper region itself of struts N-L-1 and N-L-2 are shown in Fig. 27. The red lines represent strains in the 

uniform diameter region adjacent to the taper and the blue lines represent strains in the taper itself. The 

strains in the uniform diameter region are uniform around the circumference and consistent between the 

two struts until immediately before failure. Nonlinearities can be seen in the strains in the taper region for 

the full loading range suggesting a redistribution of load in the corrugations as load is increased. 

-1,200 -800 -400 0 400

-2,500

-2,000

-1,500

-1,000

-500

0

SG#51 on 13 SG#52 on 14 SG#53 on 15 SG#51 on 13 SG#52 on 14 AU

Midlength axial strain, inches/inch

Load, 

lb

N-L-1

N-L-2

 -0.0012 -0.0008 -0.0004 

0 degrees

120 degrees

240 degrees

Test data at three 

points for each strut

Figure 26. Midlength strain for lightly loaded struts.

Figure 27. Strain in taper region of struts of type N-L.
14

The bending component at the three circumferential locations for strut N-L-1 is shown in Fig. 28. The 

bending strain is small for low loads but up to approximately 50 percent of the total strains as the loading is 

approaching the buckling load. The nonlinear region of the curves represents post-buckling behavior, 

where bending strain has the largest magnitude. 

C. Failure

Photographs of failed heavily loaded specimens are shown in Figs. 29 and 30. Both struts with the 

bonded end fittings failed the composite-to-fitting joint as shown in Fig. 29. The failure of the bond was 

evident around the entire circumference of the strut. Despite the fact that the bond area was maximized by 

locating it in the maximum diameter region of the strut, the adhesive bond did not sustain the required load. 

Visible failures in the corrugated-fitting struts were generally in the fitting region and of the type shown in 

-1,200 -800 -400 0 400

-2,500

-2,000

-1,500

-1,000

-500

0

Load, lb

Measured strain

Measured strain-average strain=bending strain

Average axial strain

Strain, inches/inch

-0.0012 -0.0008 -0.0004 0 0.0004

Test data at three 

points for each strut

0 degrees

120 degrees

240 degrees

average 

Figure 28. Measured and bending strains for strut N-L-1.

Fig. 29. Failure of a bonded fitting strut. Fig 30. Failure of a corrugated fitting strut.
15

figure 30 but not all struts displayed visible failure. Failure of the corrugated fitting struts were all at loads 

greater than DUL. Lightly loaded specimens buckled when they reached their maximum load and left little 

evidence of fiber damage after loading was removed. Photographs of strut N-L-2 showing its buckling at 

its maximum loading are shown in Fig. 31. A summary of these failures is presented in Table 6. NDE was 

not conducted after test to find any non-visible damage.

Table 6. Failure mode of tested struts.

Strut ID Failure mode

N-H-1 and N-H-2 Overwrap plies in fitting area

N-H-3 Along length and around circumference through impact site 

N-L-1 and N-L-2 Buckling, no significant visible strength failure

B-H-1 and B-H-2 Failure of bond between fitting and composite

X. Application to Altair Lunar Lander

Before this study was initiated, a great deal of work had been performed in examining lightly loaded 

struts but not as much in highly loaded struts. Design, analysis and testing had not been sufficient to have 

confidence that highly loaded composite struts were a feasible choice for the structure for the Altair lunar 

lander. For this reason, the technology readiness level (TRL)18 was assumed to be at a level of 4 based on 

the work on more lightly loaded struts for other applications. This study involved full-scale struts loaded to 

the maximum required uniaxial loading. In this study, a strut concept was proven to meet the loading 

requirements in both tension and compression. The design, analysis, and testing effort in this study 

demonstrated that the strut-based design is a viable concept. However, since no thermal or vibration 

conditions have been examined, as would be necessary to simulate a space and launch environment, the 

Figure 31. Buckling of strut N-L-1 under maximum loading. 
16

TRL has only been raised to 5. Further studies involving temperature extremes appropriate for orbit and 

long-term lunar exposure and vibrations similar to those anticipated on launch would be required to raise 

the TRL to 6.

