Elevated temperature mechanical_characterization_of_isogrid_booms

ELEVATED TEMPERATURE MECHANICAL CHARACTERIZATION OF ISOGRID BOOMS
Stephen E. Scarborough* and David P. Cadogan† – ILC Dover, Inc., Frederica, DE
Lauren M. Pederson‡ and Joseph R. Blandino§* – James Madison University, Harrisonburg, VA
Gary L. Steckel∋ and Wayne K. Stuckey¶ – The Aerospace Corporation, El Segundo, CA
1
Abstract
Structurally efficient isogrid booms, manufactured
from rigidizable composite materials, are becoming an
enabling technology for spacecraft structures because
of their high packing efficiency. Selection of the
materials used in the construction of rigidizable space
structures is commonly driven by mechanical
performance properties at elevated temperatures.
Mechanical properties testing was performed on
composite tow samples and on an isogrid boom at
various temperatures. To characterize elevated
temperature behavior, the isogrid booms, and its sub-element
composite tows were manufactured from
ILC’s TP283E shape memory polymer (SMP) matrix
resin and a carbon reinforcement. Both the flexural
modulus and the tensile modulus of the composite tow
samples were determined as a function of temperature.
These values were compared to the calculated values
for the composite based on rule of mixtures analysis.
The predicted rule of mixtures composite modulus is
used in ILC’s isogrid analytical code to predict the
structural properties of the isogrid boom. A number of
composite tow samples were fabricated by ILC and
mechanically characterized by the Aerospace
Corporation to gather independent performance data.
An isogrid boom was fabricated by ILC and
mechanically characterized at elevated temperatures by
James Madison University (JMU). JMU tested this
boom in tension, compression, and also performed
preliminary creep testing at various temperatures. A
similar isogrid boom was fabricated by ILC and tested
by The Aerospace Corporation for composite CTE
performance. This paper discusses the results of both
the composite tow testing and the isogrid boom testing
in pre- and post-packing conditions. A discussion of
the correlation between the predicted values and the
actual test values is also presented.
Introduction
NASA and DoD space missions in the near future will
require much larger satellites, the sizes of which will
be beyond the capabilities of current technologies. The
types of Gossamer
spacecraft that will be
needed include
antennas, solar arrays,
sunshields, solar sails,
and telescopes (Figs.
1-2). Some systems
being considered are
hundreds of meters in
size to accomplish
mission goals. Due to
the increase in payload
size required,
innovative support
structures, which can
be packed into the
faring of available
launch vehicles, must
be developed. In recent
years, research and
development work has
been performed in this
area1-3. Of the
available options, one
of the most promising
technological advancements is the rigidizable
inflatable structure. A rigidizable inflatable structure is
one that is fabricated on Earth, packed into the launch
container, and inflated for deployment once on orbit1,3.
After deployment, the material is rigidized, or
hardened, to form a stiff composite structure that no
longer needs the inflation gas for support. This class of
structures has unique benefits such as low packing
volume, reduced mass, and in most cases, very high
deployed structural efficiency3,4.
Several types of construction can be used in a
rigidizable inflatable including monocoque, isogrid,
IsoTruss, and truss-frame booms. Each composite
structure can be fabricated into a varying geometric
shapes utilizing any number of resin and fiber types5.
The fibrous reinforcement can be in tow or woven
fabric form. In order to optimize the structure, the sizes
of the tows and the weave styles of the fabrics can be
varied5. It is also possible to manufacture near-zero
coefficient of thermal expansion (CTE) booms through
the fiber and resin selection and by optimizing the
volume fractions of each5. However, key to all
mechanical performance properties is the ability to fold
and 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
American Institute of Aeronautics and Astronautics
Figure 1. ½ Scale Next
Generation Space
Telescope Sunshield
Figure 2. ILC 3.2m Diameter
TSU Hexapod Testbed

