AIAA_SciTech_2014

Analysis of Reusable Integrated Thermal Protection

Panel Elements with Various Insulating Core Options

Fang Jiang∗, Wenbin Yu†, Zheng Ye‡, Ronald Kerans §¶, Ming Y. Chen‖

A finite element analysis (FEA) tool was developed for evaluating the effects of basicdesign features and materials choices on a basic panel for an Integrated Thermal ProtectionSystem (ITPS). A set of practical optimization problems regarding different insulating coreoptions was solved by utilizing the commercial FEA software ANSYS. The core optionsrepresent five different layouts for the insulation layer: bonded/unbonded foam with bladestiffeners, bonded/unbonded foam with hat-section stiffeners, and simple bonded foam. Aconventional design with parasitic insulation tiles was also analyzed as a reference. Figuresof Merit (FoMs) identifying the combination of load-bearing capability and the mass of theITPS were defined. Using these FoMs, the optimization objective functions were createdso as to consider both the insulation performance and structural strength and stiffness.Some examples of FoMs of each design candidate were optimized and compared with eachother to identify the best structural layout. Finally, the effects of materials properties onthe best core options were investigated.

I. Introduction

Thermal Protection Systems (TPS) are the heat shields attached to the surfaces of high speed air vehiclesto limit the temperatures of underlying structure. In general, TPS approaches include both ablative andreusable systems, depending on requirements, such as Apollo Avcoat ablator and NASA LI-2200 Shuttle tiles,respectively.1,2 TPS concepts are also divided into categories of passive, semi-passive, and active,3 in whichthe passive ones are regarded as the most weight efficient and generally the safest,4 and further classifiedinto load-carrying and nonload-carrying TPS.5 Because incident heating rates vary across a vehicle surface,different types of TPS are generally used on the same vehicle.4

From the perspective of reducing the expense of orbital transportation, there is interest in hypersonicvehicles with reusable TPS including Reusable Launch Vehicles (RLVs),6,7 miliary spaceplane,8 spaceplanesfor tourism,9 space trucks,10 suborbital package delivery vehicles,11 and hypersonic air breathing vehicles.12

Material type is a concern. Metallics are robust and waterproof, but heavy and of either limited temper-ature capability or with poor environmental resistance.13 The nose cap and wing leading-edge of a spacecraftoften reach the highest temperature. For these parts, some C-C and SiC/SiC composites can sustain hightemperature and have various applications.14 To keep the inner temperature of the Space Shuttle Orbiter lessthan 450 Kelvin, reusable surface insulation tiles were used primarily on the windward surface.15 Developmentof improved ceramic TPS was an active topic at the NASA Ames Research Center for many years.16 TheAlumina Enhanced Thermal Barrier (AETB) is a typical type of ceramic tile,17 which is usually used withToughened Uni-piece Fibrous Insulation (TUFI) coating and Reaction Cured Glass (RCG) coating. Thelatter versions are believed to be significantly stronger and more resistant to rain erosion than the originalShuttle tiles,16 and to have improved dimensional stability at high temperatures — to 2600 Fahrenheit andabove.13 However, these are relatively poor structural materials.18

∗Graduate Research Assistant, Department of Mechanical and Aerospace Engineering, Utah State University, Logan, Utah84322-4130.†Associate Professor, School of Aeronautics and Astronautics, Purdue University, West Lafayette, Indiana 47907-2045.

Associate Fellow, AIAA; Fellow, ASME; Member AHS.‡Reliability Engineer, Baker Hughes, Claremore, Oklahoma 74017. Student Member of AIAA.§Air Force Research Laboratory, AFRL/RXCC, Wright-Patterson AFB, Ohio 45433-7750¶UES, Inc., Beavercreek, OH 45433‖Air Force Research Laboratory, AFRL/RXCC, Wright-Patterson AFB, Ohio45433-7750

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American Institute of Aeronautics and Astronautics

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55th AIAA/ASMe/ASCE/AHS/SC Structures, Structural Dynamics, and Materials Conference

13-17 January 2014, National Harbor, Maryland

AIAA 2014-0351

Copyright © 2014 by Fang Jiang. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.

