NASA CR·175022 

i 

\ 

\ NASA-CR-175022 

\ 19860011026 

L------~ 

Rocket Thrust Chamber 

Thermal Barrier Coatings· 

Final Report 

under contract NAS3·23262 

m SOLAR 

TURBINES 

INCORPORAtED 

-111111111111111111111111111111111111111111111 .. 

NF01233 

LIBRARY COpy 

MAR 26 19815 

.LAI~GlEY rlESE,\RCH CENTER 

LIBRARY, NASA 

Ht\Mt:TOf'!!. VIRGIWI!. 
NASA CR-175022 

Rocket Thrust Chamber 

Thermal Barrier Coatings 

Fi nal Report 

under contract NAS3·23262 

by 

A.P. Batakis 

J.W. Vogan 

for 

National Aeronautics & Space Administration 

Communications & Propulsion Center 

21000 Brookpark Road 

Cleveland, OH 44135 

July 1985 

SR84-R-5052-16 

m SOLAR 

TURBINES 

INCORPOR~TEC 

SUBSIDIARY OF CATERPILLAR TRACTOR CO. 

P.O. Box 85376. San Diego. CA 92138-5376 

The words SOI.f, Mil'S, Cllnteur, S.turn and Tit.n are Trademarks of Solar Turbines Incorporated. 

tB is a Trademark 0' CaterpIllar Tractor Co. SpecifiCations subject to chatlge wilhoul nollce. 

Printed in U.S.A. / 

0' 

1. Report No. 2. Government Accession No. 

CR-175022 

4. Title and Subtitle 

Rocket Thrust Chamber Thermal Barrier Coatings 

7. Author(s) 

A. P. 8atakis and J. W. Vogan 

9. PerforminQ.. Organization Name and Address 

Solar lurbines Incorporated 

P.O. Box 85376 

San Diego, CA 92138-5376 

3. Recipient's Catalog No. 

5. Report Date 

July 1985 

6. Performing Organization Code 

8. Performing Organization Report No. 

SR85-R-5052-16 

10. Work Unit No. 

11. Contract or Grant No. 

NAS 3-23262 

13. Type of Report and Period Covered 

~12~.~Sp-o-ns-o-'ri-ng-A~g-en-c-y~N-am-e-a-nd77Ad~d~re-ss--------------------------------I Final Report 

National Aeronautics & Space Administration Dec. 1981 - Oct. 1984 

Communications & Propulsion Center 

21000 Brookpark Road 

14. Sponsoring Agency Code 

Cleveland, OH 44135 5310 

15. Supplementary Notes 

16. Abstract • A research program was conducted to generate data and develop analytlcal techniques to predict the performance and reliability of ceramic thermal barrier coatings in high heat flux environments. A finite element model was used to analyze 

the thermomechanical behavior of coating systems in rocket thrust chambers. Candidate coating systems (using a copper substrate, NiCrA1Y bond coat and Zr02.8Y203 

ceramic overcoat) were sel ected for detail ed study based on photomicrograpni c 

evaluations of experimental test specimens. The effects of plasma spray application 

parameters on the material properties of these' coatings were measured and the 

effects on coating performance evaluated using the finite element model. Coating 

design curves which define acceptable operating envelopes for selected coating systems were constructed based on temperature and strain limitations. 

Spray gun power level was found to have the most significant effect on coating 

structure. Three coating systems were selected for study usi ng different power 

levels. Thermal conductivity, strain tolerance, density, and residual stress were 

measured for these coatings. Analyses indicated that extremely thin coatings 

«0.02 mm) are required to accommodate the high heat flux of a rocket thrust chamber 

and ensure structural integrity. High heat flux coatings are extremely sensitive 

to material property and thickness variations; thus precise control of these variables is essential to successful coating utilization. Coating tests indicate that 

the bond-substrate interface may be the weak link in the coating system. Further 

development of the analytical model and the plasma spray techniques are recommended. 

