Jeffrey J. DeMange

The University of Toledo, Toledo, Ohio

Robert J. Bott

Case Western Reserve University, Cleveland, Ohio

Patrick H. Dunlap

Glenn Research Center, Cleveland, Ohio

Investigation of the Mechanical Performance

of Compliant Thermal Barriers

NASA/TM—2011-217142

December 2011

AIAA–2011–2388
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Jeffrey J. DeMange

The University of Toledo, Toledo, Ohio

Robert J. Bott

Case Western Reserve University, Cleveland, Ohio

Patrick H. Dunlap

Glenn Research Center, Cleveland, Ohio

Investigation of the Mechanical Performance

of Compliant Thermal Barriers

NASA/TM—2011-217142

December 2011

AIAA–2011–2388

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

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San Francisco, California, April 11–14, 2011
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Level of Review: This material has been technically reviewed by technical management. 
Compliant thermal barriers play a pivotal role in the thermal protection systems of

advanced aerospace vehicles. Both the thermal properties and mechanical performance of

these barriers are critical in determining their successful implementation. Due to the custom nature of many thermal barriers, designers of advanced spacecraft have little guidance

as to the design, selection, and implementation of these elements. As part of an effort to

develop a more fundamental understanding of the interrelationship between thermal barrier design and performance, mechanical testing of thermal barriers was conducted. Two

different types of thermal barriers with several core insulation density levels ranging from

62 to 141 kg/m3 were investigated. Room-temperature compression tests were conducted

on samples to determine load performance and assess thermal barrier resiliency. Results

showed that the loading behavior of these thermal barriers was similar to other porous,

low-density, compliant materials, such as elastomeric foams. Additionally, the insulation

density level had a significant non-linear impact on the stiffness and peak loads of the

thermal barriers. In contrast, neither the thermal barrier type nor the level of insulation

density significantly influenced the room-temperature resiliency of the samples.

Nomenclature

A footprint area (m2)

f generic experimental measurement

F unit load (N/mm)

l thermal barrier length (mm)

m mass of thermal barrier insulation (kg)

P load (N)

V volume of thermal barrier core (m3)

w thermal barrier width (from contact footprint) (mm)

δn+1 permanent set, cycle n+1 (mm)

∆n maximum displacement, cycle n (mm)

ρrel relative density of the thermal barrier core insulation

ρ

∗ density of the thermal barrier core insulation (kg/m3

)

ρs density of the aluminosilicate from which the thermal barrier core insulation is made (kg/m3)

NASA/TM—2011-217142 1

Investigation of the Mechanical Performance of 

Compliant Thermal Barriers 

Jeffrey J. DeMange 

The University of Toledo 

Toledo, Ohio 43606 

Robert J. Bott 

Case Western Reserve University 

Cleveland, Ohio 44106 

Patrick H. Dunlap 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Abstract 
σ compressive stress (kP a)

As cross-sectional area of thermal barrier core (m2)

Rn resiliency, cycle n (%)

U

b

f

bias error in function or measurement

U

p

f

precision error in function or measurement

U

t

f

total uncertainty in function or measurement

U

t

ρ∗ total uncertainty in density

Ux,y,...,z individual uncertainties for the measurements making up a given function

I. Introduction

C

ompliant thermal barriers are an integral component of the thermal protection systems (TPS) for nearly

all aerospace vehicles. These barriers (Fig. 1), usually made from high-temperature ceramic-based fibrous

materials, are typically installed in the interface gaps between various components and substructures. Example locations include areas around access panels and doors, control surfaces, gaps between TPS tiles/panels,

perimeters of vents and instrumentation components, and within propulsion systems (Fig. 2). The primary

role of the thermal barriers is to help protect the underlying components and structures from excessive

heating which may be encountered during high-speed vehicle travel. In addition, the barriers must accommodate structural deflections resulting from thermal, aerodynamic, and/or reaction loads without damaging

surrounding materials.

Figure 1. Examples of compliant thermal barriers: blanket thermal barrier (left) and hybrid thermal barrier

(right).