AIAA SciTech

A possible improvement might be an Integrated Thermal Protection System (ITPS) — a reusable load-carrying passive insulation system. The entire ITPS would have some structural capability and wouldcarry load so the structure underneath it can be lighter. The materials considered for the external surfaceof the ITPS are often Ceramic Matrix Composites (CMC) due to their excellent thermal and structuralperformance.19 Ceramic foams are used as insulating materials and can be combined with CMC stiffeners ascandidates for the insulation core. For the internal layer, which is used as a part of the structure, polymermatrix composites (PMC) such as the T650-35 fiber reinforced PMR-15 polyimide resin are regarded ascandidate aeroshell structural material for their good mechanical properties and light weight.20

Sizing of panels can be conducted via an optimization18,21 that includes both thermal analysis andstructural analysis. For load-carrying passive ITPS, the change of insulation layer thickness will influencestructural weight, strength, and temperature simultaneously, which likely results in groups of local optimum.21

So the parameter selection for design variables for sizing optimization is an issue.22 FEA software is oftenutilized for analysis of the complex structures such as honeycombs and corrugated stiffeners.23,24 In somecases, the thermal analysis model can be simplified from 3D to 2D and/or 1D after homogenization of theinsulation layer,18,25–27 which is often followed by a static structural analysis only under the mechanicalloading conditions separately.

In this paper, a basic panel for an ITPS consisting of CMC, insulation core and PMC was studied. Anactual design of a TPS system is strongly dependent upon the nature of the vehicle, its mission, and thedetails of the vehicle itself, and is well outside the scope of this work. The goal here was to build theframework for, and examine preliminary results of, a study to compare the basic characteristics of severalpossible panel options. To reveal the potential of the panel designs, optimizations of the different designcandidates were conducted by using the ANSYS Parametric Design Language (APDL). Then the resultswere compared to conclude a preferred design strategy considering both thermal and structural conditions.Lastly, a preliminary analysis of material properties on the FoMs of the panel options were studied.

II. Core Options of ITPS

The panel studied in this paper is a sandwich-like structure consisting of three layers; an outer layer ofCMC, an inner layer of PMC, and an insulation core. The area of the panel considered is 0.762 × 0.762 m2

(30 × 30 in2). The core is mainly responsible for thermal insulation, but in an ITPS system, it must alsocarry structural loads. In this study, there are 5 core options:

(1) Option 1: Blade stiffeners with unbonded foam.In this case, blade-shaped stiffeners of CMC carry primary loads across the core. The foam is divided

into separate cells, as shown in Figure 1(a). The arrangement of the blade stiffeners is shown in Figure1(b). The stiffeners are assembled so that those running in the x direction are continuous and those in thez direction are in 3 pieces each. The foams are not bonded to CMC, PMC or stiffeners.

(a) ITPS with core option 1 (b) Arrangement of blade stiffeners

Figure 1. Foam with blade stiffeners as insulation core.

(2) Option 2: Hat-section stiffeners with unbonded foams.In this case, hat-section stiffeners made of CMC carry loads across the core. The foam is in at least 15

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separate pieces, as shown in Figure 2 (a). The arrangement of the hat section stiffeners is shown in Figure 2(b). The assembly of hat-sections consists of long ones in the x direction separating segmented ones in thez direction. The foams are not bonded to CMC, PMC or stiffeners.

(a) ITPS with core option 2 (b) Arrangement of hat-section stiffeners

Figure 2. Foam with hat-section stiffeners as insulation core.

(3) Option 3: Blade stiffeners with bonded foam This option is the same as the option 1 except that theinsulating foam is bonded onto all walls of the chamber.

(4) Option 4: Hat-section stiffeners with bonded foamThis option is the same as the option 2 except that the insulating foam is bonded onto all walls of the

chamber.(5) Option 5: Bonded foamIn this option, the insulation layer composes of foam only and it is bonded directly to the CMC and

PMC layer to form the sandwich structure, as shown in Figure 3.

