elevated temperature mechanical_characterization_of_isogrid_booms

American Institute of Aeronautics and Astronautics1

ELEVATED TEMPERATURE MECHANICAL CHARACTERIZATION OF ISOGRID BOOMS

Stephen E. Scarborough* and David P. Cadogan† – ILC Dover, Inc., Frederica, DELauren M. Pederson‡ and Joseph R. Blandino§* – James Madison University, Harrisonburg, VA

Gary L. Steckel∋ and Wayne K. Stuckey¶ – The Aerospace Corporation, El Segundo, CA

AbstractStructurally efficient isogrid booms, manufacturedfrom rigidizable composite materials, are becoming anenabling technology for spacecraft structures becauseof their high packing efficiency. Selection of thematerials used in the construction of rigidizable spacestructures is commonly driven by mechanicalperformance properties at elevated temperatures.Mechanical properties testing was performed oncomposite tow samples and on an isogrid boom atvarious temperatures. To characterize elevatedtemperature behavior, the isogrid booms, and its sub-element composite tows were manufactured fromILC’s TP283E shape memory polymer (SMP) matrixresin and a carbon reinforcement. Both the flexuralmodulus and the tensile modulus of the composite towsamples were determined as a function of temperature.These values were compared to the calculated valuesfor the composite based on rule of mixtures analysis.The predicted rule of mixtures composite modulus isused in ILC’s isogrid analytical code to predict thestructural properties of the isogrid boom. A number ofcomposite tow samples were fabricated by ILC andmechanically characterized by the AerospaceCorporation to gather independent performance data.An isogrid boom was fabricated by ILC andmechanically characterized at elevated temperatures byJames Madison University (JMU). JMU tested thisboom in tension, compression, and also performedpreliminary creep testing at various temperatures. Asimilar isogrid boom was fabricated by ILC and testedby The Aerospace Corporation for composite CTEperformance. This paper discusses the results of boththe composite tow testing and the isogrid boom testingin pre- and post-packing conditions. A discussion ofthe correlation between the predicted values and theactual test values is also presented.

IntroductionNASA and DoD space missions in the near future willrequire much larger satellites, the sizes of which willbe beyond the capabilities of current technologies. The

types of Gossamerspacecraft that will beneeded includeantennas, solar arrays,sunshields, solar sails,and telescopes (Figs.1-2). Some systemsbeing considered arehundreds of meters insize to accomplishmission goals. Due tothe increase in payloadsize required,innovative supportstructures, which canbe packed into thefaring of availablelaunch vehicles, mustbe developed. In recentyears, research anddevelopment work hasbeen performed in thisarea1-3. Of theavailable options, oneof the most promisingtechnological advancements is the rigidizableinflatable structure. A rigidizable inflatable structure isone that is fabricated on Earth, packed into the launchcontainer, and inflated for deployment once on orbit1,3.After deployment, the material is rigidized, orhardened, to form a stiff composite structure that nolonger needs the inflation gas for support. This class ofstructures has unique benefits such as low packingvolume, reduced mass, and in most cases, very highdeployed structural efficiency3,4.

Several types of construction can be used in arigidizable inflatable including monocoque, isogrid,IsoTruss, and truss-frame booms. Each compositestructure can be fabricated into a varying geometricshapes utilizing any number of resin and fiber types5.The fibrous reinforcement can be in tow or wovenfabric form. In order to optimize the structure, the sizesof the tows and the weave styles of the fabrics can bevaried5. It is also possible to manufacture near-zerocoefficient of thermal expansion (CTE) booms throughthe fiber and resin selection and by optimizing thevolume fractions of each5. However, key to allmechanical performance properties is the ability to foldand tightly pack the material.

*Member AIAA† Associate Fellow AIAA‡Undergraduate Research Assistant, Dept. of Int. Science and Tech.§Associate Professor, Dept. of Int. Science and Tech.∋Senior Scientist, Materials Sciences Dept.¶Distinguished Scientist, Space Materials Lab

Figure 2. ILC 3.2m DiameterTSU Hexapod Testbed

Figure 1. ½ Scale NextGeneration Space

Telescope Sunshield


Isogrid Design and ConstructionOne of the most advanced inflatable rigidizablestructures is the isogrid boom, which consists of a grid-work of equilateral triangles1,3,6. These equilateraltriangles give the overall structure isotropic mechanicalproperties1,3. One of the isogrid booms fabricated forthis study is shown in Figure 3 (patent pending). Thistype of boom has a circular cross-section and is encased on both sideswith a polymeric film such aspolyimide1,3. The inner film layeracts as a bladder or gas-retaininglayer for inflation. The outer later,called the anti-blocking layer, is usedto prevent the structure fromadhering (blocking) to itself when itis in the packed configuration1,3. Theouter layer can also act as the firstlayer in a mult-layered insulation(MLI) blanket. ILC Dover hasdeveloped analytical modelingtechniques to predict the propertiesof this structure1,3,6. ILC has studiedthe room temperature mechanicalproperties of this structureextensively, but there is still a needfor further work to be performed tofully characterize the performance ofthe structure, especially at elevatedtemperatures1,3.