17. Key Words (Suggested by Author(s)) 18. Distribution Statement 

19. Security Classif. (of this report) 20. Security Classlf. (of this page) 21. No. of pages 22. Price" 

Uncl ass ifi ed Un c 1 ass if i ed 164 

"For sale by the National Technical Information Service, Springfield, Virginia 22161 

TABLE OF CONTENTS 

Section 

1 INTRODUCTION 1 

3 

3 

2 PROCEDURES 

2.1 Laboratory Materials and Procedures 

2.1.1 Powder Specification and Preparation 

2.1.2 Plasma Variables 

2.1.3 Property Measurement 

2.1.4 Calorimeter Testing 

2.1.5 Residual Stress and Stress Free Temperature 

2.1.6 Substrate Temperature 

2.2 Analytical Procedures 

2.2.1 Coating System Performance Analysis 

2.2.2 Coating Structure Analysis 

2.2.3 Failure Strain Analyses 

2.3 Coating of Subscale Thrust Chamber 

3 RESULTS 

3.1 Plasma Deposition Variables 

3 

4 

5 

7 

8 

10 

10 

10 

14 

16 

17 

21 

21 

3.1.1 Effect of Plasma Gas Flow on Coating Characteristics 21 

3.1.2 Standoff Distance Effects 28 

3.1.3 Arc Voltage Effects on Coating Structures 31 

3.1.4 Effect of Arc Current on Structure 38 

3.1.5 Anode Selection Effects 39 

3.1.6 Effect of Traverse Rate 43 

3.1.7 Powder Feed Rate 43 

3.2 Property Measurement 

3.2.1 Density 

3.2.2 Thickness Control 

3.2.3 Thermal Diffusivity and Conductivity 

3.2.4 Surface Finish 

3.2.5 Crystal Structure 

iii 

45 

47 

48 

48 

48 

48 
Section 

TABLE OF CONTENTS (Continued) 

3.2.6 Calorimeter Tests 

3.2.7 Residual Stress and Stress Free Temperature 

3.3 Analytical Modeling 

3.3.1 Thrust Chamber Model Development 

3.3.2 Coating Structure Model Development 

3.3.3 Discussion of Model Development 

3.4 Computer Analyses 

3.4.1 Coating Strength Data Analyses 

3.4.2 Performance Analyses of Actual TBCs 

3.4.3 Coating Design Technique 

3.4.4 Discussion for Analytical Results 

3.5 Test Thrust Chamber Coating 

4 CONCLUSIONS 

5 RECOMMENDATIONS 

REFERENCES 

APPENDIX 1 - Coating Stress/Strain Analysis for Four Point 

Fl exure Test 

APPENDIX 2 - Examples of Finite Element Thrust Chamber Analyses 

Input and Output 

APPENDIX 3 - Photographs of Test Specimens 

iv 

Page 

51 

53 

56 

56 

62 ~) 

73 

74 

74 

83 

94 

96 

97 

107 

109 

111 

113 

117 

159 
LIST OF FIGURES 

Figu~ Page 

1 Schematic of Four Point Flexure Test 6 

2 Calorimeter 7 

3 Residual Stress Model 8 

4a Rocket Thrust Chamber 11 

4b Thrust Chamber Model 11 

5 Finite Element Model of a Porous Coating 15 

6 Two-Gun Configuration 18 

7 Coating Fixture 19 

8 NiCrA1Y Coat Applied for 12 Cycles at a Gas Flow of 1.70 

m3/Hour 23 

9 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

1.84 m3/Hour 23 

10 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

1. 98 m3 /Ho ur 24 

11 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.12 m3/Hour 24 

12 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.15 m3/Hour 25 

13 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.26 m3/Hour 25 

14 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.40 m3/Hour 26 

15 NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.40 m3/Hour 26 

v 
Figure 

16 

17 

18 

19 

20 

21 

22 

23 

24 

25 

26 

27 

28 

29 

30 

31 

32 

33 

34 

LIST OF FIGURES (Continued) 

NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas Flow of 

2.69 m3/Hour 

Baseline Thermal Barrier Coating 

Thermal Barrier Coating with Gun Stand-Off Distance of 

9.0 cm 

Thermal Barrier Coating with Gun Stand-Off Distance of 

·7.6 cm 

Thermal Barrier Coating with Gun Stand-off Distance 

of 11.4 cm 

Thermal Barrier Coating with Plasma Arc Voltage of 

66 Volts 

Thermal Barrier Coating with Plasma Arc Voltage of 

72 Volts 

Substrate After 12 Traverses at 45 Volts 

Substrate After 12 Traverses at 50 Volts 

NiCrA1Y Deposit After 12 Traverses at 55 Volts 

NiCrA1Y Deposit After 12 Traverses at 60 Volts 

NiCrA1Y Deposit After 12 Traverses at 65 Volts 

NiCrA1Y Deposit After 12 Traverses at 70 Volts 

NiCrA1Y Deposit After 12 Traverses at 75 Volts 

Thermal Barrier Coating with Plasma Arc Current of 

350 Amperes 

Thermal Barrier Coating with Plasma Arc Current of 

475 Amperes 

Baseline Coating Applied with a GH Anode 

Bond Coat Deposited at 42 kW with a Type 700 Anode 

Bond Coat Applied at 45 kW with a Type 703 Anode 

vi 

27 

28 

29 

30 

30 

32 

32 

34 

35 

35 

36 

36 

37 

37 

38 

39 

41 

41 

42 
Figure 

35 

36 

37 

38 

39 

40 

41 

42 

43 

44 

45 

46 

47 

48 

49 

50 

51 

52 

53 

LIST OF FIGURES (Continued) 

Zirconia Applied to the Baseline Bond Coat with a 

704 Anode at 58 kW 

Bond Coat Applied with a Type 704 Anode at 58 kW 

Thermal Barrier Coating with Powder Feed Rate of 

0.04 gms/sec. 

Thermal Barrier Coating with Powder Feed Rate of 

0.062 gms/sec. 

Thermal Barrier Coating with Powder Feed Rate of 

0.09 gms/sec. 

Thermal Barrier Coating Applied at 20 kW 

Thermal Barrier Coating Applied at 32 kW 

Thermal Barrier Coating Applied at 40 kW 

Thermal Diffusivity of Zr02Y203 Layers 

Thermal Conductivity of Zr02Y203 Layers 

Water-Cooled Calorimeter Probe 

Calorimeter Face in Guard Shield 

Calorimeter Face with 40 kW Zr02Y203 Coating 0.10 mm 

Thick Exposed to a Heat Flux of 9.15 watts/mm2 

Heat Flux from a Plasma Torch as a Function of Distance 

at Selected Power Levels 

Heat Flux Transmitted to a Copper Calorimeter at 32 kW 

Through a 0.10 mm, 20 kW Zirconia Coating 

Heat Flux Transmitted to a Copper Calorimeter at 32 kW 

Through a 0.10 mm, 32 kW Zirconia Coating 

Heat Flux Transmitted to a Copper Calorimeter at 32 kW 

Through a 0.10 mm, 40 kW Zirconia Coating 

Temperature Gradient for 1650°C Wall Temperature 

Axial Stress Gradient for 1650°C Wall Temperature 

vi i 

Page 

42 

42 

43 

44 

44 

46 

46 

47 

50 

50 

51 

52 

52 

53 

54 

54 

55 

57 

58 
Figure 

54 

55 

56 

57 

58 

59 

60 

61 

62 

63 

64 

65 

66 

67 

68 

69 

70 

71 

72 

LIST OF FIGURES (Continued) 

Thrust Chamber Radial Stress Gradient for 1650°C Inner 

Wall Temperature 

Thrust Chamber Temperature Gradient for 1650°C Inner 

Wall Temperature 

Thrust Chamber Axial Stress Gradient for 1650°C Inner 

Wall Temperature 

Relationship of Coating Surface Temperature and Stress 

to Coating Thickness at Constant Heat Flux 

Initial Predicted Heat Flux vs. Coating Thickness at 

Constant tlT. 