XI. Concluding Remarks

Composite struts were designed with the objective of meeting the loading requirements for the most 

heavily loaded and the most lightly loaded struts for the Altair lunar lander. Seven minimum-weight 

carbon-epoxy struts were fabricated and experimentally evaluated through uniaxial loading. All struts 

contained titanium end fittings. Five struts were fabricated with a tapered composite body and a corrugated

end fitting where carbon fibers were wrapped around the corrugations to provide a mechanical connection 

between the fitting and the carbon-epoxy body of the strut. Two struts were fabricated using an adhesively 

bonded step lap joint between the carbon-epoxy body of the strut and the end fitting. The fittings were 

tapered in the struts with the step lap joint so the composite body could have a uniform diameter. Uniaxial 

tension and compression loading was applied to each type of strut to determine their strain, displacement, 

buckling and failure characteristics. 

Struts fabricated using the corrugated-fitting concept met the required compressive loading for heavily 

and lightly loaded struts. Heavily loaded struts of this type did not buckle and displayed failures in the 

fitting region where failures could be seen. Lightly loaded struts buckled into a single half-wave and 

displayed no evidence of strength failures. However, struts fabricated using the bonded-fitting concept 

failed at significantly lower loads than the design requirement. Damage to the bond began to initiate at a 

load significantly less than DLL in the tensile loading. Neither bonded fitting strut sustained load greater 

than DLL. A bonded joint depending only on the adhesive might be a viable concept but the struts in this 

study did not support the required load. Both bonded-fitting struts failed at the adhesive bond between the 

titanium and composite. Only the corrugated-fitting struts evaluated in this study meet the requirements 

defined by the Altair lunar lander project. In addition, one severely impact-damaged heavily loaded strut 

sustained barely visible damage but failed through the impact site at a load of approximately 20 percent of 

the failure load of the nominally identical pristine struts.

These results indicate that the use of highly loaded struts for a lander design carrying heavy loads 

through the primary structure during launch is feasible. As with any design, significant validation testing 

would be required. Because highly loaded struts tested here failed in the region of the end fittings, it is 

reasonable to suggest that end-fitting design may be a key area in the design of heavily loaded struts for 

Altair-type applications. The fitting-to-composite joints are critical both because it is a challenge to 

achieve sufficient strength at the tube-fitting interface, and because the mass of the joint may be a 

significant portion of the total strut assembly, reducing the mass benefit advantage of utilizing a truss-based 

design. 

References

1. Wagner, H., “Remarks on Airplane Struts and Girders under Compressive and Bending Stresses. 

Index Values.” NACA TM 500. 1929.

2. Dow, N. F., and Hickman, W. A., “Design Charts for Flat Compression Panels Having Longitudinal 

Extruded Y-Section Stiffeners and Comparison with Panels Having Formed Z-Section Stiffeners,” NACA 

NT 1389, 1947.

3. Williams, J., Anderson, M., Rhodes, M., Starnes, J., and Stroud, J., “Recent Developments in the 

Design, Testing and Impact-Damage Tolerance of Stiffened Composite Panels,” NASA TM 80077, 1979.

4. Williams, J. G., and Stein, M., “Buckling Behavior and Structural Efficiency of Open-Section 

Stiffened Composite Compression Panels.” AIAA Journal, Vol. 14, No. 11, 1976, pp. 1618-1626.

5. Williams, J. G., and Mikulas, M. M., Jr. “Analytical and Experimental Study of Structurally Efficient 

Composite Hat-Stiffened Panels. Presented at the ASME/AIAA SAE 16th Structures, Structural Dynamics 

and Materials Conference, Denver Colorado, 1995. AIAA Paper No. 75-754.