MaterialsA leading rigidizable material candidate is the shapememory composite 1-5, 7,8. The composite consists of afibrous reinforcement, such as carbon, and a polymericmatrix resin such as polyurethane or epoxy. The resinis initially consolidated at a high temperature, calledthe set temperature, to form a high modulus, rigidstructure. This initial heating and consolidation eventdefines the shape of the structure. Subsequent heatingevents above the material’s glass transitiontemperature (Tg) lowers the modulus of the materialsignificantly and allows the structure to be tightlypackaged (Fig. 4). Ifconstrained and cooled belowTg while packed, thecomposite modulus willincrease and the material willretain the packed shape, evenif the package isunconstrained. Upon re-heating, the materialexperiences a large decreasein modulus and can thereforebe deployed back to its as-manufactured state. Oncethe deployed structure cools below the Tg, the modulus

will increase and the inflation gas is no longer requiredfor structural support. This process of packing anddeploying the structure is repeatable which allowsflight hardware to be packed and deployed forevaluation several times during ground test prior tolaunch and deployment in space. The thermoplasticresin has some degree of shape memory upon heating,which causes the structure to attempt to return to its as-manufactured state. However, this shape memory forceis weak relative to the force required to deploy thecomposite structure and associated systems, thereforethis force must usually be augmented by inflation gas.

Material PropertiesIn this study, ILC Dover’s TP283E epoxy resin wasused to manufacture all of the test samples. The resinsystem has reduced cross-link functionality in order toadd increased flexibility above the Tg as compared totypical epoxy resin systems. TP283E therefore exhibitsthermoplastic behavior with the only exception beingthat it does not have a melt temperature. ILC choseHexcel’s IM9 carbon fiber as the reinforcement for allof the test samples discussed herein. IM9 was chosenbecause of its balance of high mechanical propertiesand high strain to failur1,3,8. The properties of this fiberare listed in Table 1.

Table 1. IM9 (12K) Carbon Fiber Properties9

Property English Units SI UnitsTensile Modulus 42.0 x 106 psi 290GpaTensile Strength 890,000 psi 6,141 MPaUltimate Elongation 2.1% 2.1.%Axial CTE∝ -0.228 ppm/oF -0.5 ppm/oCDensity 0.0650 lb/in3 1.80 g/cm3Filament Diameter 0.175 mil 4.4 micronsFilament shape Round RoundWeight/Length 18.8 x 10-6 lb/in 0.335 g/mFiber Cross-Sectional Area 2.89x10-4in2 0.19mm2

∝Estimate

In earlier studies, assumptions were made based on testdata from similar resins to predict the mechanicalproperties of TP283E resin in order to eliminate theoverall amount of testing required in initialdevelopment1,3,8. To determine the actual properties ofthe resin using an independent test lab, a resin samplewas manufactured by ILC and sent to The AerospaceCorporation for mechanical characterization. Theelastic modulus of the neat resin was tested using the 3-point bending method on a TA Instruments Inc.,Dynamic Mechanical Analysis (DMA) system. Fromthis testing, the average elastic modulus at 23°C for 5samples of TP283E resin was determined to be 500 ±20 ksi (Fig. 5).

Figure 3. ILCIsogrid Boom:

Figure 4. Z-Folded7” Diameter, 39”

Long IsogridBoom


Using the cantilever bending function on the DMA, TheAerospace Corporation determined from the lossmodulus peak that the Tg of the resin was 55oC at afrequency of 1 Hz. Three tow samples extracted froman isogrid boom were also tested using DMA (Fig. 6).The results from these tests indicate that the Tg isbetween 57oC-61oC. Earlier Differential ScanningCalorimetry (DSC) testing performed at ILC indicatedthat TP283E had a Tg of 48oC1. Differences in the testmethods and slight deviations in the set temperaturesare the likely causes of the variation in the Tg results.Therefore, depending on the set temperatures, TP283Ecan be made to have a Tg in the range of 48-61oC.

The flexural modulus of TP283E resin over thetemperature range from -150°C to +50°C was alsotested. Tests were performed on sample numbersB3−B7, which were approximately 2.3-inches long,0.35-inches wide, and 0.12-inches thick. They weretested using the 2-in. span of the DMA 3-Point Bendfixture. Peak loads were varied for the different testtemperatures because the neat resin samples hadincreasingly lower moduli at higher temperatures.

The temperature sequence was 23, 30, 35, 40, and50°C. At the end of each test the force was set to 0 N.The force was then maintained at 0 N during heating tothe next test temperature and during the 5-minutethermal equilibration at the test temperature. After the5-minute thermal equilibration at the test temperature,the static force was then set to 0.01 N and immediatelyramped to the final force. Good modulus data wereobtained for all test temperatures for sample B7 (Fig.7). Modulus values for sample B7 were similar to thosefor sample B6 for all test temperatures.

After the modulus testing was completed, the validmodulus data for samples B3−B7 were tabulated andplotted as functions of temperature. It was determinedthat two linear curve fits could be used to describe thedata. A straight line with a relatively low slope (-1.3ksi/°C) was fit to the data for -150 to 23°C and astraight line with a much higher slope (18.7 ksi/°C) wasfit to the data for 23 to 50°C (Fig. 8). The two curvesintersect at 25°C.