2
Isogrid Design and Construction
One of the most advanced inflatable rigidizable
structures is the isogrid boom, which consists of a grid-work
of equilateral triangles1,3,6. These equilateral
triangles give the overall structure isotropic mechanical
properties1,3. One of the isogrid booms fabricated for
this study is shown in Figure 3 (patent pending). This
type of boom has a circular cross-section
and is encased on both sides
with a polymeric film such as
polyimide1,3. The inner film layer
acts as a bladder or gas-retaining
layer for inflation. The outer later,
called the anti-blocking layer, is used
to prevent the structure from
adhering (blocking) to itself when it
is in the packed configuration1,3. The
outer layer can also act as the first
layer in a mult-layered insulation
(MLI) blanket. ILC Dover has
developed analytical modeling
techniques to predict the properties
of this structure1,3,6. ILC has studied
the room temperature mechanical
properties of this structure
extensively, but there is still a need
for further work to be performed to
fully characterize the performance of
the structure, especially at elevated
temperatures1,3.
Materials
A leading rigidizable material candidate is the shape
memory composite 1-5, 7,8. The composite consists of a
fibrous reinforcement, such as carbon, and a polymeric
matrix resin such as polyurethane or epoxy. The resin
is initially consolidated at a high temperature, called
the set temperature, to form a high modulus, rigid
structure. This initial heating and consolidation event
defines the shape of the structure. Subsequent heating
events above the material’s glass transition
temperature (Tg) lowers the modulus of the material
significantly and allows the structure to be tightly
packaged (Fig. 4). If
constrained and cooled below
Tg while packed, the
composite modulus will
increase and the material will
retain the packed shape, even
if the package is
unconstrained. Upon re-heating,
the material
experiences a large decrease
in modulus and can therefore
be deployed back to its as-manufactured state. Once
the deployed structure cools below the Tg, the modulus
will increase and the inflation gas is no longer required
for structural support. This process of packing and
deploying the structure is repeatable which allows
flight hardware to be packed and deployed for
evaluation several times during ground test prior to
launch and deployment in space. The thermoplastic
resin 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 force
is weak relative to the force required to deploy the
composite structure and associated systems, therefore
this force must usually be augmented by inflation gas.
Material Properties
In this study, ILC Dover’s TP283E epoxy resin was
used to manufacture all of the test samples. The resin
system has reduced cross-link functionality in order to
add increased flexibility above the Tg as compared to
typical epoxy resin systems. TP283E therefore exhibits
thermoplastic behavior with the only exception being
that it does not have a melt temperature. ILC chose
Hexcel’s IM9 carbon fiber as the reinforcement for all
of the test samples discussed herein. IM9 was chosen
because of its balance of high mechanical properties
and high strain to failur1,3,8. The properties of this fiber
are listed in Table 1.
Table 1. IM9 (12K) Carbon Fiber Properties9
Property English Units SI Units
Tensile Modulus 42.0 x 106 psi 290Gpa
Tensile Strength 890,000 psi 6,141 MPa
Ultimate Elongation 2.1% 2.1.%
Axial CTE∝ -0.228 ppm/oF -0.5 ppm/oC
Density 0.0650 lb/in3 1.80 g/cm3
Filament Diameter 0.175 mil 4.4 microns
Filament shape Round Round
Weight/Length 18.8 x 10-6 lb/in 0.335 g/m
Fiber Cross-Sectional Area 2.89x10-4in2 0.19mm2
∝Estimate
In earlier studies, assumptions were made based on test
data from similar resins to predict the mechanical
properties of TP283E resin in order to eliminate the
overall amount of testing required in initial
development1,3,8. To determine the actual properties of
the resin using an independent test lab, a resin sample
was manufactured by ILC and sent to The Aerospace
Corporation for mechanical characterization. The
elastic modulus of the neat resin was tested using the 3-
point bending method on a TA Instruments Inc.,
Dynamic Mechanical Analysis (DMA) system. From
this testing, the average elastic modulus at 23°C for 5
samples of TP283E resin was determined to be 500 ±
20 ksi (Fig. 5).
Figure 3. ILC
Isogrid Boom:
Figure 4. Z-Folded
7” Diameter, 39”
Long Isogrid
Boom

3
2
1.8
1.6
1.4
1.2
1
0.8
0.6
0.4
0.2
Using the cantilever bending function on the DMA, The
Aerospace Corporation determined from the loss
modulus peak that the Tg of the resin was 55oC at a
frequency of 1 Hz. Three tow samples extracted from
an isogrid boom were also tested using DMA (Fig. 6).
The results from these tests indicate that the Tg is
between 57oC-61oC. Earlier Differential Scanning
Calorimetry (DSC) testing performed at ILC indicated
that TP283E had a Tg of 48oC1. Differences in the test
methods and slight deviations in the set temperatures
are the likely causes of the variation in the Tg results.
Therefore, depending on the set temperatures, TP283E
can be made to have a Tg in the range of 48-61oC.
Tg = 57.55oC
The flexural modulus of TP283E resin over the
temperature range from -150°C to +50°C was also
tested. Tests were performed on sample numbers
B3−B7, which were approximately 2.3-inches long,
0.35-inches wide, and 0.12-inches thick. They were
tested using the 2-in. span of the DMA 3-Point Bend
fixture. Peak loads were varied for the different test
temperatures because the neat resin samples had
increasingly lower moduli at higher temperatures.
The temperature sequence was 23, 30, 35, 40, and
50°C. At the end of each test the force was set to 0 N.
The force was then maintained at 0 N during heating to
the next test temperature and during the 5-minute
thermal equilibration at the test temperature. After the
5-minute thermal equilibration at the test temperature,
the static force was then set to 0.01 N and immediately
ramped to the final force. Good modulus data were
obtained for all test temperatures for sample B7 (Fig.
7). Modulus values for sample B7 were similar to those
for sample B6 for all test temperatures.
1.6
1.4
1.2
1.0
0.8
0.6
0.4
0.2
After the modulus testing was completed, the valid
modulus data for samples B3−B7 were tabulated and
plotted as functions of temperature. It was determined
that two linear curve fits could be used to describe the
data. A straight line with a relatively low slope (-1.3
ksi/°C) was fit to the data for -150 to 23°C and a
straight line with a much higher slope (18.7 ksi/°C) was
fit to the data for 23 to 50°C (Fig. 8). The two curves
intersect at 25°C.
1000000
900000
800000
700000
600000
500000
400000
300000
200000
100000
0
0.0 0.1 0.2 0.3 0.4 0.5
Flexural Strain, %
Flexural Stress, ksi
No. 1
No. 2
No. 3
No. 4
No. 5
E = 535 ksi
E = 501 ksi
E = 493 ksi
E = 490 ksi
E = 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 Curve
TP283E Sample No. B7
0.0
0 0.002 0.004 0.006
Flexural Strain, in/in
25C, E = 445 ksi
32C, E = 362 ksi
36C, E = 261 ksi
41C, E = 72 ksi
45C, E = 49 ksi
50C, E = 22 ksi
Figure 7. TP283E Elastic Modulus Vs. Temperature
TP283E Elastic Modulus versus Temperature
0
-150 -125 -100 -75 -50 -25 0 25 50
Temperature, oC
Elastic Modulus, psi
Sample Nos. B3-B7
Bi-linear Curve Fit
Bi-linear Curve Fit
E = -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