Figure 3. Bonded foam as insulation core.

S200 CMC manufactured by COI Ceramics, Inc.28 is used as the external face sheet of the ITPS. Thereinforcing fiber is Ceramic Grade Nicalon (CG-Nicalon, NL-201), and the SiC matrix is made by using a

Polymer Infiltration Pyrolysis (PIP) process. The density is 2000 kg/m3

at 293 Kelvin and 1900 kg/m3

at 3000 Kelvin. The Young’s modulus is 96 GPa at 293 Kelvin and 90 GPa at 1900 Kelvin, and thePoisson’s ratio is 0.27, not changing with respect to temperature. The temperature dependent specific heatcapacity, thermal conductivity, and coefficient of thermal expansion (CTE) are shown in Figure 4. Theproportional limit and the onset of significant matrix cracking, and hence of environmental degradation is inthe neighborhood of 0.1% strain for most SiC based CMCs. Consequently, that is taken to be the allowablestrain in this work.

SiC foam produced by Ultramet29 with density of 320 kg/m3

is used in the insulation core. The Young’smodulus of this foam is 2.873 GPa, and the Poisson’s ratio is 0.22. In addition, the specific heat capacity is1422.56 kJ/kg-K. The other temperature dependent material properties are shown in Figure 5.

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(a) Specific heat capacity (b) Thermal conductivity (c) CTE

Figure 4. Material properties of CMC.

(a) Thermal conductivity (b) CTE

Figure 5. Material properties of Foam.

The PMC is the 4-ply-fabric T650 laminate20 of which the Young’s modulus is 77 GPa at 293.15 Kelvinand 81 GPa at 616.15 Kelvin, and the Poisson’s ratio is 0.08. The density is 1900 kg/m

3, and the specific

heat capacity is 1200 kJ/kg-K. The other temperature dependent material properties of the PMC are shownin Figure 6.

(a) Thermal conductivity (b) CTE in plane (c) CTE through thickness

Figure 6. Material properties of PMC.

AETB tiles with TUFI coating protecting a PMC structure was modeled and analyzed as a referenceconventional system. The AETB ceramic tile with TUFI coating was developed at the NASA Ames ResearchCenter as an improvement to the LI-900 tile. The system is composed of an 8×8 in2 insulation tiles mountedon a felt Strain Isolation Pad (SIP) by using Room Temperature Vulcanizing (RTV) adhesive, as shown inFigure 7. This title is parasitic as the AETB carries no mechanical loads, which means only the PMClayer contributes the mechanical properties for the system as whole. FoMs of this parasitic case are used asreferences to assess the FoMs of ITPS with the five core options.

For efficient optimization, the configurations in Figure 1a, 2a and 3 are considered as the geometricalshapes of the FEA models in ANSYS. In order to eliminate the edge effect of the panel, the elements and

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Figure 7. Structure of ITPS with parasitic insulation tiles.

nodes in the central area of these configurations are selected as the sample structures to obtain the maximalvon Mises strains (see Conclusions regarding the choice of strain criteria). These sample structures of coreoption 3, 4, 5 are shown in Figure 8, 9, and 10 respectively.

Note for Option 1 and Option 2, the foam material is unbonded and thus will not contribute to loadbearing. For simplicity, in the thermomechanical analysis, we assume them to be bonded but with very smallYoung’s modulus. No gaps are considered in the thermal analysis.

(a) CMC (b) Foam (c) PMC

Figure 8. Finite element models of the sample structures of core option 3.





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III. Figures of Merit (FoMs)

A notional temperature cycle with a maximum value of 1324.37 Kelvin (1050 C, 1924 F) is added on theexternal surface of CMC for a duration is of 2400 seconds, as shown in Figure 11.

Figure 11. Temperature cycle applied to the external surface of CMC.