(a) Shuttle external tank door (b) X-38 rudder (c) Shuttle reaction control system

Figure 2. Examples of locations on aerospace vehicles requiring thermal barriers.

NASA/TM—2011-217142 2
to custom-tailor their performance. Due to the multitude of configurations and often limited performance

data, vehicle designers have little guidance on which thermal barrier design to incorporate in a given interface.

While it is intuitively obvious the designer needs to assess thermal performance, one must also consider the

mechanical behavior of the barriers. With the exception of other application-specific studies conducted by

the authors, 1–5 no significant fundamental information on thermal barrier mechanical properties exists in

the open literature. An understanding of the load characteristics of the barrier is not only important during

a mission, but must also be accounted for in ground operations (e.g., assembly). Excessive loads during any

phase of the operations can result in vehicle damage.

In cooperation with the NASA Glenn Research Center, researchers have embarked on an effort to develop

a more fundamental understanding of the interrelationship between thermal barrier design and performance.

As part of this effort, an investigation was undertaken to evaluate the effect of insulation density on the

mechanical properties of thermal barriers. Room-temperature compression tests were performed on two

configurations of samples containing several different insulation density levels with load and resiliency results

reported herein.

II. Experimental Procedure

A. Test Specimen Description

1. Blanket Thermal Barriers

The blanket thermal barrier (BTB) test specimens were fabricated from two concentric outer sheaths of an

aluminoborosilicate woven fabric and a core filled with an aluminosilicate insulation blanket (Fig. 3) at four

different density levels: 62, 80, 123, 141 kg/m3.

Figure 3. Blanket thermal barrier.

In its as-received form, the aluminosilicate insulation was constructed as a laminated blanket with polycrystalline fibers (mean diameter of 3.0 - 3.5 μm) oriented randomly within the layers (Fig. 4). According

to the vendor, the as-received blanket had a nominal density of around 96 kg/m3. To achieve the desired

density levels in the thermal barrier specimens used for the current tests, the blanket was cut to appropriate

sizes and then inserted into the sheath layers. The blanket laminate layers were oriented in a slightly curved

vertical orientation, as shown in Fig. 3 (i.e., layers predominantly oriented parallel to loading direction). The

thermal barriers were also stitched through the thickness using aluminoborosilicate thread. This stitching

was implemented as a means to help retain the insulation should the barrier be damaged during an actual

flight.

After fabrication was completed, five samples of each density level were cut to length using a dry diamond

wheel and then heat-cleaned at 482±11 ◦C for 15 hrs to remove residual organic sizing. Two to three shorter

samples of each density level were also cut and heat-cleaned for density verification. All samples were weighed

before and after heat cleaning. The samples were dimensionally measured (width and height) before and

after compression testing using a non-contact laser system (Keyence LJ-G200) accurate to within 96 μm.

Sample lengths were measured using calipers accurate to 20 μm.

2. Hybrid Thermal Barriers

The hybrid thermal barrier (HTB) test specimens (Fig. 5) were designed to combine the insulating performance of the blanket thermal barrier with the high-temperature resiliency of a metallic spring tube thermal

Thermal barriers can be fabricated in a multitude of configurations and several parameters can be varied

NASA/TM—2011-217142 3
.

Figure 4. Aluminosilicate fibers in blanket insulation (Courtesy Saffil Ltd.).

barrier. The samples were fabricated by inserting a Rene 41-knitted spring tube (Fig. 6) filled with the aluminosilicate insulation into two outer sheaths of an aluminoborosilicate-woven fabric. The remaining interior

volume of the sheath layers was also filled with the insulation. Thermal barrier samples were fabricated using

this configuration at two insulation density levels, 76 and 127 kg/m3. Note that this was the overall average

density level for each HTB configuration, inclusive of both the insulation inside and outside of the spring

tube. For the insulation outside the spring tube, the layers were generally oriented in a slightly curved vertical configuration (Fig. 5). The insulation inside the spring tubes exhibited more variation between samples

in terms of layer orientation.