4
The Aerospace Corporation also tested the coefficient
of thermal expansion (CTE) of the TP283E neat resin.
CTE measurements were made on three TP283E epoxy
samples over the temperature range of -100 to +50°C
using a TA Instruments, Inc., Thermal Mechanical
Analyzer (TMA). The three samples were prepared
with the sample length parallel to the length, width, and
thickness, respectively, of the ILC TP283E cast panel.
The thermal expansion curves for the three samples
were essentially identical and the data for the 3 samples
were combined into one file. Various methods of curve
fitting the CTE data for the 3 samples were attempted
with the best curve fit obtained by two 2nd order
polynomial equations. A curve fit of the data from -100
to +20°C gave a CTE = 0.135T + 59.35 ppm/°C. A
curve fit of the data from +20 to +50°C gave a CTE =
2.91T – 2.43 ppm/°C (Fig. 9).
TP283E Thermal Expansion
CTE = 2.91T - 2.43 ppm/oC
(20 to 50oC)
Rule of Mixtures Predictions
4000
2000
0
-2000
-4000
-6000
Rule of mixtures (ROM) CTE and modulus calculations
were made for unidirectional IM9/TP283E composites
using the fiber properties listed in Table 1 and the
TP283E epoxy matrix modulus and CTE values given
by the aforementioned equations. The ROM equations
used are listed as equations (1) and (2). The ROM
predictions 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).
1.1
0.9
0.7
0.5
0.3
0.1
-0.1
-0.3
ROM Elastic Modulus Prediction for
Unidirectional IM9/TP283E
Ef = 42 msi, CTEf = -0.5 ppm/oC
30
25
20
15
10
5
Ef = 42 Msi
Em = -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)
As can be seen from Figure 10, the composite modulus
is not affected appreciably by the modulus of the resin;
instead it is fiber modulus and volume fraction
dominated. Therefore, even though the modulus of
TP283E decreases significantly at 50oC, the composite
modulus does not. Figure 11 illustrates the point that
rigidizable materials can be made to exhibit near zero
CTE through proper material selection and fiber
volume fraction control. It is apparent from this graph
that composite CTE is more dependent on the modulus
of the resin than is composite modulus.
Tow Tensile Test Results
The Aerospace Corporation also mechanically
characterized composite tow samples during this study.
The tow samples were manufactured by ILC using the
same methods as the isogrid booms. All samples
discussed herein are made from TP283E/IM9 (48K)
carbon fiber tows. The fiber volume fraction of each
sample was measured non-destructively by accurately
measuring the mass and length of each sample and then
using the resin and fiber densities to calculate the
:= (2)
-8000
-150 -100 -50 0 50 100
Temperature, oC
(L - L23)/L23, ppm
Data for 3 Samples
Curve Fits
CTE = 0.135T + 59.35 ppm/oC
(-100 to 20oC)
Figure 9. Curve Fit of Avg. TP283E CTE Results
ROM CTE Predictions for Unidirectional
IM9/TP283E
-0.5
-100 -75 -50 -25 0 25 50
Temperature, oC
Composite CTE, ppm/oC
35 vol.% Fiber 40 vol.% Fiber
Figure 11. Rule of Mixtures CTE Predictions
0
-150 -125 -100 -75 -50 -25 0 25 50
Temperature (oC)
Composite Modulus
(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)
Figure 10. Rule of Mixtures Modulus Predictions
E c := E f⋅v f + Em⋅v m (1)
α c
α f⋅E f⋅v f + α m⋅Em⋅v m
E f⋅v f + Em⋅v m
where αc = axial composite CTE
αf and αm = fiber and matrix CTE
Ef,Em,Ec = fiber, matrix, and axial composite modulus
vf and vm = fiber and matrix volume fraction