The mechanical loads, such as the bending and tensile boundary conditions, are also applied to the panelsresulting in stresses and strains that add to the stresses and strains due to the temperature distributiongenerated by the thermal loading. As the mechanical loads can be applied on either CMC and PMC layers,we consider two types of loading: (1) Inner-sheet loading, the tensile and bending loads are applied on PMClayer only; (2) Fully loading, the tensile and bending loads are applied on both CMC and PMC layers.

To evaluate the performance of the panels with five core options and the parasitic TPS, we define threedifferent FoMs as follows:

A. FoM of bending stiffness

The bending stiffness of a panel, S, is defined as

S =Mmaxke

(1)

where ke is the effective bending curvature of the loaded panel, Mmax is the max bending moment the panelcan sustain before any point in the sample structure reaches the failure strain. For this study, we set vonMises strain equal to 0.1% as the failure strain.

The effective curvature of the bent panel can be expressed by

ke =d2v

dx2(2)

Figure 12 is a schematic plot of the bending of a panel with the bending moment positive with respectto the z axis. Before deformation, the configuration of the structure is represented by black solid outlines,and after deformation it is plotted with red dashed outlines.

Assuming linear behavior, the displacement in the x direction, u, can be calculated by

u = −ycΦz (3)

where yc is the distance from the point having zero displacement. The strain in x direction can be calculatedby

εx =d(−ycΦz)

dx= −yc

d2v

dx2= −ycke (4)

which implies that the effective curvature is the changing ratio of the εx verses yc. In FEA, we compute theeffective curvature as:

ke = −∆εx∆y

= − εCMCx − εPMC

x

yCMC − yPMC(5)

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Figure 12. Geometry of the deformed differential panel segment

where εCMCx is the averaged nodal εx on the external CMC surface, εPMC

x is the averaged nodal εx onthe internal PMC surface, yCMC is the y-coordinate of the external CMC surface, and the yPMC is they-coordinate of the internal PMC surface. Finally, the bending stiffness of the panel can be calculated byusing Eq. (1) and Eq. (5) together and the FoM of bending stiffness is defined as

FoMS =S

ρa(6)

where the ρa is effective areal density of the panel which is equal to the total weight of the panel divided byits area.

B. FoM of bending strength

The FoM of bending strength can be easily calculated by

FoMB =Mmaxρa

(7)

C. FoM of tensile strength

The FoM of tensile strength is defined as the max tensile force T applied to the ends of the panel dividedby the effective areal density, which is expressed as

FoMT =Tmaxρa

(8)

IV. Optimization Algorithm

If the entire structure of a panel, including the insulation core and the CMC, has some load-bearing capa-bility, the structure underneath it (in this simple case, the PMC layer) can be lighter, thereby compensatingfor the increased weight of the core and CMC.

For optimization, some definitions have to be introduced. Firstly, a statement variable (SV) is definedas the constrained conditions which the optimized results must satisfy. During comparison of the five ITPSoptions, the statement variable is the maximum strain in the CMC and PMC layers, which is limited to beno more than 0.1%. The reason we do not limit the strain of the foam is that we shall have insight to thefoam’s material properties after we figure out the best ITPS core option. Another statement variable is thetemperature on the top of the PMC layer which is limited to be not higher than 560.928 Kelvin, the glasstransition temperature of the PMC.

Secondly, a design variable (DV) is defined as the parameter which influences the optimized objective.If we want to optimize the FoMs defined by Eq. (7) and Eq. (8), the loads applied on the panel shouldalso be considered as design variables under the ANSYS conventions. When we are optimizing the FoMs ofthe five ITPS core options, in order to focus on determining the best structural scheme, all of the material

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properties are specifically fixed, and the thickness of the core is considered as a design variable. After thebest core option was selected according to the FoMs, the thickness of the CMC and the material propertiesof both CMC and foam were taken into consideration. The parameters used in the optimization are shownin Table 1. Note the max mechanical load can be chosen from two values to avoid excessive time spent inoptimization as it is expected that option 5 without stiffeners will not be able to sustain as large a load asthe stiffened or bonded panels.