Influence of Microstructure on Effective Thermal 

Conductivity 

Finite Element Analysis of Porosity Effect on 

Effective Thermal Conductivity of Coatings 

Illustration of Mixture Having a Continuous Major 

Phase and Dispersed Minor Phase 

Radiation 

Theoretical Thermal Conductivity of Porous Coating 

Coating Strength Data from Four-Point Flexure Tests 

Average Coating Failure Strain Versus Plasma Spray 

Power Level 

Survival Probability Versus Coating Strain 

Coating Strength Weibull Plots 

Predicted Coating Temperature Profile 

Predicted Hoop and Axial Coating Stress Profile 

predicted Radial Stress Profile through Coating Using Average Measured Thermal Conductivity 

Coating Temperature Profile 

Hoop and Axial Coating Stress 

viii 

58 

60 

60 

61 

62 

63 

65 

66 

67 

71 

76 

80 

82 

82 

85 

86 

87 

87 

88 
73 

74 

75 

76 

77 

78 

79 

80 

81 

82 

83 

84 

85 

86 

87 

88 

89 

90 

91 

92 

LIST OF FIGURES (Continued) 

Effect of Coating Thickness on Surface Temperature for 

Fixed Heat Flux 

Effect of Coating Thickness on Maximum Compressive 

Coati ng Strai n 

Effect of Thickness Variations on Surface Temperature 

Coating Surface Temperature Versus Coating Thickness for 

Various Heat Fluxes Using Average Measured Thermal 

Conductivity 

Coating Strain Versus Coating Thickness for Various 

Heat Fl uxes 

Heat Flux Versus Coating at Constant ~T 

Heat Flux Versus Coating Thickness for Two ~Ts 

Maximum Heat Flux Versus Coating Thickness at 1% Strain 

NASA Thrust Chamber for Test Coating 

Photomicrograph of 42 kW Bond Coat (Ring 49) 

Photomicrograph of 42 kW Zirconia-8% Yttria Coating 

for Ring 47 

Overall Thrust Chamber Coating Set-Up 

Coating a Thrust Chamber 

First Coating Applied to a Thrust Chamber with a 7MBT 

Angle Extension 

Coating Applied with Periodic Reversaing of Cylinder 

Coating Applied to a Planned Thickness of 0.013 mm 

Stress/Strain Model for Plastically Deforming Four Point 

Bend Specimen 

Specimens Prior to Testing 

Two Bend Specimens After Test 

Compressive Coating Failure 

ix 

89 

90 

91 

92 

93 

94 

95 

96 

98 

100 

101 

103 

103 

104 

104 

105 

116 

162 

162 

163 
Fi_gur~_ 

93 

94 

95 

LIST OF FIGURES (Continued) 

Tensile Coating Cracks 

Specimens After Test 

Bend Specimen After Test 

x 

Pag_~ 

163 

164 

164 
Table 

1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14a 

14b 

15 

16a 

16b 

17 

LIST OF TABLES 

Material Properties from the Literature (Refs. 10-16) 

Basic NiCrAl Test Parameters 

Effect of Plasma Gas Flow on Deposit Rate of NiCrA1Y 

Baseline Ceramic Coating Spray Parameters 

Baseline, Angle Expansion Spray Parameters 

Effect of Voltage on Bond Coat Structure 

Plasma Spray Anode Performance Characteristics 

Effect of Anode Selection 

Plasma Spray Parameters Selected for Coating Property 

Measurements 

Thermal Diffusivity and Conductivity of Zr02Y203 Layers 

Residual Stress and Stress Free Temperature Resutls From 

Coated Flat Specimens 

Values of ¢m for Various Mixtures 

Values of KE for Various Mixtures 

Four-Point Bend Specimen Results (Data Set 1) 

Four-Point Bend Specimen Results (Data Set 2) 

Coating Strength Data for Statistical t-Test Analysis 

Matched Pair t-Test Calculations (Data Set 1) 

Matched Pair t-Test Calculations (Data Set 2) 