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.0 0.1 0.2 0.3 0.4 0.5

Flexural Strain, %

Fle

xura

l St

ress

, ks

iNo. 1No. 2No. 3No. 4No. 5

E = 535 ksiE = 501 ksiE = 493 ksiE = 490 ksiE = 482 ksi

Average E = 500 + 20 ksi(0.04 to 0.25% Strain)

Figure 5. Elastic Modulus at 23oC of TP283E Resin

Flexural Stress-Strain CurveTP283E Sample No. B7

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

0 0.002 0.004 0.006

Flexural Strain, in/in

Flex

ural

Str

ess,

ksi

25C, E = 445 ksi

32C, E = 362 ksi36C, E = 261 ksi

41C, E = 72 ksi45C, E = 49 ksi

50C, E = 22 ksi

Figure 7. TP283E Elastic Modulus Vs. Temperature

TP283E Elastic Modulus versus Temperature

0

100000

200000

300000

400000

500000

600000

700000

800000

900000

1000000

-150 -125 -100 -75 -50 -25 0 25 50Temperature, oC

Ela

stic

Mod

ulus

, psi

Sample Nos. B3-B7Bi-linear Curve Fit

Bi-linear Curve FitE = -1340T + 481,000 psi (-150 to 25oC)E = -18,700T + 919,000 psi (25 to 50oC)

Figure 8. Bi-linear Curve Fit of TP283E Modulus

Figure 6. TP283E/IM9 Tow #29 DMA Tg Results

Tg = 57.55oC


The Aerospace Corporation also tested the coefficientof thermal expansion (CTE) of the TP283E neat resin.CTE measurements were made on three TP283E epoxysamples over the temperature range of -100 to +50°Cusing a TA Instruments, Inc., Thermal MechanicalAnalyzer (TMA). The three samples were preparedwith the sample length parallel to the length, width, andthickness, respectively, of the ILC TP283E cast panel.The thermal expansion curves for the three sampleswere essentially identical and the data for the 3 sampleswere combined into one file. Various methods of curvefitting the CTE data for the 3 samples were attemptedwith the best curve fit obtained by two 2nd orderpolynomial equations. A curve fit of the data from -100to +20°C gave a CTE = 0.135T + 59.35 ppm/°C. Acurve fit of the data from +20 to +50°C gave a CTE =2.91T – 2.43 ppm/°C (Fig. 9).

Rule of Mixtures PredictionsRule of mixtures (ROM) CTE and modulus calculationswere made for unidirectional IM9/TP283E compositesusing the fiber properties listed in Table 1 and theTP283E epoxy matrix modulus and CTE values givenby the aforementioned equations. The ROM equationsused are listed as equations (1) and (2). The ROMpredictions were made over the temperature range of –100 to +50°C for fiber contents of 35, 40, 45, 50, 55,and 60 percent fiber volume fraction (Figs. 10 and 11).

As can be seen from Figure 10, the composite modulusis not affected appreciably by the modulus of the resin;instead it is fiber modulus and volume fractiondominated. Therefore, even though the modulus ofTP283E decreases significantly at 50oC, the compositemodulus does not. Figure 11 illustrates the point thatrigidizable materials can be made to exhibit near zeroCTE through proper material selection and fibervolume fraction control. It is apparent from this graphthat composite CTE is more dependent on the modulusof the resin than is composite modulus.

Tow Tensile Test ResultsThe Aerospace Corporation also mechanicallycharacterized composite tow samples during this study.The tow samples were manufactured by ILC using thesame methods as the isogrid booms. All samplesdiscussed herein are made from TP283E/IM9 (48K)carbon fiber tows. The fiber volume fraction of eachsample was measured non-destructively by accuratelymeasuring the mass and length of each sample and thenusing the resin and fiber densities to calculate the

TP283E Thermal Expansion

-8000

-6000

-4000

-2000

0

2000

4000

-150 -100 -50 0 50 100Temperature, oC

(L -

L23

)/L 2

3,

ppm

Data for 3 SamplesCurve Fits

CTE = 0.135T + 59.35 ppm/oC(-100 to 20oC)

CTE = 2.91T - 2.43 ppm/oC(20 to 50oC)

Figure 9. Curve Fit of Avg. TP283E CTE Results

ROM CTE Predictions for Unidirectional IM9/TP283E

-0.5

-0.3

-0.1

0.1

0.3

0.5

0.7

0.9

1.1

-100 -75 -50 -25 0 25 50Temperature, oC

Com

posi

te C

TE, p

pm/o C

35 vol.% Fiber 40 vol.% Fiber45 vol.% Fiber 50 vol.% Fiber55 vol.% Fiber 60 vol.% Fiber

Ef = 42 msi, CTEf = -0.5 ppm/oCEm = -0.00134T + 0.481 msi (-150 to 25oC); = -0.0187T + 0.919 msi (26 to 50oC)

CTEm = 0.135T + 59.35 ppm/oC (-100 to 25oC); = 2.91T - 2.43 ppm/oC (26 to 50oC)

Figure 11. Rule of Mixtures CTE Predictions

ROM Elastic Modulus Prediction for Unidirectional IM9/TP283E

0

5

10

15

20

25

30

-150 -125 -100 -75 -50 -25 0 25 50Temperature (oC)

Com

posi

te M

odul

us

(Msi

)

Vf=35% Vf=40%Vf=45% Vf=50% Vf=55% Vf=60%

Em = -1340T + 481,000 psi (-150 to 25oC)Em = -18,700T + 919,000 psi (25 to 50oC)

Ef = 42 Msi

Figure 10. Rule of Mixtures Modulus Predictions

E c E f v f⋅ E m v m⋅+:= (1)

α cα f E f⋅ v f⋅ α m E m⋅ v m⋅+

E f v f⋅ E m v m⋅+:= (2)

where αc = axial composite CTEαf and αm = fiber and matrix CTEEf,Em,Ec = fiber, matrix, and axial composite modulusvf and vm = fiber and matrix volume fraction


volume fraction. Thefiber volume fractionswere recorded prior totesting in order tonormalize the data and tocorrelate it with thepredicted ROM values.After manufacturing, thetow samples were bondedinto G10 grip tabs asshown in Figure 12.