Table 1. Parameters used in optimization

Parameters Meaning Type Maximum Minimum

εPMC Max von Mises strain in PMC layer SV 0.001 0

εCMC Max von Mises strain in CMC layer SV 0.001 0

Tmax(K) Max temperature of PMC top surface SV 560.928 293.15

Mmax(kNm) Max bending moment DV 5.0, 55.0 0

Pmax(kN) Max tensile force DV 350.0, 500.0 0

tcore(m) Thickness of insulation core DV 0.2 0.125

Thirdly, the objective function (OBJ) is defined as to be minimized in ANSYS and only one objectivefunction can be set in one optimization. In this study, we define the objective functions of bending andtensile loading cases using Eq. (9) and Eq. (10) respectively.

OBJB =MmaxA

mmin− FoMB (9)

OBJT =TmaxA

mmin− FoMT (10)

where the A is the area of the panel structure and mmin is the minimal weight of the panel.We use the ANSYS optimization module to maximize the FoMs. First we need to carry out the transient

heat conduction analysis to obtain the time history of the temperature distribution within the panel andcheck whether the temperature of the PMC top surface satisfies the max PMC temperature constraint. Ifyes, we will continue to conduct a one-way coupled thermomechanical analysis under both the temperaturedistribution and applied mechanical load. A check is then performed to see if the max von Mises strain issmaller than the failure strain, if yes, the FoMs are computed. The optimization procedure needs to loop inthe heat conduction analysis and the one-way coupled thermomechanical analysis to optimize the dimensionof the panel and meanwhile find the max applied loads corresponding to max FoMs.

The optimization was conducted in steps with different optimizing methods, as shown in Figure 13. Thefirst step was to perform a random search of the design space of the variables; 50 random searches were usedin this work. This step was terminated after 25 feasible solutions were found, otherwise the total 50 analyseswere conducted. The feasible design sets were kept and others were removed before the second step, the firstorder iteration. This step is accurate but time-consuming because there are 4 to 6 analyses in each iterationin order to find the best trend of the design variables. If the result of the first iteration is not satisfactory,the code keeps the best sets and uses them to run the first order iteration again.

Figure 13. Optimization steps.

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The curve of optimized feasible FoM of inner-sheet tensile load for core option 5 is shown in Figure 14.By applying the optimization steps in Figure 13, the APDL code first scans the whole design variable spacewith a random search method, which results in feasible design sets quickly. After 34 loops of random search,the unfeasible solutions are deleted, and the first order iteration method starts with the best feasible setsand provides an accurate and fine convergence to the final FoM value.

Figure 14. FoMs of tensile strength of inner-sheet-loaded ITPS with core option 5.

Another important thing is the strain-to-failure. In any core option, one of the things of interest is varyingthe material properties of the foam to see if useful properties are in the realm of feasibility. For example, itis useful to know the maximum strain value in the foam when the CMC is at max allowable strain. Afterobtaining the best solution from the five core options and determining the optimized dimensions, the straindistribution of the foam with different material properties can be evaluated.

V. FoMs of Thermal Protection Systems

The FoMs can help us determine the best scheme from the five options. Please notice here the FoMsin this section are only comparable among the five options without the parasitic case, because we have notdetermined the competitive material properties of the CMC and foam.

Firstly, the inner-sheet loading (ISL) tests are conducted. Secondly, because the CMC in ITPS is also acompetitive load-bearing materials, the fully loading (FL) tests are analyzed too. And actually we expectthe fully loaded panel could obtain better FoMs as an advantage to the old parasitic baseline case. Table 2shows the optimized design variables and FoMs of core option 5.