Average Coating Strength as a function of Plasma Spray 

Power Level 

xi 

13 

22 

22 

27 

33 

33 

40 

40 

45 

49 

56 

69 

70 

75 

75 

78 

78 

79 

81 
Table 

18 

19 

LIST OF TABLES (Continued) 

Specimen Definition in Final Ring Coating Tests 

Plasma Spray Parameters for Flat and Ring Specimens 

xii 

Page 

99 

102 
1 

INTRODUCTION 

The objective of this program is to develop a zirconia based thermal barrier 

coating (TBC) system for use in high heat flux rocket applications such as 

the Space Shuttle main engine (SSME) thrust chamber. The technology developed 

is also pertinent to other appl ications such as hot components in the gas 

turbine engine. Previous TBC development and test has concentrated on ceramic 

overlays ranging from 0.25 mm to 0.38 mm in thickness. The high heat fluxes 

encountered in high pressure rocket engines dictate that thinner, 0.012 mm to 

0.10 mm, coatings be used to avoid surface melting and mechanical failure of 

the coating. 

The plasma sprayed TBC required in this application is a co~plex structure. 

No currently accepted model is known that relates coating performance to 

thickness and to controllable process variables. The use of very thin coatings, such as those under consideration, is particularly complicated. Four 

examples of this complexity are given below. These illustrate some of the 

difficulties that were anticipated in the practical and the analytical 

development program. 

1. Structures are frequentl y two phase (e.g., monoel ini c pl us transformabl e tetragonal in the case of zi rconi a) and may vary from 

point to point. 

2. Each successively deposited grain impacts a surface whose temperature may increase as coating thickness increases. 

3. Defects (e.g.,pores) are of significant size in relation to coating 

thi ckness so that bul k or average coating properties may not be 

appropriate. 

4. Individual grain size of the plasma spray powder approaches total 

coating thickness as does the normal surface roughness of plasma 

sprayed coatings. 

1 

2 

PROCEDUR·ES 

2.1 LABORATORY MATERIALS AND PROCEDURES 

The coating system selected for evaluation in this study was a copper substrate, NiCrA1Y bond coat and yttria stabilized zirconia ceramic thermal 

barrier. This bond coat/ceramic combination has been extensively investigated 

as a TBC for heat engines and shows great promise as a reliable coating. 

The coating deposition performed in this study was accomplished using a Metco 

7M plasma system. During the program, deposition parameters such as gas flow, 

power levels, standoff, etc. were varied. The methods used to define these 

vari at; ons, and tests used to establ ish the resulti ng effects on coati ng 

properties, are given in this section. 

2.1.1 Powder Specification and Preparation 

The powder materials used in this study were as follows: 

Bond Coat = Amdry powder #961 

18.3% Cr, 5.8% Al , 0.5% Y, bal. Ni 

-170 + 325 mesh 

Ceramic TBC = Cerac Zirconium Oxide-Yttria Stabilized powder 

Zr02. 8Y 203 

-325 mesh 

The ceramic powder was determined to have the following size distributions: 

+100 mesh - trace 

-100 + 140 mesh - 4% 

-140 + 200 mesh - 30% 

-200 + 325 mesh - 61% 

-325 mesh - 5% 

Early tests indicated that narrowing the particle size range within that of 

the starting material had a negligible effect on coating structure. Thus, 

the as-received material was used in this study. 

3 
2.1.2 Plasma Variables 

The following deposition parameters were varied during the test program. The 

effects of each variable on coating structure were evaluated by metallurgical examination. 

Arc Current 

During the program, current to the plasma gun was varied over a range of 300 

to 1000 amperes. The current was measured at the control console using a 50 

millivolt shunt and meter. Calibration tests indicated that, with manual 

correct i on for drift by the operator, an accuracy of +25 amperes coul d be 

maintained during a test run. 

Arc Voltage 

Arc vol tage was al so moni tored at the control conso1 e and these val ues are 

reported to an accuracy of +2.5 volts. Voltage measurements at the gun input 

connections showed a voltage drop of 1 to 3 volts in the water cooled power cables. The voltage drop to the gun was dependent upon the current. 