5
volume fraction. The
fiber volume fractions
were recorded prior to
testing in order to
normalize the data and to
correlate it with the
predicted ROM values.
After manufacturing, the
tow samples were bonded
into G10 grip tabs as
shown in Figure 12.
The tow tensile tests were performed using an Instron
universal testing machine set up with 1,000-lb grips
with 2-inch long serrated grip faces and a 1,000-lb load
cell. An extensometer was placed on each sample
during testing to measure strain. The samples were
tested at a crosshead rate of 0.1 in./min (strain rate ≤
0.033 in./in./min). Young’s modulus, ultimate tensile
strength, and failure strain were measured. Young’s
modulus, E, was calculated from a linear regression of
stress versus strain data from 0.05 to 0.50% strain. The
position of the extensometer was varied during the
testing. 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 modulus
measurements. The modulus test values were then
normalized to 60% fiber volume fraction and compared
to ROM predictions.
Six pristine tows (no folding and deployment cycles)
were tensile tested during this study (Table 2). When
normalized to 60% fiber volume fraction, the average
room temperature modulus of the pristine TP283E/IM9
(48K) twisted tows was 24.8-Msi ± 0.3-Msi, which
compares well with the rule of mixtures in Figure 10.
Table 2. TP283E/IM9 (48K) Pristine Tow Tensile Results
No.# Vf E, Measured,
160
140
120
100
80
60
40
20
Msi
E, Normalized to
60% Vf
% ROM Em =
0.5
3S 0.385 15.5 23.3 94
3R 0.385 15.3 23.0 93
3R 0.385 15.4 23.1 93
6R 0.397 16.4 24.8 97
6S 0.397 16.4 24.8 97
6R 0.397 16.4 24.8 97
9R 0.380 15.8 24.9 97
9S 0.380 15.9 25.1 98
9S 0.380 15.8 24.9 97
7aE 0.394 16.2 24.7 96
7aR 0.394 16.0 24.4 95
7aR 0.394 16.4 25.0 97
7bS 0.411 17.3 25.3 99
7bR 0.411 17.0 24.8 97
7bR 0.411 16.9 24.7 96
1a 0.337 14.1 25.1 97
30 0.465 18.9 24.4 95
#Extensometer on smooth side of tow (S), rough side (R), or on edge (E) between
smooth and rough sides
Five of the six pristine tensile samples were tested to
failure and had tensile strengths within 11% of the fiber
ultimate tensile strength (UTS). Samples 3, 6, 9, 7 and
1a failed at 315-ksi, 343-ksi, 305-ksi, 311-ksi, and 320-
ksi, respectively.
The Aerospace Corporation also investigated the effect
of temperature on the tensile modulus of the tow
samples (Fig. 13). The results of this experiment show
that over the measured range the composite tow tensile
modulus is not effected by temperature, which agrees
well with the rule of mixtures predictions in Figure 10.
Note that some bending occurred at low loads at 80 and
96°C and was most likely due to the weight of the
extensometer.
Vf=48%
The effect of folding on
composite tow samples
was also examined.
After being processed at
their initial set
temperature, tow
samples were heated to a
temperature of 100-
105oC and folded 180
degrees over a 5/32-inch
(4-mm) radius as seen in
the X-ray photograph in
Figure 14. The samples
were constrained in the folded position as they cooled.
Once cooled to room temperature, they were released
from the constraining mechanism. Samples were then
heated again and allowed to return to shape via the
shape memory recovery force of the material (Fig. 15).
This process was repeated 1, 3, or 5 times in order to
determine the possible effects of packaging and
deploying structures fabricated from these materials.
Note that the nano-X-ray of the TP283E/IM9 (48K) tow
shown in Figure 14, which was folded one time over a
15/32-inch radius, revealed no fiber damage.
Figure 12. Tow Tensile
Test Samples
0
0.0 0.2 0.4 0.6 0.8 1.0
Tensile Strain, %
Tensile Stress, ksi
48C
62C
80C
96C
29C
Figure 13. Tensile Stress vs. Strain: Sample 10a
Figure 14. X-ray of Folded
TP283E/IM9 (48K) Tow