Table 2. Optimization data of ITPS with core option 5

Optimized variable Unit Downwards bending Upwards bending Tension

FL / ISL FL / ISL FL / ISL

Tmax K 377.69 / 377.66 545.59 / 548.91 525.70 / 550.32

tcore m 0.19997 / 0.19999 0.13882 / 0.13806 0.14355 / 0.13774

m kg 45.926 / 45.926 34.561 / 34.420 35.440 / 34.361

Mmax kNm 34.698 / 35.098 24.508 / 24.442 NA

ke 10−3·m−1 5.631 / 5.675 8.880 / 8.917 NA

S kNm2 6161.897 / 6184.525 2759.820 / 2740.962 NA

FoMS kNm4/kg 231.087 / 231.919 137.527 / 137.145 NA

FoMB Nm3/kg 1301.28 / 1316.17 1221.29/ 1222.98 NA

Pmax kN NA NA 558.64 / 391.46

FoMT kNm2/kg NA NA 28.460 / 19.620

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From Table 2, we can see that for insulation core option 5, the downwards bending potential is betterthan the upwards bending potential with simultaneous thermal loadings. In addition, the bending behaviorsof the inner-sheet loaded panel and the fully loaded panel do not differ from each other. However, the fullyloaded panel has better tensile performance than the inner-sheet loaded panel.

In the parasitic case, the AETB tiles and their TUFI coatings contribute to the total weight of thissystem. Therefore, in the optimization the thickness of the ATEB layer is set to be a DV and the resultsare shown in Table 3.

Table 3. Optimized data of parasitic case

Optimized variable Unit Downwards bending Upwards bending Tension

m kg 11.73 11.12 11.16

Mmax kNm 428.848 428.685 NA

FoMS kNm4/kg 146.479 146.479 NA

FoMB Nm3/kg 62.978 66.419 NA

Pmax kN NA NA 317.88

FoMT kNm2/kg NA NA 49.051

The optimizations of panels with stiffened core options do not have feasible solutions. Max thermalstrains in option 1 are shown in Figure 15. The contour plots of corresponding strain distribution in thesample structures are shown in Figure 16 for CMC thickness of 0.002 m. Thickening the CMC face sheetand blade stiffeners does not influence the max thermal strains in CMC but raises the max thermal strainsin PMC layer. The max strain exceeding the 0.1% failure strain mainly occurs in the blade stiffeners, and itis the reason that there is no feasible solution of core option 1 in optimization.

(a) Maximal thermal strains in CMC (b) Maximal thermal strains in PMC

Figure 15. Time dependent maximal thermal strains in the sample structures with core option 1.

The maximum thermal strains in panels with core option 2 are shown in Figure 17. The correspondingstrain distribution in the sample structures with CMC thickness of 0.002 m are shown in Figure 18. Similarly,thickening the CMC face sheet and hat-section stiffeners did not influence the max thermal strains in theCMC, but raised the max thermal strains in the PMC layer. The strains that exceeded the 0.1% failurestrain mainly occurred in the hat-section stiffeners, and are the reason that there was no feasible solution.of core option 2 in optimization.

The maximum thermal strains in the panel with core option 3 are shown in 19. The corresponding contourplots of von Mises strains in the sample structures are shown in Figure 20 for CMC thickness of 0.002 m.Because the PMC is always safe in thermomechanical analysis under pure thermal loads, the contours of thePMC were not included. Thickening the CMC face sheet and blade stiffeners had slight effect on the thermalstrains in the CMCs, especially when the thickness is 0.002 m. Increasing the CMC thickness decreased themax thermal strains in the foam and raised the max thermal strains in the PMC. The strains that exceededthe limit of 0.1% mainly occurred in the blade stiffeners, and were the reason that there is no feasible solutionfor core option 3 in optimization.

The maximum thermal strains in panels with core option 4 are shown in Figure 21. The correspondingcontour plots of von Mises strains in the sample structures with CMC thickness of 0.002 m are shown in

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(a) Von Mises thermal strains in CMC (b) Von Mises thermal strains in PMC

Figure 16. Contour plots of thermal strains at 1200 second in the sample structures with core option 1.

(a) Maximal thermal strains in CMC (b) Maximal thermal strains in PMC


(a) Von Mises thermal strains in CMC (b) Von Mises thermal strains in PMC


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(a) Maximal thermal strains in CMC (b) Maximal thermal strains in PMC (c) Maximal thermal strains in PMC


(a) Von Mises thermal strains in CMC at 2400 second (b) Von Mises thermal strains in foam at 1200 second

Figure 20. Contour plots of thermal strains in the sample structures with core option 3.