Plasma and Carrier Gas Flows 

The flow of pl asma on carri er Porter gas was control 1 ed by Fi sher purge 

meters. Pressure at the meters was maintained at 689 kPa. Flow was regulated 

by needle valves giving a reproducibility of +5%. 

Powder Feed Rate 

Powder feed control util ized a Model 1250 Plasmadyne hopper. This unit 

incorporates a metering disc with a digital speed readout to introduce powder 

into the carri er gas stream. Del ;very rates were peri odi cally verifi ed by collecting and weighing the powder from the delivery tube over a period of 

thirty minutes. In addition, the initial weight of powder introduced into 

the cannister was recorded and at the completion of the test, the weight of 

powder remaining was established. The difference in these two weights was 

recorded. The latter method was subject to errors resulting from material 

retained in the powder feed system after emptying the cannister. 

4 
Deposition Rates 

The rate of coating deposition was established by direct measurement. The 

plasma gun was traversed at a constant rate for a specified number of cycles. 

Thickness measurements were made after every five cycles to determine if 

deposit rates were constant. At completion of the test, metallurgical sections were made to verify coating thickness. 

Nozzle Effects 

The effect of changing the pl asma nozzl e (anode) were determined using the 

vari ous confi gurati ons suppl i ed for the Metco 7M system. No modifi cat ions 

were made in the base or configuration. Data suppl ied by the manufacturer 

were used as a guide in establishing basic operational characteristics. 

2.1.3 Property Measurement 

Density/Porosity 

Five methods were used to determine coating density. In initial trials, test 

specimens were weighed before and after coating. The increased weight was 

attributed to the coating deposit. Coating thickness and area measurements 

were then made, the apparent volume calculated, and the apparent density 

(weight divided by volume) reported. 

Porosity of the coating was also determined by immersing pre-weighed coated 

spec imens in boi 1 i ng water. The heat was removed and the wate r coo 1 ed to 

room temperature. The specimens were then removed from the water, excess 

water removed and the specimen rapidly weighed. The gain in weight was attributed to water entrapped in pores, the pore volume was then calculated using the known density of water. 

The third and fourth methods used for determining coating density required the preparation of metallurgical coating sections. Test specimens were 

sectioned and photomicrographs were made, usually at lOOX magnification. 

From these photomicrographs, porosity was established by photometrically 

comparing average reflectance of the coating section with the nonporous substrate and by direct measurement of the average pore area in a unit area of 

coating. 

The fifth method used to determine density was by direct measurement. The 

area of a rectangular coated specimen was measured. The substrate and bond 

coat (if used) were carefully removed by chemical etching and the coating was 

weighed after rinsing and drying. The specimen was then mounted and polished 

for metallographic examination and coating thickness measurement. The volume 

of the coating was then calculated from this data and the density calculated 

using the previously determined coating weight. 

5 
Coating Thickness 

Because of the extremely thin coatings being evaluated, thickness measurement 

required a high degree of precision. During the coating process, measurement 

of coating thickness was approximated using a micrometer. These measurements 

were limited to +0.013 mm by the roughness and irregularities in the coating. 

More precise measurements were made after coating the specimens using a metallurgical section and a microscope with a calibrated eyepiece. 

Coating Strength 

Coating strength was determined using a four paint bend test (see Fig. 1). 

The analytical methods used to determine the coating stresses developed 

during this test are described in Appendix 1. 

Copper specimens 0.81 mm thi ck x 13 mm wi de x 76 mm long were coated on one 

s ide with a 0.03 mm Ni CrAlY bond coat. A yttri a stabil i zed TBC was then 

applied over this bond coat at selected thicknesses ranging from 0.02 mm to 

0.15 mm. The specimen was then placed on two knife edge supports spaced a 

distance of 51 mm apart. For tensile loading of the coating the coated 

surface was placed on the supports and for compressive coating loads the 

coated surface was placed opposite from the supports. The specimen was then 

loaded symmetrically by knife edges spaced 25 mm apart at a rate of 0.08 

mm/second until coating failure occurred as evidenced by an abrupt decrease 

in the load required to maintain a constant rate of strain. Deflection was 

initially measured using a dial gauge mounted directly below the mid-point of 

the specimen. Subsequent tests used crosshead travel during loading to 

monitor specimen deflection. 