Figure 5. Hybrid thermal barrier. Figure 6. Spring tube (shown with insulation).

Five samples of each density level were cut to length using a dry diamond wheel and heat-cleaned at

482±11 ◦C for 15 hrs to remove residual organic sizing. Two to three shorter samples of each density level were

also cut and heat-cleaned for density verification. All samples were weighed before and after heat cleaning

and dimensionally measured (width and height) using a non-contact laser system (Keyence LJ-G200) before

and after compression testing. Sample lengths were measured using calipers.

B. Test Facility Description

An MTS Model 318 servohydraulic load frame with custom test fixturing was used to evaluate the load

performance of the thermal barrier samples (Fig. 7). The load frame incorporated a 2.2 kN (accurate to

within ±0.6 N) load cell to monitor loads during testing. The samples were compressed between two smooth

metallic platens. Displacement between the loading platens was measured using a Beta LaserMike Intelliscan

50 laser extensometer (accurate to within approximately ±6.4 μm).

C. Test Procedure

1. Density Estimation

The insulation density of each configuration was determined by sampling several short lengths for each design.

After heat-cleaning, weighing, and measuring the dimensions, each density sample was carefully dissected,

and all the insulation was removed and weighed. Based upon the laser measurements, a cross-sectional area

was estimated assuming an elliptical shape of the thermal barrier core cross section. The corresponding

elliptical shapes were then visually compared against the photographs to evaluate the relative accuracy of

NASA/TM—2011-217142 4
Figure 7. Thermal barrier compression test rig.

the shape assumption. As a confirmation, areas were also measured directly using a commercially available

photographic software package. However reference measurements were not always available to accurately

scale the areas. In addition, the ends of the thermal barriers often exhibited some local bulging and fraying

due to the cutting process and may not have accurately reflected the true cross section of the samples.

Additional errors may have been introduced by non-normality of the thermal barrier cross section relative

to the camera perspective. Based upon the insulation mass, core area estimate, and the measured length of

the sample, the insulation density was calculated according to Eq. (1).

ρ

∗ =

m

V

=

m

As·l

(1)

2. Compression Tests

The thermal barrier samples were compressed at room temperature (approximately 20 ◦C) between two flat

platens to a final gap which corresponded to approximately 30% of their average initial height (i.e., 70%

compression). During testing, the samples were loaded to approximately 0.9 N to achieve initial uniform

contact and then compressed at a strain rate of roughly 5.0 × 10−3to the final gap. After dwelling at this

gap for approximately 60 s, the samples were unloaded at the same rate to the initial contact point. This

load-unload cycle was repeated four times. Prior to the first cycle, a carbon transfer tape was inserted

between the top of the sample and the upper platen to generate a contact footprint for compressive stress

calculations. After the first cycle, the tape was removed and the setup was adjusted to account for the

thickness of the tape. Five samples were tested for each thermal barrier type and density level.

One of the primary responses from the compression tests was load behavior, calculated both in terms

of unit loads and stresses. The unit loads were calculated based upon the load and initial length of the

specimen (Eq. (2)). The compressive stresses were determined from the loads and contact footprint areas,

as per Eq. (3).

F =

P

l

(2)

NASA/TM—2011-217142 5
σ =

P

A

=

P

l·w

(3)

Thermal barrier resiliency, which is a measure of how well the barrier springs back to its original undeformed size, was determined after each load cycle using Eq. (4).

Rn =

∆n − δn+1

∆n

∗ 100 (4)

A schematic further illustrating the calculation of resiliency is shown in Fig. 8. As shown in this schematic, the

resiliency is the amount of compression recovered after unloading from a full cycle of compression. The “permanent set” in the sample on subsequent cycles is defined as the stroke at which the sample reaches the predefined preload value. While resiliency at high temperature is often of prime interest, ambient-temperature

resiliency can also be important in terms of ground operations and repeated assembly/disassembly cycles.

(a) Illustration of terms for determining resiliency (b) Determination of permanent set

Figure 8. Example plots illustrating calculation of resiliency.