The tow tensile tests were performed using an Instronuniversal testing machine set up with 1,000-lb gripswith 2-inch long serrated grip faces and a 1,000-lb loadcell. An extensometer was placed on each sampleduring testing to measure strain. The samples weretested at a crosshead rate of 0.1 in./min (strain rate ≤0.033 in./in./min). Young’s modulus, ultimate tensilestrength, and failure strain were measured. Young’smodulus, E, was calculated from a linear regression ofstress versus strain data from 0.05 to 0.50% strain. Theposition of the extensometer was varied during thetesting. The gauge length of the extensometer was 1.0-inches. All samples were loaded to 400-lb (130 ksi,0.8% strain) and unloaded twice for the initial modulusmeasurements. The modulus test values were thennormalized to 60% fiber volume fraction and comparedto ROM predictions.

Six pristine tows (no folding and deployment cycles)were tensile tested during this study (Table 2). Whennormalized to 60% fiber volume fraction, the averageroom temperature modulus of the pristine TP283E/IM9(48K) twisted tows was 24.8-Msi ± 0.3-Msi, whichcompares well with the rule of mixtures in Figure 10.

Table 2. TP283E/IM9 (48K) Pristine Tow Tensile ResultsNo.# Vf E, Measured,

MsiE, Normalized to

60% Vf

% ROM Em =0.5

3S 0.385 15.5 23.3 943R 0.385 15.3 23.0 933R 0.385 15.4 23.1 936R 0.397 16.4 24.8 976S 0.397 16.4 24.8 976R 0.397 16.4 24.8 979R 0.380 15.8 24.9 979S 0.380 15.9 25.1 989S 0.380 15.8 24.9 977aE 0.394 16.2 24.7 967aR 0.394 16.0 24.4 957aR 0.394 16.4 25.0 977bS 0.411 17.3 25.3 997bR 0.411 17.0 24.8 977bR 0.411 16.9 24.7 961a 0.337 14.1 25.1 9730 0.465 18.9 24.4 95

#Extensometer on smooth side of tow (S), rough side (R), or on edge (E) betweensmooth and rough sides

Five of the six pristine tensile samples were tested tofailure and had tensile strengths within 11% of the fiberultimate tensile strength (UTS). Samples 3, 6, 9, 7 and1a failed at 315-ksi, 343-ksi, 305-ksi, 311-ksi, and 320-ksi, respectively.

The Aerospace Corporation also investigated the effectof temperature on the tensile modulus of the towsamples (Fig. 13). The results of this experiment showthat over the measured range the composite tow tensilemodulus is not effected by temperature, which agreeswell with the rule of mixtures predictions in Figure 10.Note that some bending occurred at low loads at 80 and96°C and was most likely due to the weight of theextensometer.

The effect of folding oncomposite tow sampleswas also examined.After being processed attheir initial settemperature, towsamples were heated to atemperature of 100-105oC and folded 180degrees over a 5/32-inch(4-mm) radius as seen inthe X-ray photograph inFigure 14. The sampleswere constrained in the folded position as they cooled.Once cooled to room temperature, they were releasedfrom the constraining mechanism. Samples were thenheated again and allowed to return to shape via theshape memory recovery force of the material (Fig. 15).This process was repeated 1, 3, or 5 times in order todetermine the possible effects of packaging anddeploying structures fabricated from these materials.Note that the nano-X-ray of the TP283E/IM9 (48K) towshown in Figure 14, which was folded one time over a15/32-inch radius, revealed no fiber damage.

Figure 12. Tow TensileTest Samples

0

20

40

60

80

100

120

140

160

0.0 0.2 0.4 0.6 0.8 1.0Tensile Strain, %

Ten

sile

Str

ess,

ksi

48C62C80C96C29C

Figure 13. Tensile Stress vs. Strain: Sample 10a

Vf=48%

Figure 14. X-ray of FoldedTP283E/IM9 (48K) Tow


Tensile testing wasperformed on fivecomposite tow samples thatwere folded and deployedone time (Table 3). Whennormalized to 60% fibervolume fraction, theaverage room temperaturemodulus of theTP283E/IM9 (48K) towsafter being packed anddeployed once over a 5/32-inch radius is 24.6 MSI ±2.5 MSI. This value isclose to the rule of mixturesprediction in Figure 10, andagrees with data collectedusing pristine tows, indicating that folding one timeover a 15/32-inch radius has no effect on the tensilemodulus of the composite tow. Two of these sampleswere tensile tested to failure. Sample 12 failed at 338-ksi (99% fiber UTS) while sample 13 failed at 292-ksi(78% fibers UTS).

Table 3. TP283E/IM9 (48K) Tow Tensile Results; 1 FoldNo.# Vf E, Measured,

msiE, Normalized to

60% Vf

% ROMEm = 0.5

11-OD 0.467 19.7 25.3 9912-OD 0.390 17.2 26.5 10312-ID 0.390 14.1 21.7 8512-E 0.390 16.9 26.0 10112-E 0.390 16.8 25.8 101

13-OD 0.428 18.9 26.5 10313-ID 0.428 15.7 22.0 8613-E 0.428 17.7 24.8 9713-E 0.428 18.3 25.7 100

15-OD 0.389 17.6 27.1 10615-ID 0.389 12.6 19.4 7615-E 0.389 17.2 26.5 103

16-OD 0.446 19.7 26.5 10416-ID 0.446 15.1 20.3 7916-E 0.446 18.4 24.8 97

# OD=extensometer on outside of tow fold diameter; ID=extensometer on inside of tow folddiameter; E=extensometer on edge between OD and ID sides

The results indicate that the composite tow tensilemodulus of the folded samples varies with theextensometer location relative to the fold direction.The modulus measured on the fold edge matchedextremely well with the ROM predictions, while themeasured modulus was much lower on the inside edge(Fig. 16).