6. Dow, N. F., and Rosen, B. W., “Structural Efficiency of Orthotropic Cylindrical Shells Subjected to 

Axial Compression,” AIAA Journal, Vol. 4, No 3, 1966, pp. 481-485.
17

7. Haynie, W. T., and Hilburger, M., “Comparison of Methods to Predict Lower Bound Buckling Loads 

of Cylinders Under Axial Compression,” Presented at the 51st AIAA/ASME/ASCE/AHS/ASC Structures, 

Structural Dynamics, and Materials Conference, 2010, Orlando, Florida, AIAA Paper No 2010-2532.

8. Constellation Program: America’s Spacecraft for a New Generation of Explorers, The Altair Lunar 

Lander. NASA facts, Lyndon B. Johnson Space Center, 2008.

9. Brewster, Jebediah, Design of Structurally Efficient Tapered Struts NASA CR 2009-8393, 

December 2009.

10. Messinger, R., “Design of Structurally Efficient Tapered Struts,” NASA CR 2010-216698, 2010.

11. Deo, R., Benner, H., Dawson, V., Olason, E., and Harrison, R. “Design of Structurally Efficient 

Tapered Struts (SETS).” NASA CR 2010-216699.

12. Messinger, R., “Design and Manufacturer of Structurally Efficient Composite Struts – Concept 1,” 

Final report for NASA contract NAS1-NNL04AA11B Task NNL09AC35T, 2009.

13. Pires, K., Benner, H., Deo, R., Grover, R., Palm, T., McLaughlin, M., Olason, E., Lucking, R., and 

Collier, C., “Design and Manufacturer of Structurally Efficient Composite Struts – Concept 2,” Final 

report for NASA contract NAS1-NNL04AA13B Task NNL09AC36T, 2010.

14. Jegley, D., Wu, K., Phelps, J, McKenney, M., and Oremont, L., “Structural Efficiency of 

Composite Struts for Aerospace Application,” AIAA paper 925421 presented at the 52nd

AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics and Materials Conference, Denver CO, 

2011.

15. Wu, K., Phelps, J, McKenney, M., and Jegley, D., “Highly Loaded Composite Strut Test 

Development,” AIAA paper 932820 presented at the 52nd AIAA/ASME/ASCE/AHS/ASC Structures, 

Structural Dynamics and Materials Conference, Denver CO, 2011.

16. ABAQUS version 6.9 EF1 On-line Documentation, ABAQUS Analysis User's Manual, Dassault 

Systemes Simulia Corp., Providence, RI, 2009.

17. McGowan, David M., Ambur, Damodar R., and McNeil, Stephen R., “Full-field Structural 

Response of Composite Structures: Analysis and Experiment,” presented at the 44th

AIAA/ASME/ASCE/AHS Structures, Dynamics and Materials Conference, AIAA 2003-1623, Norfolk, 

VA, April 2003.

18. Mankins, John C., “Technology Readiness Levels, A White Paper,” April 6, 1995.
REPORT DOCUMENTATION PAGE Form Approved

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Evaluation of Long Composite Struts

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Jegley, Dawn C; Wu, K. Chauncey; Phelps, James E.; McKenney, Martin 

J.; Oremont, Leonard; Barnard, Ansley

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NASA Langley Research Center

Hampton, VA 23681-2199

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14. ABSTRACT

Carbon-epoxy tapered struts are structurally efficient and offer opportunities for weight savings on aircraft and spacecraft 

structures. Seven composite struts were designed, fabricated and experimentally evaluated through uniaxial loading. The 

design requirements, analytical predictions and experimental results are presented. Struts with a tapered composite body and 

corrugated titanium end fittings successfully supported their design ultimate loads with no evidence of failure. 

15. SUBJECT TERMS

Composites; Bonded joint; Buckling; Carbon-epoxy; Corrugated fitting; Graphite; Lunar lander

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