P/2 1_"'''-- 2.5 em -_.-1 P/2 Coating Test Section 

0.025 

to F================================i.========1r=:::I:::==:::;:= 0.100 em 

Copper Substrate 

1-... -------5.1 cm------'., __ I 

P = Applied Load 

Side View Front View 

Figure 1. Schematic of Four Point Flexure Test 

6 
Thermal Diffusivity and Conductivity 

Thermal conductivity of the coatings were measured at Purdue Laboratories 

using a laser heat pulse technique. A detailed discussion of this technique 

is given by Taylor (Ref. 1). 

For thi s measurement, copper specimens 13 mm x 13 mm x 1.6 mm thi ck were 

coated on one side with the TBC to be tested. Coating thickness was a nominal 

0.18 mm. A thermocouple was attached to the uncoated surface of the specimen, 

the assembly furnace heated to the desired test temperature and a laser heat 

pulse applied to the coating. Temperature rise of the back face as a function 

of time was measured. This, combined with coating and substrate thicknesses, 

densities and specific heats enabled the calculation of diffusivity and 

conductivity using the methods described by Taylor in the previously cited 

reference. 

2.1.4 Calorimeter Testing 

A water cool ed calorimeter was used to measure average heat fl ux from the 

plasma torch, heat flux during coating deposition and the reaction of the 

coatings to high heat inputs. This calorimeter, shown in Figure 2, consists 

of two water cooled copper components. The fi rst is the cent ra 1 test a rea 

and the second is the outer guard ring used to shield areas other than the 

test surface from external influences. In operation, a constant metered flow 

of water is supplied to the test area, and water inlet and outlet temperatures 

are monitored. The calorimeter is inserted into the plasma flame and the 

heat absorbed by the cool i ng water is presumed to be the total heat input 

over the test area. 

The calorimeter data is then converted to heat flux in the following manner: 

Q = (0.279) (W) (,1T) 