It was determined that the folding and deploymentprocess using shape memory resulted in fiber kinks onthe compression side of the fold, which causes a largedisplacement on the inside of the fold at low loads. Theoutside of the tows did not have the same displacementeffects because the fibers are placed in tension in thisarea of the tow. The kinks in the fiber on the inside of

the fold were removed at approximately 30-ksi, whichis a high stress relative to the expected loads in actualapplication. Tow samples from tubes that weredeployed via inflation were also tested and did notexhibit this fiber kinking behavior due to the relativelyhigh stress of the inflation pressure as compared to theshape memory recovery stress of the resin.

In order to determine the stress required to remove thefiber kinks of the packed and deployed tows at differenttemperatures, one composite tow sample was tensiletested at temperatures up to 96oC (Fig. 17). The resultsindicate that fiber kinks are removed at lower stressesas the temperature increases. Also, high temperatureloading permanently removes fiber kinks as illustratedby the nearly linear stress-strain curve at 24°Cfollowing loading at 96°C. Note that large deformationsat low loads at high temperatures may be due to theweight of the extensometer. These results indicate thatit is possible to remove any fiber kinks in the tows ofthe isogrid boom by carefully controlling thedeployment temperature and the inflation pressure.

0

20

40

60

80

100

120

140

160

0.0 0.2 0.4 0.6 0.8 1.0Tensile Strain, %

Ten

sile

Str

ess,

ksi

Fold Outside

Fold Inside

Fold Edge

E = 16.9 msi (102% ROM)extensometer on fold edge

Figure 16. Tensile Stress vs. Strain: Sample 12,Packed and Deployed Via Shape Memory Once

0

20

40

60

80

100

120

140

160

0 0.2 0.4 0.6 0.8 1Tensile Strain, %

Ten

sile

Str

ess,

ksi

35C60C82C96C24C

Extensometer on ID side of fold

Figure 17. Tensile Stress vs. Strain at ElevatedTemp.; Tow 16, 1 fold and deployment cycle

Figure 15. TP283E/IM9(48K) Tow SamplesReturned to Shape

After Folding


Three composite tow samples were tensile tested afterbeing packed and deployed three times (Table 4). Theresults indicate that there are fiber kinks on the inside ofthe fold similar to that seen in the results for one foldand deployment cycle (Fig. 18). The average roomtemperature tensile modulus of the three TP283E/IM9(48K) tows after three folds is 23.7 MSI ± 1.8 MSIwhen the data is normalized to 60% fiber volumefraction. All three samples were tensile tested tofailure. Samples 31, 32, and 33 failed at 340-ksi (91%fiber UTS), 389-ksi (101% fiber UTS), and 392-ksi(103% fiber UTS), respectively. The results of thesetests indicate that there is no degradation in tensileproperties from three 180-degree folding anddeployment cycles over a 15/32-inch radius.

Table 4. TP283E/IM9 (48K) Tensile Test Results; 3 foldsNo.# Vf E, Measured,

msiE, Normalized

to 60% Vf

% ROMEm = 0.5

31-OD 0.427 18.2 25.6 10031-ID 0.427 15.2 21.4 8331-E 0.427 16.7 23.5 92

31-ID-2 0.427 15.4 21.6 8532-OD 0.441 18.8 25.6 10032-ID 0.441 15.5 21.1 8232-E 0.441 17.8 24.2 95

33-OD 0.434 18.8 26.0 10233-ID 0.434 17.1 23.6 9233-E 0.434 17.5 24.2 95


Three composite tow samples were tensile tested afterbeing packed and deployed five times (Table 5). Theresults show that there is residual bending on the insideof the fold similar to that seen in the results for one foldand three fold and deployment cycles (Fig. 19). Theaverage room temperature tensile modulus of the threeTP283E/IM9 (48K) tows after five folds is 23.9 MSI ±1.8 MSI when the data is normalized to 60% fibervolume fraction.

Table 5. TP283E/IM9 (48K) Tensile Test Results; 5 folds

No.# Vf E, Measured,msi

E, Normalizedto 60% Vf

% ROMEm = 0.5

34-OD 0.431 18.6 25.9 10134-ID 0.431 16.9 23.5 9234-E 0.431 18.0 25.1 98

35-OD 0.484 20.3 25.2 9935-ID 0.484 16.7 20.7 8135-E 0.484 19.3 23.9 94

36-OD 0.450 18.9 25.2 9936-ID 0.450 16.1 21.5 8436-E 0.450 18.4 24.5 96


All three samples were tensile tested to failure.

Samples 34, 35, and 36 failed at 380-ksi (100% fiberUTS), 324-ksi (77% fiber UTS), and 381-ksi (97% fiberUTS), respectively. The results of these tests indicatethat there is no degradation in tensile modulus from five180-degree folding and deployment cycles over a15/32-inch radius. However, one tow (#35) had visualfiber damage after five folds and experienced a 25%reduction in tensile strength.