6
Tensile testing was
performed on five
composite tow samples that
were folded and deployed
one time (Table 3). When
normalized to 60% fiber
volume fraction, the
average room temperature
modulus of the
TP283E/IM9 (48K) tows
after being packed and
deployed once over a 5/32-
inch radius is 24.6 MSI ±
2.5 MSI. This value is
close to the rule of mixtures
prediction in Figure 10, and
agrees with data collected
using pristine tows, indicating that folding one time
over a 15/32-inch radius has no effect on the tensile
modulus of the composite tow. Two of these samples
were 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 Fold
160
140
120
100
80
60
40
20
0
Fold Outside
Fold Inside
Fold Edge
E = 16.9 msi (102% ROM)
extensometer on fold edge
0.0 0.2 0.4 0.6 0.8 1.0
Tensile Stress, ksi
Figure 16. Tensile Stress vs. Strain: Sample 12,
Packed and Deployed Via Shape Memory Once
160
140
120
100
80
60
40
20
msi
E, Normalized to
60% Vf
% ROM
Em = 0.5
11-OD 0.467 19.7 25.3 99
12-OD 0.390 17.2 26.5 103
12-ID 0.390 14.1 21.7 85
12-E 0.390 16.9 26.0 101
12-E 0.390 16.8 25.8 101
13-OD 0.428 18.9 26.5 103
13-ID 0.428 15.7 22.0 86
13-E 0.428 17.7 24.8 97
13-E 0.428 18.3 25.7 100
15-OD 0.389 17.6 27.1 106
15-ID 0.389 12.6 19.4 76
15-E 0.389 17.2 26.5 103
16-OD 0.446 19.7 26.5 104
16-ID 0.446 15.1 20.3 79
16-E 0.446 18.4 24.8 97
# OD=extensometer on outside of tow fold diameter; ID=extensometer on inside of tow fold
diameter; E=extensometer on edge between OD and ID sides
The results indicate that the composite tow tensile
modulus of the folded samples varies with the
extensometer location relative to the fold direction.
The modulus measured on the fold edge matched
extremely well with the ROM predictions, while the
measured modulus was much lower on the inside edge
(Fig. 16).
It was determined that the folding and deployment
process using shape memory resulted in fiber kinks on
the compression side of the fold, which causes a large
displacement on the inside of the fold at low loads. The
outside of the tows did not have the same displacement
effects because the fibers are placed in tension in this
area of the tow. The kinks in the fiber on the inside of
the fold were removed at approximately 30-ksi, which
is a high stress relative to the expected loads in actual
application. Tow samples from tubes that were
deployed via inflation were also tested and did not
exhibit this fiber kinking behavior due to the relatively
high stress of the inflation pressure as compared to the
shape memory recovery stress of the resin.
Tensile Strain, %
In order to determine the stress required to remove the
fiber kinks of the packed and deployed tows at different
temperatures, one composite tow sample was tensile
tested at temperatures up to 96oC (Fig. 17). The results
indicate that fiber kinks are removed at lower stresses
as the temperature increases. Also, high temperature
loading permanently removes fiber kinks as illustrated
by the nearly linear stress-strain curve at 24°C
following loading at 96°C. Note that large deformations
at low loads at high temperatures may be due to the
weight of the extensometer. These results indicate that
it is possible to remove any fiber kinks in the tows of
the isogrid boom by carefully controlling the
deployment temperature and the inflation pressure.
0
0 0.2 0.4 0.6 0.8 1
Tensile Strain, %
Tensile Stress, ksi
35C
60C
82C
96C
24C
Extensometer on ID side of fold
Figure 17. Tensile Stress vs. Strain at Elevated
Temp.; Tow 16, 1 fold and deployment cycle
Figure 15. TP283E/IM9
(48K) Tow Samples
Returned to Shape
After Folding

7
Three composite tow samples were tensile tested after
being packed and deployed three times (Table 4). The
results indicate that there are fiber kinks on the inside of
the fold similar to that seen in the results for one fold
and deployment cycle (Fig. 18). The average room
temperature tensile modulus of the three TP283E/IM9
(48K) tows after three folds is 23.7 MSI ± 1.8 MSI
when the data is normalized to 60% fiber volume
fraction. All three samples were tensile tested to
failure. 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 these
tests indicate that there is no degradation in tensile
properties from three 180-degree folding and
deployment cycles over a 15/32-inch radius.
Table 4. TP283E/IM9 (48K) Tensile Test Results; 3 folds
400
300
200
100
0
Fold Outside
Fold Inside
Fold Edge
UTS = 380 ksi (99% ROM)
E = 18.0 msi (99% ROM)
0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2
Tensile Stress, ksi
Figure 19. Tensile Stress vs. Strain: Tow 34; 5 folds
UTS = 389 ksi (99% ROM)
msi
E, Normalized
to 60% Vf
% ROM
Em = 0.5
31-OD 0.427 18.2 25.6 100
31-ID 0.427 15.2 21.4 83
31-E 0.427 16.7 23.5 92
31-ID-2 0.427 15.4 21.6 85
32-OD 0.441 18.8 25.6 100
32-ID 0.441 15.5 21.1 82
32-E 0.441 17.8 24.2 95
33-OD 0.434 18.8 26.0 102
33-ID 0.434 17.1 23.6 92
33-E 0.434 17.5 24.2 95
400
300
200
100
Three composite tow samples were tensile tested after
being packed and deployed five times (Table 5). The
results show that there is residual bending on the inside
of the fold similar to that seen in the results for one fold
and three fold and deployment cycles (Fig. 19). The
average room temperature tensile modulus of the three
TP283E/IM9 (48K) tows after five folds is 23.9 MSI ±
1.8 MSI when the data is normalized to 60% fiber
volume fraction.
Table 5. TP283E/IM9 (48K) Tensile Test Results; 5 folds
msi
E, Normalized
to 60% Vf
% ROM
Em = 0.5
34-OD 0.431 18.6 25.9 101
34-ID 0.431 16.9 23.5 92
34-E 0.431 18.0 25.1 98
35-OD 0.484 20.3 25.2 99
35-ID 0.484 16.7 20.7 81
35-E 0.484 19.3 23.9 94
36-OD 0.450 18.9 25.2 99
36-ID 0.450 16.1 21.5 84
36-E 0.450 18.4 24.5 96
All three samples were tensile tested to failure.
Tensile Strain, %
Samples 34, 35, and 36 failed at 380-ksi (100% fiber
UTS), 324-ksi (77% fiber UTS), and 381-ksi (97% fiber
UTS), respectively. The results of these tests indicate
that there is no degradation in tensile modulus from five
180-degree folding and deployment cycles over a
15/32-inch radius. However, one tow (#35) had visual
fiber damage after five folds and experienced a 25%
reduction in tensile strength.
Tow Flexural Modulus Test Results
The composite tow flexural modulus tests were
performed using the 3-point bending fixture of the
(DMA) using a similar procedure as that used to test the
neat resin samples. The tow used for this test had a
fiber volume fraction of 59% and was cut from the end
of the isogrid test tube (#042) characterized by JMU.
The results in Figure 20 show that the flexural modulus
of the composite tow sample decreases 86% at 100oC,
which is a similar result to that obtained during the neat
TP283E resin flexural modulus testing. Table 6 shows
the strain ranges that were used to calculate the
modulus values at the different test temperatures.
0
0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2
Tensile Strain, %
Tensile Stress, ksi
Fold Outside
Fold Inside
Fold Edge
E = 17.8 msi (95% ROM)
Figure 18. Tensile Stress Vs. Strain: Sample 32; 3
folds