3. Uncertainty Estimation

An uncertainty analysis is an important aspect of any experimental value. The total uncertainties for

each measure or calculation (f) were estimated by combining the precision (random) error and the bias

(systematic) error, according to Eq. (6).

f = f(x, y, . . ., z) (5)

U

t

f =

q

(U

p

f

)

2 + (U

b

f

)

2 (6)

For these analyses, the bias and precision error were assumed to be independent.

The precision error for each measure or calculation was determined based upon the standard deviation

of the tested samples weighted at a 90% confidence interval. The bias error was estimated by using Eq. (7).

U

b

f =

s

∂f

∂xUx

2

+



∂f

∂y Uy

2

+ . . . +



∂f

∂z Uz

2

(7)

Note that Ux,y,...,z are the individual uncertainties for the measurements making up a given function (typically

taken as the accuracy of the instrument or estimated based upon prior experience).

NASA/TM—2011-217142 6
III. Results and Discussion

A. Density Estimation

Density results from sampling of the thermal barriers are presented in Tables 1 and 2 for the blanket

thermal barriers and hybrid thermal barriers, respectively. While a range of density values was obtained, the

estimated uncertainties illustrate the potential for some overlap between levels. The uncertainty estimations

(U

t

ρ∗ ) are likely somewhat conservative, but also follow from a high sample variability in some cases combined

with a limited sample size. This was especially true for the blanket thermal barrier with a core density of

123 kg/m3 and the hybrid thermal barrier at a density level of 127 kg/m3. The uncertainties were also

fairly high in general due to the difficulty in obtaining accurate volumetric measurements. This is illustrated

in Fig. 9, which shows representative photographs of the ends of the thermal barriers with superimposed

ellipse data, as determined from the laser measurements. In some cases, the ellipse assumption appeared

to accurately recreate the thermal barrier cross-sectional shape (Fig. 9(a)), while in others, the assumption

appeared to overestimate the area (Fig. 9(b)), and thus underestimate the density. Consequently, the high

estimated uncertainties are warranted in most cases and indeed seem to be consistent with previous anecdotal

experience.

Table 1. Average estimated insulation density results for blanket thermal barriers.

ρ

∗ Ut

ρ∗ ρrel

kg

m3

kg

m3

62 16 0.016

80 26 0.021

123 62 0.031

141 34 0.036

Table 2. Average estimated insulation density results for hybrid thermal barriers.

ρ

∗ Ut

ρ∗ ρrel

kg

m3

kg

m3

76 13 0.019

127 54 0.032

(a) Acceptable fit (b) Imperfect fit

Figure 9. Examples of conformance of assumed ellipse shape for thermal barrier volume determination (Note:

dimensions shown are normalized dimensions).

The tables also include a relative density for the core insulation, which is defined according to Eq. (8)

ρrel =

ρ

∗

ρs

(8)

This measure is a fundamental property often used to help characterize the behavior of low-density materials,

NASA/TM—2011-217142 7
such as honeycombs, foams, fiber-based materials, etc. 6 The relative densities presented in the tables are

similar to the values typically observed for low-density polymeric foams used for cushioning, packaging, and

insulation. 6

B. Blanket Thermal Barrier Compression Results

Results from the compression tests for the blanket thermal barriers are presented in Fig. 10 through 12. Note

the error bars in each graph denote total uncertainties at 90% confidence intervals. Figures 10(a) and 10(b)

show plots of peak unit load and peak compression stress at 70% compression, respectively, as functions of

compression cycle for the four insulation density levels. These graphs can aid the designer in conducting

structural analyses and determining what insulation density level should be used for thermal barriers placed

against specific TPS materials, given material strengths.

(a) Peak unit load (b) Peak stress

Figure 10. Blanket thermal barrier peak loads vs. compression cycle.