Tow Flexural Modulus Test ResultsThe composite tow flexural modulus tests wereperformed using the 3-point bending fixture of the(DMA) using a similar procedure as that used to test theneat resin samples. The tow used for this test had afiber volume fraction of 59% and was cut from the endof the isogrid test tube (#042) characterized by JMU.The results in Figure 20 show that the flexural modulusof the composite tow sample decreases 86% at 100oC,which is a similar result to that obtained during the neatTP283E resin flexural modulus testing. Table 6 showsthe strain ranges that were used to calculate themodulus values at the different test temperatures.

0

100

200

300

400

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2Tensile Strain, %

Ten

sile

Str

ess,

ksi

Fold Outside

Fold Inside

Fold Edge


UTS = 380 ksi (99% ROM)

Figure 19. Tensile Stress vs. Strain: Tow 34; 5 folds

0

100

200

300

400

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2Tensile Strain, %

Ten

sile

Str

ess,

ksi

Fold Outside

Fold Inside

Fold Edge


UTS = 389 ksi (99% ROM)

Figure 18. Tensile Stress Vs. Strain: Sample 32; 3folds


Table 6. Tow Flexural Modulus vs. TemperatureTemp. oC E, Msi Strain Range, %

27 13.1 0 - 0.5-150 13.8 0 - 0.5-100 13.6 0 - 0.5-50 13.4 0 - 0.523 12.6 0 - 0.550 8.8 0 - 0.175 4.1 0 - 0.05

100 1.8 0 - 0.05

Tow CTE Test ResultsPMIC of Corvallis, Oregon tested several TP283E/IM9(48K) composite tow samples to determine their CTEvalues. Tow number 40 (Vf=61%), which was cut fromthe end of boom #042 had a CTE of 0.3 ppm/oC (-100o

to 20oC) and 0.4 ppm/oC (-20o to 40oC). Anothercomposite tow sample (# 012-180+4, Vf=55%) was cutfrom an isogrid boom and was tested in to understandsome of the tube level effects on CTE. This tow had aCTE of = 0.7 ppm/oC (-100o to –20oC) and 1.0 ppm/oC(-20o to 40oC). Although both of these tests values areextremely low and are be acceptable for mostapplications, they are slightly higher than the ROMpredictions of Figure 11. A number of root causes wereidentified during this testing that directly affect the CTEof the tows. These fundamental issues will beoptimized in the future in order to achieve lower CTE’s,which will be closer to the ROM predictions.

Isogrid Boom TestingTwo booms were fabricated for this study. One boom(#042) was loaded in compression, tension, tested forcreep, and then failed in compression at an elevatedtemperature. The other boom was tested for CTE (#25).A section of another boom (#12) was also tested forCTE. The isogrid design chosen for this research wasthe Inflatable Solar Array Experiment II (ISAE-II)

baseline design. This particular configuration waschosen to allow for comparison between isogrid boomsfabricated by ILC in the past with both thermoplasticand UV epoxy resins. The fiber volume fractions of thebooms were calculated based on the average value fromfive tows cut from the ends of the booms aftermanufacturing. The properties of the booms are shownin Table 7. Like the composite tow samples, the isogridbooms were fabricated from TP283E epoxy resin and48K IM9 carbon fibers. The booms were manufacturedusing a modified filament winding technique. Theactual test booms are show in Figures 3 and 21.

Table 7. Isogrid Boom Design Dimensions Parameter Value

Lead Angle 30o

Lead of the Helix 12.697 inchesIsogrid Base Length 1.374 inchesBoom Inner Diameter 7.05 inchesNumber of Longitudinals 16Number of Helicals 16Tow Size (IM9) 48KAverage Tow Diameter 0.0534 inchesResin TP283EMass of Boom #042 199.5 gramsLength of Boom #042 48.375 inchesBoom Test Length #042 46.75 inchesAverage Fiber Volume Fraction #042 57.8%

A laser interferometer was used to characterize the CTEof the isogrid tube (#25) fabricated for this study alongwith a tube sample (#12)extracted from a longerisogrid boom. Tube #25was first tested in thepristine condition. Theapproximate average CTE(–40oC to +40°C) in thepristine condition of tube#25 is 1.15 ppm/°C. AfterZ-folding tube #25 over a15/32-inch radius one timeand deploying it usinginflation pressure, the boomhad an average CTE (-40oCand +40°C) ofapproximately 1.15 ppm/°C. Therefore it appears thatfolding has no effect on the CTE of the boom, butfurther testing is required to verify this initial result,especially with respect to the orientation of the fold lineof the tube during the testing.

The results from the section of isogrid boom (#12) cutfrom a longer boom are slightly lower than tube #25.After four thermal cycles, the average CTE (-40 and+40°C) is approximately 0.82 ppm/°C, while themaximum CTE was approximately 0.95 ppm/°C afterthe fourth thermal cycle (Fig. 22). From these results, it

Flexural Stress-Strain CurveIM9 (48K)/TP283E Tow No. 37

0

20

40

60

80

100

0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 0.008Flexural Strain, in/in

Flex

ural

Str

ess,

ksi 27C -150C

-100C -50C 23C 50C 75C 100C

Analysis for round cross sectionFiber Content = 59 vol.%

Tow extracted from tube prior to patch application

Figure 20. Tow Flexural Stress vs. Strain

Figure 21. Isogrid TubeDuring CTE Testing


is apparent that thermal cycling stabilizes the CTE ofthe isogrid boom. Like the tow CTE testing, the resultsof these tests revealed a number of issues that will beoptimized in the future in order to further reduce theCTE of the structure, thereby making it closer to theROM predictions in Figure 11.