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Figure 22. The strains that exceeded the limit of 0.1% mainly occurred in the hat-section stiffeners, andwere the reason there was no feasible solution for core option 4 in optimization.

It is clear that for cores with stiffeners, the stiffeners become a limiting part of the ITPS. This might beobviated in several ways. The allowable strain to avoid microcracking used here is well below the ultimatestrain capability of the CMC. If it were felt allowable to have modest microcracking in a reasonably smallregion of the stiffeners, then these designs would be acceptable.

(a) Maximal thermal strains in CMC (b) Maximal thermal strains in PMC (c) Maximal thermal strains in PMC


(a) Von Mises thermal strains in CMC (b) Von Mises thermal strains in foam


VI. Effects of Material Properties

It has been observed that the material requirements for thermal insulation and load-bearing capabilitysometimes conflict..18 To better understand this conflict, and perhaps identify paths to resolution, one cansimply vary the material properties, for example the Young’s moduli of the CMC and the foam. Here theinner-sheet bending of option 5 is used as an example to illustrate this feature of the developed APDLframework. The time history of max strain is plotted in Figure 23. Contour plots of von Mises straindistributions in the sample CMC and foam at the time points of overall maximum are shown in Figure 24aand 24b, respectively.

The effect of changing the elastic modulus of the CMC, with the same core thickness and loading, isshown in Figure 25. With increasing Young’s modulus in the CMC, the maximum strain in the CMC wasreduced while the strains in foam and PMC were not significantly affected.

In contract, Figure 26 shows that enhancing the Young’s modulus of the foam did not improve thepotential of the bonded foam panel for a specific design layout.

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Figure 23. Maximal strain during thermomechanical analysis of core option 5.

(a) Sample CMC at 1320 sec (b) Sample foam at 2400 sec

Figure 24. Element average von Mises strains distribution of core option 5.

(a) Maximal strains in sample CMC (b) Maximal strains in sample foam (c) Maximal strains in sample PMC

Figure 25. Max strains of core option 5 with different CMC Young’s Modulus.

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(a) Maximal strains in sample CMC (b) Maximal strains in sample foam (c) Maximal strains in sample PMC

Figure 26. Max strains of core option 5 with different foam Young’s Modulus.

VII. Conclusion

A framework for the design and analysis of sandwich panels that could form the basis for an IntegratedThermal Protection System (ITPS) has been developed using the ANSYS parameteric design language(APDL). Figures of Merit (FoMs) were defined to evaluate the thermomechanical performance of panelswith various options for the layout of the insulation core. All of the sandwich panels were found to fallshort of meeting the basic requirements. By optimizing the FoMs, the core option of simple bonded foamappeared to be the best design strategy, as it obtained the best FoMs among the core options. However thatresult depended upon the relaxation of the ultimate strain requirement of the foams. Identification of thenecessary requirements of a foam to satisfy ITPS requirements can easily be pursued using this approach.Determination of those requirements and comparison to what seems technologically feasible is a likely futuregoal. The parasitic case containing AETB tiles with representative configuration of 0.762 × 0.762 m2 areawas also analyzed. By evaluating the FoMs, the ITPS (with relaxed foam requirements) has better bendingpotentials but worse tensile performance than the parasitic case. Evaluation of varying the elastic modulus ofthe CMC as an approach to enhancing the FoMs of the ITPS panels was performed. Changing the thicknessof CMC was found to have little effect on the ITPS performance. Evaluating other parameters in the sameway should be a good use of this approach in the future. While a von Mises criterion provides certaininsights, on reflection, the authors now feel that it is not the best choice for the materials considered. Aclear topic for further work is to revisit the analysis using principal strain criteria.

Acknowledgements

This work is supported by the Air Force Research Laboratory Rapid Development and Insertion ofHypersonic Materials program. The views and conclusions contained herein are those of the authors andshould not be interpreted as necessarily representing the official policies or endorsement, either expressed orimplied, of the funding agency.

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