8
Flexural Stress-Strain Curve
IM9 (48K)/TP283E Tow No. 37
100
80
60
40
20
Tow extracted from tube prior to patch application
Analysis for round cross
section
Fiber Content = 59 vol.%
Table 6. Tow Flexural Modulus vs. Temperature
Temp. 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 - 0.5
23 12.6 0 - 0.5
50 8.8 0 - 0.1
75 4.1 0 - 0.05
100 1.8 0 - 0.05
Tow CTE Test Results
PMIC of Corvallis, Oregon tested several TP283E/IM9
(48K) composite tow samples to determine their CTE
values. Tow number 40 (Vf=61%), which was cut from
the end of boom #042 had a CTE of 0.3 ppm/oC (-100o
to 20oC) and 0.4 ppm/oC (-20o to 40oC). Another
composite tow sample (# 012-180+4, Vf=55%) was cut
from an isogrid boom and was tested in to understand
some of the tube level effects on CTE. This tow had a
CTE of = 0.7 ppm/oC (-100o to –20oC) and 1.0 ppm/oC
(-20o to 40oC). Although both of these tests values are
extremely low and are be acceptable for most
applications, they are slightly higher than the ROM
predictions of Figure 11. A number of root causes were
identified during this testing that directly affect the CTE
of the tows. These fundamental issues will be
optimized in the future in order to achieve lower CTE’s,
which will be closer to the ROM predictions.
Isogrid Boom Testing
Two booms were fabricated for this study. One boom
(#042) was loaded in compression, tension, tested for
creep, and then failed in compression at an elevated
temperature. The other boom was tested for CTE (#25).
A section of another boom (#12) was also tested for
CTE. The isogrid design chosen for this research was
the Inflatable Solar Array Experiment II (ISAE-II)
baseline design. This particular configuration was
chosen to allow for comparison between isogrid booms
fabricated by ILC in the past with both thermoplastic
and UV epoxy resins. The fiber volume fractions of the
booms were calculated based on the average value from
five tows cut from the ends of the booms after
manufacturing. The properties of the booms are shown
in Table 7. Like the composite tow samples, the isogrid
booms were fabricated from TP283E epoxy resin and
48K IM9 carbon fibers. The booms were manufactured
using a modified filament winding technique. The
actual 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 inches
Isogrid Base Length 1.374 inches
Boom Inner Diameter 7.05 inches
Number of Longitudinals 16
Number of Helicals 16
Tow Size (IM9) 48K
Average Tow Diameter 0.0534 inches
Resin TP283E
Mass of Boom #042 199.5 grams
Length of Boom #042 48.375 inches
Boom Test Length #042 46.75 inches
Average Fiber Volume Fraction #042 57.8%
A laser interferometer was used to characterize the CTE
of the isogrid tube (#25) fabricated for this study along
with a tube sample (#12)
extracted from a longer
isogrid boom. Tube #25
was first tested in the
pristine condition. The
approximate average CTE
(–40oC to +40°C) in the
pristine condition of tube
#25 is 1.15 ppm/°C. After
Z-folding tube #25 over a
15/32-inch radius one time
and deploying it using
inflation pressure, the boom
had an average CTE (-40oC
and +40°C) of
approximately 1.15 ppm/°C. Therefore it appears that
folding has no effect on the CTE of the boom, but
further testing is required to verify this initial result,
especially with respect to the orientation of the fold line
of the tube during the testing.
The results from the section of isogrid boom (#12) cut
from 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 the
maximum CTE was approximately 0.95 ppm/°C after
the fourth thermal cycle (Fig. 22). From these results, it
0
0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 0.008
Flexural Strain, in/in
27C -150C
-100C -50C
23C 50C
75C 100C
Figure 20. Tow Flexural Stress vs. Strain
Figure 21. Isogrid Tube
During CTE Testing