As shown in both plots, the peak loads increased significantly with increasing core density. For example,

peak loads for the first compression cycle escalated by a factor of 6X when the core density increased from

62 to 141 kg/m3(i.e., 2.3X increase). This trend is more clearly illustrated in Fig. 11, which shows peak

loads as a function of core density for compression cycles 1 and 5. As illustrated in the graph, the loads

appeared to increase in a non-linear fashion, especially as the insulation density approached the highest level

(141 kg/m3). With increasing insulation density, core porosity decreased and therefore resulted in more

fiber-to-fiber contact (and subsequently higher loads) as the sample was compressed. This type of behavior

is also typically observed for low-density elastomeric foams. 6

Figures 10(a) and 10(b) also show that peak loads decreased with each subsequent compression cycle. This

was likely the result of some compaction or “set” in the thermal barriers, possibly due to fiber entanglement,

inter-fiber friction, or fiber breakage as the samples were repeatedly compressed. For the blanket thermal

barriers, this effect appeared to be more acute at higher core densities. For example, the peak load decreased

by approximately 39% after five load cycles for samples with an average core density of 141 kg/m3. In

contrast, peak loads diminished by approximately 25% for samples with an average insulation density of

62 kg/m3.

Plots showing load behavior as a function of percent compression during the first load cycle for the

different insulation density levels are presented in Figures 12(a) and 12(b). As shown in Fig. 12(b), both the

load and displacement were normalized to produce what essentially amounts to a stress-strain curve. Due to

the porous nature of the thermal barriers and the somewhat compliant behavior of the fiber-based blanket,

the compressive response of these barriers was similar to that exhibited by low-density foam materials, such

as polyethylene. 7–9 For open-celled polymeric foams, the stress-strain curves exhibit shapes similar to that

shown in Fig. 13. This response is composed of a small region of linear elasticity, followed by a plateau

NASA/TM—2011-217142 8
Figure 11. Blanket thermal barrier peak unit load vs. core insulation density.

regime, and finally a densification region. The deformation mechanics for these foams have been shown to

involve cell wall bending (linear-elastic region), followed by wall buckling (plateau region), and then finally

cell collapse and densification. 6

(a) Unit load (b) Stress

Figure 12. Blanket thermal barrier load vs. percent compression.

This same behavior is evident in Fig. 12(b) for the thermal barriers, though the influence of the insulation

blanket layers plays an additional role at a macro-level. In addition, there is not a clear differentiation between

the linear-elastic region and the plateau region in many cases. A closer examination of the deformation

mechanics of the thermal barrier (Fig. 14) illustrated that as the barrier was compressed, the blanket layers

further bent over. With continued deformation, the layers rotated to an orientation in which the throughthickness direction was nearly parallel to the loading direction. Similar to conventional foams, the porosity

in the samples increasingly collapsed and more fiber-to-fiber contact occurred as the specimens were further

compressed. Correspondingly, the stiffness increased with additional compression as the insulation material

NASA/TM—2011-217142 9
Figure 13. Compression performance of polyethylene as a function of relative density. 6

densified. As observed with other low-density porous materials, the stiffness increased more rapidly for

samples with higher core densities (and therefore lower initial porosity).

It should be noted that it was only possible to determine the contact area at maximum compression

(i.e., the carbon transfer tape only captured a final footprint). Therefore the compression stresses at lower

compression levels were slightly underestimated. Nevertheless, the impact of insulation density on load

behavior is clearly evident in Fig. 12(b). Additionally this data can guide maximum acceptable thermal

barrier compression levels given the mating material compressive strength.

Blanket thermal barrier resiliency results (after being compressed by 70% during each load cycle) are

presented in Fig. 15. As illustrated in this plot, the blanket thermal barriers exhibited approximately 20-

30% permanent set after the first load cycle at ambient temperatures. The samples continued to compact

with additional load cycling. This behavior would need to be considered if multiple assembly/disassembly

cycles which repeatedly compress the thermal barriers to high levels would occur during ground operations.

Although the plots appear to suggest improved resiliency with increased core density, this effect is not

statistically significant when the uncertainty in the data is considered. Therefore thermal barrier resiliency

likely cannot be improved by increasing insulation density.