The test stand used to obtain the tensile andcompressive elastic modulus, as well as the preliminarycreep data is described in detail elsewhere4. The boomis mounted horizontally on the test frame inside an 18-inch diameter by 52-inch long radiant heater. Axial andcompressive loads are applied using a computercontrolled, servomotor driven, linear motion system.Fiber optic strain gages were used because they werebetter suited for mounting to the ribs of the isogridstructure than typical bonded resistance type straingages. Four strain gages were mounted 90o apart at thecenter of the boom. The use of fiber optic strain gagesand associated signal conditioning necessitated usingtwo different data acquisition systems, one for the straingages and one for the load cell and thermocouples.Twelve thermocouples were used to measuretemperature. They were arranged in three groups offour gages. The gage groups were locatedapproximately six inches from each end of the tube andin the center. The four gages at each location werelocated 90° apart. For elastic modulus testing and creeptesting the tube was loaded with approximately 50-lbf(1400-psi). For elastic modulus testing a cross-headrate of 0.05-in/min was used to apply the load. The loadwas then held and the tube allowed to relax until thestrain gages stabilized. Tensile and compression testswere performed between 25-64oC. When datacollection was completed at the test temperature, thetube was unloaded and allowed to cool. Betweenelevated temperature tests, a room temperature test wasperformed to ensure that there were no structuralchanges in the tube caused by the heating. The results

from the tensile and compression modulus testing areshown in Figures 23 and 24.

The average tensile modulus of the boom from 28-47oCis 24.5-Msi ± 0.3-Msi. Normalized to 60% fibervolume fraction, the tensile modulus of the boom is25.5-MSI. These compressive modulus test valuescompare favorably (within 1%) with the predictedROM modulus as seen in Figure 10. The averagecompressive modulus of the boom from 25-58oC is21.8-Msi ± 1.5-Msi. Normalized to 60% fiber volumefraction, the compressive modulus of the tube was 22.6-Msi. This value also compares favorably (within 10%)with the rule of mixtures predictions in Figure 10.

The behavior of the boom changes dramatically atapproximately 55-59oC. At these temperatures, forboth the compressive and tensile cases, the boomexhibited viscoelastic behavior. As seen in Figure 25for the compressive loading case at 64oC, strainrelaxation occurs where the load was held at 50-lbf.This behavior occurs because of the large drop inmodulus of the resin above Tg. At temperatures above55oC the modulus becomes temperature and timedependent. These tests indicate that the viscoelastic

CTE ≅ 0.82 ppm/°CCTE ≅ 0.95 ppm/°C

Figure 22. Isogrid CTE Results; Tube #012

Figure 23. Isogrid Tensile Modulus Results

Tensile Modulus Vs. Temperature IM9 (48K)/TP283E Isogrid Boom #042

0

5

10

15

20

25

30

25 30 35 40 45 50Temperature (OC)

Mod

ulus

(MSI

)

Vf=57.8

Figure 24. Isogrid Compression Modulus Results

Compression Modulus Vs. Temperature IM9 (48K)/TP283E Isogrid Boom #042

05

1015202530

25 35 45 55Temperature (OC)

Mod

ulus

(MSI

) Vf=57.8%


behavior of the boom must be considered during designand deployment.

Since understanding the viscoelastic behavior of theboom is fundamental to understanding boomperformance, preliminary creep tests were performed.Strain relaxation data were obtained at fourtemperatures between 46oC and 65 °C. The timeconstant for the relaxation at each temperature wasobtained from equation (3).

where ε(t) is the strain at any instant in time, ε∞ is thestrain at steady state, t is time, and τ is the timeconstant. A plot of the time constant vs. 1/T is shownin Figure 26.

The trend line represents the behavior of the materialpredicted by the Arrhenius equation (4).

Where T is the temperature in Kelvin, and RE∆

is a

constant related to the activation energy and has unitsof K. The Arrhenius equation can be used to predict thecreep behavior of the boom outside of the temperaturerange measured. For example, a typical operatingtemperature for the boom is 10°C (283 K). Thepredicted time constant at this temperature is 94,504seconds or 26.25 hours. This is the time required for63% of the strain relaxation to occur. 98% of the strainrelaxation will occur at 4 time constants or 105 hours.

The final boom testperformed at JMU was thecompression to failure atelevated temperature test.The expected maximum usetemperature of a boom ofthis construction is 35oC.With a safety factor of 5oC,it was decided that the tubeshould tested atapproximately 40oC. Forthis test the tube was loadedin compression at a strainrate of 0.05 in/min. The tubefailed in rib buckling at 418.2-lbf, 11681.6-psi incompression at 40.6oC (Figure 27). These resultscompare favorably to the predictions obtained by usingILC’s isogrid analytical model1,3. Using the inputs inTable 9 and the ROM modulus at 57.8% Vf of 24.5-Msi, the predicted compressive failure load is 443-lbf inrib buckling, which is 5.6% above the actual load.