9
is apparent that thermal cycling stabilizes the CTE of
the isogrid boom. Like the tow CTE testing, the results
of these tests revealed a number of issues that will be
optimized in the future in order to further reduce the
CTE of the structure, thereby making it closer to the
ROM predictions in Figure 11.
CTE ≅ 0.82 ppm/°C
CTE ≅ 0.95 ppm/°C
Figure 22. Isogrid CTE Results; Tube #012
The test stand used to obtain the tensile and
compressive elastic modulus, as well as the preliminary
creep data is described in detail elsewhere4. The boom
is mounted horizontally on the test frame inside an 18-
inch diameter by 52-inch long radiant heater. Axial and
compressive loads are applied using a computer
controlled, servomotor driven, linear motion system.
Fiber optic strain gages were used because they were
better suited for mounting to the ribs of the isogrid
structure than typical bonded resistance type strain
gages. Four strain gages were mounted 90o apart at the
center of the boom. The use of fiber optic strain gages
and associated signal conditioning necessitated using
two different data acquisition systems, one for the strain
gages and one for the load cell and thermocouples.
Twelve thermocouples were used to measure
temperature. They were arranged in three groups of
four gages. The gage groups were located
approximately six inches from each end of the tube and
in the center. The four gages at each location were
located 90° apart. For elastic modulus testing and creep
testing the tube was loaded with approximately 50-lbf
(1400-psi). For elastic modulus testing a cross-head
rate of 0.05-in/min was used to apply the load. The load
was then held and the tube allowed to relax until the
strain gages stabilized. Tensile and compression tests
were performed between 25-64oC. When data
collection was completed at the test temperature, the
tube was unloaded and allowed to cool. Between
elevated temperature tests, a room temperature test was
performed to ensure that there were no structural
changes in the tube caused by the heating. The results
from the tensile and compression modulus testing are
shown in Figures 23 and 24.
Tensile Modulus Vs. Temperature IM9
(48K)/TP283E Isogrid Boom #042
25 30 35 40 45 50
Figure 23. Isogrid Tensile Modulus Results
30
25
20
15
10
5
0
Compression Modulus Vs. Temperature
IM9 (48K)/TP283E Isogrid Boom #042
30
25
20
15
10
5
The average tensile modulus of the boom from 28-47oC
is 24.5-Msi ± 0.3-Msi. Normalized to 60% fiber
volume fraction, the tensile modulus of the boom is
25.5-MSI. These compressive modulus test values
compare favorably (within 1%) with the predicted
ROM modulus as seen in Figure 10. The average
compressive modulus of the boom from 25-58oC is
21.8-Msi ± 1.5-Msi. Normalized to 60% fiber volume
fraction, 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 at
approximately 55-59oC. At these temperatures, for
both the compressive and tensile cases, the boom
exhibited viscoelastic behavior. As seen in Figure 25
for the compressive loading case at 64oC, strain
relaxation occurs where the load was held at 50-lbf.
This behavior occurs because of the large drop in
modulus of the resin above Tg. At temperatures above
55oC the modulus becomes temperature and time
dependent. These tests indicate that the viscoelastic
Temperature (OC)
Modulus (MSI)
Vf=57.8
Figure 24. Isogrid Compression Modulus Results
0
25 35 45 55
Temperature (OC)
Modulus (MSI)
Vf=57.8%

10
behavior of the boom must be considered during design
and deployment.
5
0
-5
-10
-15
-20
-25
Since understanding the viscoelastic behavior of the
boom is fundamental to understanding boom
performance, preliminary creep tests were performed.
Strain relaxation data were obtained at four
temperatures between 46oC and 65 °C. The time
constant for the relaxation at each temperature was
obtained from equation (3).


= −
ε ε τ
where ε(t) is the strain at any instant in time, ε∞ is the
strain at steady state, t is time, and τ is the time
constant. A plot of the time constant vs. 1/T is shown
in Figure 26.
1000
100
10
The trend line represents the behavior of the material
predicted by the Arrhenius equation (4).
RT (4)
Where T is the temperature in Kelvin, and R
ΔE
is a
constant related to the activation energy and has units
of K. The Arrhenius equation can be used to predict the
creep behavior of the boom outside of the temperature
range measured. For example, a typical operating
temperature for the boom is 10°C (283 K). The
predicted time constant at this temperature is 94,504
seconds or 26.25 hours. This is the time required for
63% of the strain relaxation to occur. 98% of the strain
relaxation will occur at 4 time constants or 105 hours.
The final boom test
performed at JMU was the
compression to failure at
elevated temperature test.
The expected maximum use
temperature of a boom of
this construction is 35oC.
With a safety factor of 5oC,
it was decided that the tube
should tested at
approximately 40oC. For
this test the tube was loaded
in compression at a strain
rate of 0.05 in/min. The tube
failed in rib buckling at 418.2-lbf, 11681.6-psi in
compression at 40.6oC (Figure 27). These results
compare favorably to the predictions obtained by using
ILC’s isogrid analytical model1,3. Using the inputs in
Table 9 and the ROM modulus at 57.8% Vf of 24.5-
Msi, the predicted compressive failure load is 443-lbf in
rib buckling, which is 5.6% above the actual load.
Summary and Conclusions
The results of the testing discussed herein clearly
illustrate the excellent agreement between rule of
mixtures mechanical predictions and ILC’s rigidizable
composite isogrid structure test values in the
temperature range of approximately –40oC to 55oC.
The isogrid boom exhibits viscoelastic behavior above
those temperatures. The high failure load of the
TP283E/IM9 (48K) isogrid boom of 418-lbf at 40oC
also shows that the booms are structurally sound up to
the expected use temperature of 35oC. ILC’s isogrid
analytical model predicts the failure load accurately at
temperatures of at least 40oC. Isogrid technology is also
scalable to tens and hundreds of feet in length and has
been demonstrated in continuous lengths up to 23 feet
(Fig. 28). The results of the composite tow tensile tests
indicate 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
0 50 100 150 200 250 300
Time (s)
Micro-Strain (avg. of 4 gages)
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)  