C. Hybrid Thermal Barrier Compression Results

Compression results for the hybrid thermal barrier samples are presented in Figures 16 and 17. In general,

these samples exhibited load behavior similar to that of the blanket thermal barriers. A higher insulation

core density resulted in higher peak loads. As an example, the peak loads for the first compression cycle

were approximately 2.5X higher when the core density increased from 76 to 127 kg/m3. By comparison,

the peak loads for the blanket thermal barriers increased by a factor of 2.0X across the same approximate

density change.

As with the blanket thermal barriers, the peak loads decreased with increasing cycling (Fig. 16). It

was initially assumed that the hybrid thermal barrier, with its integral spring tube, would exhibit better

resiliency and maintain peak loads better than the blanket thermal barriers. Surprisingly, the average load

decrease after five compression cycles for HTB samples with an average density of 76 kg/m3 was 36%. As

shown in Table 3, this was a greater load drop than that observed with the blanket thermal barriers (27%

load decrease) at a similar density level.

As indicated previously, Figure 16 illustrates that at similar density levels, the hybrid thermal barriers

generally exhibited higher peak loads than the blanket thermal barriers. This was most evident for the

higher density samples (120 - 130 kg/m3). The additional stiffening of these types of thermal barriers was

NASA/TM—2011-217142 10
(a) Initial (undeformed) (b) 10% Compression

(c) 30% Compression (d) 70% Compression

Figure 14. Representative blanket thermal barrier at various compression levels.

Figure 15. Blanket thermal barrier resiliency vs. compression cycle.

NASA/TM—2011-217142 11
(a) Peak unit load (b) Peak stress

Figure 16. Blanket thermal barrier (BTB) and hybrid thermal barrier (HTB) peak loads vs. compression

cycle.

Table 3. Average decrease in load after five cycles for blanket thermal barriers and hybrid thermal barriers.

ρ

∗ Unit Load, N/mm Decrease

kg

m3 Cycle 1 Cycle 5 %

Blanket Thermal Barriers

62 2.6 2.0 23

80 4.9 3.6 27

123 10.0 6.5 35

141 15.4 9.4 39

Hybrid Thermal Barriers

76 7.0 4.5 36

127 17.6 10.7 39

likely a result of the spring tube stiffness 4 as well as the additional constraint imposed by the spring tube

on the insulation as the barriers were deformed. This would seem to coincide with the effect becoming more

acute at higher density levels.

Plots of unit load and compressive stress as functions of percent compression are presented in Figures 17(a)

and 17(b), respectively. These graphs compare the effect of similar density levels for the blanket thermal

barriers and hybrid thermal barriers. In general, the loading behavior of the hybrid thermal barriers was

similar to the blanket thermal barriers, indicating similar deformation mechanisms were occuring. The

augmented stiffness of the hybrid thermal barriers when compared to the blanket thermal barriers is further

illustrated in these plots.

Figure 18 compares the room-temperature resiliency between the two types of thermal barriers. As

illustrated, no significant improvement in resiliency was observed for the hybrid thermal barrier at room

temperature, even though they incorporated a resilient spring tube element. The reasons for this behavior

are unclear, but may be related to an insufficient restoring force in the spring tube when the surrounding

NASA/TM—2011-217142 12
(a) Unit load (b) Stress

Figure 17. Blanket thermal barrier (BTB) and hybrid thermal barrier (HTB) loads vs. percent compression.

volumes are filled with insulation. Therefore the resiliency in the thermal barrier appeared to be dominated at

room temperature by the performance of the insulation. The high-temperature resiliency would be expected

to demonstrate a performance advantage for the hybrid thermal barrier. Though no significant improvement

in resiliency was observed, recent testing of these barriers in high-temperature arc jet testing simulating

reentry has indicated some thermal advantage of the hybrid version. 10

Figure 18. Blanket thermal barrier (BTB) and hybrid thermal barrier (HTB) resiliency vs. compression cycle.