Summary and ConclusionsThe results of the testing discussed herein clearlyillustrate the excellent agreement between rule ofmixtures mechanical predictions and ILC’s rigidizablecomposite isogrid structure test values in thetemperature range of approximately –40oC to 55oC.The isogrid boom exhibits viscoelastic behavior abovethose temperatures. The high failure load of theTP283E/IM9 (48K) isogrid boom of 418-lbf at 40oCalso shows that the booms are structurally sound up tothe expected use temperature of 35oC. ILC’s isogridanalytical model predicts the failure load accurately attemperatures of at least 40oC. Isogrid technology is alsoscalable to tens and hundreds of feet in length and hasbeen demonstrated in continuous lengths up to 23 feet(Fig. 28). The results of the composite tow tensile testsindicate that the modulus the TP283E/IM9 composite is

Micro-Strain vs. Time for Compression Test at 64oC for Isogrid Boom #042; Vf=57.8%

-30

-25

-20

-15

-10

-5

0

5

0 50 100 150 200 250 300

Time (s)

Mic

ro-S

trai

n (a

vg. o

f 4 g

ages

)

At 234.5 seconds the load was held at 50-lbf and the boom was allowed to relax

Constant Load

Strain Fluctuations correspond to cycling of heaters

Figure 25. Isogrid Boom Strain Relaxation Curve

(3)

−=

−

∞τεεt

et 1)(Figure 27. IsogridBoom #042 During

Compressive Failure

(4)RTE

oe∆

−= ττ

τ= 2E-19e15427/T

R2 = 0.9836

1

10

100

1000

0.00295 0.003 0.00305 0.0031 0.003151/T (1/K)

τ (s

)

Figure 26. Relationship between the Time Constant forStrain Relaxation and Temperature


unaffected by up to five 180 degree folding anddeployment cycles over a 15/32-inch radius. Futuretesting will be conducted to reduce this fold radius.These tests also indicate that theultimate tensile strength of the towsis within 25% of the ultimate tensilestrength of the IM9 fibers. This tightfolding can be achieved by the largereduction in flexural modulus of thetows at elevated temperatures (65%reduced at 75oC, 80% reduced at100oC, and 95% reduced at 125oC).These folding tests indicate the highpacking efficiency that can beobtained with ILC’s isogrid boomswith no reduction in tensile modulus.The CTE tests indicate that ILC’sTP283E/IM9 isogrid booms have anear-zero CTE of approximately 1ppm/oC. The initial results of thesetests also indicate that there are noapparent folding effects on the CTEof isogrid. Root causes at thestructural and tow level wereidentified to reduce the CTE of thebooms in the future and bring themcloser to rule of mixtures thermalpredictions. The strain relaxationtesting on the isogrid booms alsoindicated that the structure willbecome stable to creep at 10oC atapproximately four days afterdeployment.

AcknowledgmentsThe authors thank Mr. Jim McManus from LunaInnovations, Inc. for the donation of the straindisplacement devices for the thermo-mechanical testingof the isogrid boom at JMU. The authors also thankJPL and DARPA for supporting the composite testingat The Aerospace Corporation.

References1. Cadogan, D.P., Lin, J.K, Sapna, G.H.,Scarborough, S.E., “Space Inflatable TechnologyDevelopment for Solar Sails and Other GossamerApplications: GR/SMP Isogrid Boom DevelopmentFinal Report,” NASA Task Order 10442, ILC Dover,Inc., October, 2001.

2. Darooka, D.K., S.E. Scarborough, and D.PCadogan, “An Evaluation of Inflatable Truss Frame ForSpace Applications,” AIAA-2001-1614, 42nd

AIAA/ASME/ ASCE/AHS/ASC Structures, StructuralDynamics, and Materials Conference and ExhibitAIAA Gossamer Spacecraft Forum, April 16-19, 2001.

3. Lin, J.K., G.H. Sapna, Cadogan, D.P., S.E.Scarborough, “Inflatable Rigidizable Isogrid BoomDevelopment,” AIAA-2002-1297, 43rd AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, andMaterials Conference and Exhibit AIAA GossamerSpacecraft Forum, April 22-25, 2002.

4. Pederson, L.M., Blandino, J.R., Woods-Vedler,J.A., “Determination of the Modulus of Elasticity as aFunction of Temperature for an Isogrid Tube,” AIAA2002-1334, Proceedings of the 43rd

AIAA/ASME/ASCE/AHS/ASC Structures, StructuralDynamics and Materials Conference, Denver, CO, 22-25 April, 2002.

5. Cadogan, D.P. and S.E. Scarborough “RigidizableMaterials for use in Gossamer Space InflatableStructures,”42nd AIAA/ASME/ASCE/AHS/ASCStructures, Structural Dynamics, and MaterialsConference & Exhibit AIAA Gossamer SpacecraftForum, Seattle, WA, April 16-19, 2001.

6. Mikulas, M.M., Jr., “Structural Efficiency of Long,Lightly Loaded Truss and Isogrid Columns for SpaceApplications,” NASA Technical Memorandum 78687,July 1978.

7. Darooka, D.K., S. Scarborough, S. Malghan, D.Cadogan, C. Knoll, “Inflatable Space Frame,” FinalReport, NASA Prime Contract Number: NAS1-99154,July 2000.

8. Cadogan, D.P., S.E. Scarborough, J.K. Lin, G.H.Sapna, “Shape Memory Polymer CompositeDevelopment For Use in Gossamer Space InflatableStructures,” AIAA-2002-1372, 43rd AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, andMaterials Conference and Exhibit AIAA GossamerSpacecraft Forum, April 22-25, 2002.

9. Hexcel Magnamite IM9 Carbon Fiber ProductData Sheet, March 2002.

Figure 28. 7”diameter, 23

foot long,Isogrid Boom

in test atNASA-LaRC

elevated temperature mechanical_characterization_of_isogrid_booms

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