 

−
∞
t
(t) 1 e
Figure 27. Isogrid
Boom #042 During
Compressive Failure
E
oe
Δ
−
τ =τ
τ= 2E-19e15427/T
R2 = 0.9836
1
0.00295 0.003 0.00305 0.0031 0.00315
1/T (1/K)
τ (s)
Figure 26. Relationship between the Time Constant for
Strain Relaxation and Temperature

11
unaffected by up to five 180 degree folding and
deployment cycles over a 15/32-inch radius. Future
testing will be conducted to reduce this fold radius.
These tests also indicate that the
ultimate tensile strength of the tows
is within 25% of the ultimate tensile
strength of the IM9 fibers. This tight
folding can be achieved by the large
reduction in flexural modulus of the
tows at elevated temperatures (65%
reduced at 75oC, 80% reduced at
100oC, and 95% reduced at 125oC).
These folding tests indicate the high
packing efficiency that can be
obtained with ILC’s isogrid booms
with no reduction in tensile modulus.
The CTE tests indicate that ILC’s
TP283E/IM9 isogrid booms have a
near-zero CTE of approximately 1
ppm/oC. The initial results of these
tests also indicate that there are no
apparent folding effects on the CTE
of isogrid. Root causes at the
structural and tow level were
identified to reduce the CTE of the
booms in the future and bring them
closer to rule of mixtures thermal
predictions. The strain relaxation
testing on the isogrid booms also
indicated that the structure will
become stable to creep at 10oC at
approximately four days after
deployment.
Acknowledgments
Figure 28. 7”
diameter, 23
foot long,
Isogrid Boom
in test at
NASA-LaRC
The authors thank Mr. Jim McManus from Luna
Innovations, Inc. for the donation of the strain
displacement devices for the thermo-mechanical testing
of the isogrid boom at JMU. The authors also thank
JPL and DARPA for supporting the composite testing
at The Aerospace Corporation.
References
1. Cadogan, D.P., Lin, J.K, Sapna, G.H.,
Scarborough, S.E., “Space Inflatable Technology
Development for Solar Sails and Other Gossamer
Applications: GR/SMP Isogrid Boom Development
Final Report,” NASA Task Order 10442, ILC Dover,
Inc., October, 2001.
2. Darooka, D.K., S.E. Scarborough, and D.P
Cadogan, “An Evaluation of Inflatable Truss Frame For
Space Applications,” AIAA-2001-1614, 42nd
AIAA/ASME/ ASCE/AHS/ASC Structures, Structural
Dynamics, and Materials Conference and Exhibit
AIAA Gossamer Spacecraft Forum, April 16-19, 2001.
3. Lin, J.K., G.H. Sapna, Cadogan, D.P., S.E.
Scarborough, “Inflatable Rigidizable Isogrid Boom
Development,” AIAA-2002-1297, 43rd AIAA/ASME/
ASCE/AHS/ASC Structures, Structural Dynamics, and
Materials Conference and Exhibit AIAA Gossamer
Spacecraft Forum, April 22-25, 2002.
4. Pederson, L.M., Blandino, J.R., Woods-Vedler,
J.A., “Determination of the Modulus of Elasticity as a
Function of Temperature for an Isogrid Tube,” AIAA
2002-1334, Proceedings of the 43rd
AIAA/ASME/ASCE/AHS/ASC Structures, Structural
Dynamics and Materials Conference, Denver, CO, 22-
25 April, 2002.
5. Cadogan, D.P. and S.E. Scarborough “Rigidizable
Materials for use in Gossamer Space Inflatable
Structures,”42nd AIAA/ASME/ASCE/AHS/ASC
Structures, Structural Dynamics, and Materials
Conference & Exhibit AIAA Gossamer Spacecraft
Forum, Seattle, WA, April 16-19, 2001.
6. Mikulas, M.M., Jr., “Structural Efficiency of Long,
Lightly Loaded Truss and Isogrid Columns for Space
Applications,” NASA Technical Memorandum 78687,
July 1978.
7. Darooka, D.K., S. Scarborough, S. Malghan, D.
Cadogan, C. Knoll, “Inflatable Space Frame,” Final
Report, NASA Prime Contract Number: NAS1-99154,
July 2000.
8. Cadogan, D.P., S.E. Scarborough, J.K. Lin, G.H.
Sapna, “Shape Memory Polymer Composite
Development For Use in Gossamer Space Inflatable
Structures,” AIAA-2002-1372, 43rd AIAA/ASME/
ASCE/AHS/ASC Structures, Structural Dynamics, and
Materials Conference and Exhibit AIAA Gossamer
Spacecraft Forum, April 22-25, 2002.
9. Hexcel Magnamite IM9 Carbon Fiber Product
Data Sheet, March 2002.

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