IV. Conclusions

Compliant thermal barriers are an integral part of the thermal protection systems of nearly all advanced

aerospace vehicles. However, designers of modern aerospace vehicles often have little guidance on the proper

design, selection, and implementation of thermal barriers. To that end, a study was initiated to begin to gain

a more fundamental understanding of thermal barrier configuration versus performance. Room-temperature

NASA/TM—2011-217142 13
compression tests up to approximately 70% compression were conducted on two types of thermal barriers:

(a) a blanket thermal barrier which contained only aluminosilicate insulation in the core and (b) a hybrid

thermal barrier which also included a metallic-knitted spring tube as part of the core, in addition to the

insulation. For each thermal barrier type, several different levels of core insulation density were evaluated,

ranging from 62 to 141 kg/m3.

Results showed that in general, the load behavior of the thermal barriers was very similar to that of

other complaint low-density materials, such as elastomeric foams. The load profiles were essentially linear

up to approximately 20-30% compression and then exhibited significant non-linear increases in loads as the

samples were compressed and the core porosity decreased. The level of core density had a significant nonlinear effect on barrier stiffness and peak loads. As an example, an increase of insulation density by 2.3X

resulted in a peak load increase of approximately 6X for the blanket thermal barrier type. In contrast, the

thermal barriers exhibited little sensitivity to type or core insulation density on barrier resiliency at ambient

temperature. These results provide a basis for proper thermal barrier design, selection, and implementation

and offer the potential for future modeling efforts.

References

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AIAA-2000-3508, AIAA, Huntsville, AL, 16-19 July 2000.

2 Dunlap, P. H., Steinetz, B. M., Curry, D. M., Newquist, C. W., and Verzemnieks, J., “Further investigations of control surface seals for the X-38 re-entry vehicle,” Proceedings of the 37th

AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit, No. AIAA-2001-3628, AIAA, Salt

Lake City, UT, 8-11 July 2001.

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Joint Propulsion Conference & Exhibit, No. AIAA-2002-3941, AIAA, Indianapolis, IN, 7-10 July 2002.

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material subject to dynamic crash loading,” International Journal of Impact Engineering, Vol. 21, No. 5,

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for the Orion Back Shell Interface Thermal Protection System Arc Jet Tests,” Tech. Rep. OTPS-0023, 7

February 2011.

NASA/TM—2011-217142 14
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01-12-2011 

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4. TITLE AND SUBTITLE

Investigation of the Mechanical Performance of Compliant Thermal Barriers 

5a. CONTRACT NUMBER

NNC08CA35C 

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6. AUTHOR(S)

DeMange, Jeffrey, J.; Bott, Robert, J.; Dunlap, Patrick, H. 

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WBS 644423.06.31.03.08.03 

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John H. Glenn Research Center at Lewis Field 

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13. SUPPLEMENTARY NOTES

14. ABSTRACT

Compliant thermal barriers play a pivotal role in the thermal protection systems of advanced aerospace vehicles. Both the thermal properties 

and mechanical performance of these barriers are critical in determining their successful implementation. Due to the cus- tom nature of 

many thermal barriers, designers of advanced spacecraft have little guidance as to the design, selection, and implementation of these 

elements. As part of an effort to develop a more fundamental understanding of the interrelationship between thermal bar- rier design and 

performance, mechanical testing of thermal barriers was conducted. Two different types of thermal barriers with several core insulation 

density levels ranging from 62 to 141 kg/m3 were investigated. Room-temperature compression tests were conducted on samples to 

determine load performance and assess thermal barrier resiliency. Results showed that the loading behavior of these thermal barriers was 

similar to other porous, low-density, compliant materials, such as elastomeric foams. Additionally, the insulation density level had a 

significant non-linear impact on the stiffness and peak loads of the thermal barriers. In contrast, neither the thermal barrier type nor the level 

of insulation density significantly influenced the room-temperature resiliency of the samples. 

15. SUBJECT TERMS

Seals; Thermal insulation; Thermal barriers; Thermal protection; Load test; Compression loads; Mechanical properties 

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