1Td 2 

Plasma Sprayed AI203 Seal 

Thermocouple 

~~~~~~~~~~~~~~=~ 7~ Water 

Guard 

Ring 

Figure 2. Calorimeter 

7 

In 

Water 

Out 
where: 

Q = heat flux (watts/mm2) 

~T = Tinlet - To~tlet (OC) W = water flow 19rams/min.) 

d = test diameter (mm) 

2.1.5 Residual Stress and Stress Free Temperature 

Residual stress in the coating was evaluated using the techniques described by 

Andrews (Ref. 2). The model is shown in Figure 3 and the symbol definitions 

are given below: 

Subscript (1) ceramic 

Subscript (2) copper 

a = coefficient of thermal expansion 

E = modulus of elasticity 

I = moment of inertia 

h = total composite thickness 

a1 = ceramic thickness 

a2 = copper thickness 

L = active length 

r = radius of curvature = L2/2 

o = tangential deflection 

S1 = stress in ceramic 

L 

Deflection 

----/ 

Coating 

Substrate 

Figure 3. Residual Stress Model 

8 
b = ~T = 

TSF = 

specimen width 

T SF - T gmbient stresS-Tree temperature 

This approach was initially developed for calculation of stresses in porcelain enamels and fits well with the requirements of thin TBC's. The methodology is 

described by the following equation: 

-20 [2 aIEl] Sl = 2 - (E1I1 + E2 12) + -

L hal 2 

(1 ) 

Since deflection (0) can be measured directly, equation (1) is easily applied. 

For most calculations Ell1 and aIEl are small and can be omitted from the 

calculation, reducing equation (1) to: 

$1 

-40 E212 = 

L2ha1 

(2 ) 

where 

I = b a3 

12 (3 ) 

The stress-free temperature of the system is calculated from the bi-metallic 

relationship: 

0 = 

where 

K = 

and 

m = 

n = 

Hence: 

~T = 

(a2 - a1) (~T)L2 

K 

h 

3(1 + m)2 

3(1 + m)2 + (1 + m) 

a 1/a2 

E1/E2 

h 

(a2 - a1)L2K 

1 

(m 2 + mn) 

9 

(4 ) 

(5 ) 

(6 ) 

(7) 

(8 ) 
and 

(9) 

For the purposes of these calculations the following constants were used: 

El = 45 x 106 kPa 

E2 = 117 x 106 ~Pa a1 = 7.15 x 10- JOC 

a2 = 17.6 x 10-6joC 

The following procedures were used to obtain the deflection values (5) used 

in these calculations. Copper test strips, 0.8 mm x 13 mm x 76 mm, were bond 

coated with 0.03 mm of NiCrA1Y on one side. The deflection, or chord height, 

if any, was determined over the 50 mm midspan (L). The specimen was then 

coated wi th the zi rconi a TBC and the measurement repeated. Any change in 

curvature was recorded and the resulting value used to calculate the residual 

stress in the coating. 

Stress free temperatures were al so calculated from this data. The results 

were then confirmed by furnace heating the specimen to the calculated temperature and determining if it returned to its original configuration. 

Similar tests were performed using copper rings 68 mm in diameter rolled and 

butt welded from 13 mm wide, 0.8 mm strips. The rings were vacuum annealed 

at 760°C after bond coating to relieve initial forming and coating stresses. 

After coating and segmenting, the change in diameter of the ring, due to the 

coating process, was established and used in Equation (1) for an alternate 

means of calculating residual stress. 

2.1.6 Substrate Temperature 

During trial coating appl ;cation, the substrate temperature was monitored 

using an Inconel sheathed chromel-alumel thermocouple, 0.76 mm in diameter. 

The thermocoupl e was install ed in the copper substrate by fi rst drill i ng a 

0.77 mm hole through the specimen and then inserting the thermocouple. The 

edges of the hole were then staked to retain the thermocouple with its tip 

flush with the surface to be coated. Then a NiCrA1Y bond coat was applied to 

bond the tip of the thermocouple to the substrate. Temperature recording was 

done with a conventional strip chart recorder. 

2.2 ANALYTICAL PROCEDURES 

2.2.1 Coating System Performance AnalysiS 

Fi nite el ement analyses, usi ng the ANSYS code, were used to eval uate coati ng 

system performance under simulated operating conditions. The coating system 

model used for these analyses (shown in Fig. 4) simulates the coated thrust 

10 
OXYGEN _ 

HYDROGENCopper 

Bond 

Ceramic 

R 

1 • 

• 

• 

• 

• 

~ 

-c::::::::: 

-=:::::: 

I. 

w 

COMBUSTION CHAMBER EXHAUST 

Figure 4a. Rocket Thrust Chamber 

TCoid 

tw 

t t Loj t t =t THot 

Q di 

... X 

Engine Centerline 

Figure 4b. Thrust Chamber Model 

chamber wall of a rocket engine. This two-dimensional, axisymmetric model 

represents a cross-section of the wall at the throat of the thrust chamber. 

The el ements used in thi s fi n;'te el ement model were two-dimensi onal, i soparametri c, axi symmetri c ri ng el ements. Each el ement was defi ned by four nodal 

11 
points and each node was defined by a radius and an axial location. For 

thermal analyses, each node had a single degree of freedom: temperature. 

For structural analyses, each node had two degrees of freedom: translations 

in the radial and axial directions. 

The finite element geometry for the thermal analyses was identical to that 

for the structural analyses. This allowed the temperature distributions 

which were derived from the thermal analyses to be directly translated to the 

structural model. 

The basic geometry of the thrust chamber model consists of a 66.04 mm internal 

di ameter (di) copper cyl i nder with a 0.889 mm wall (tw). Attached to the 

inner surface of this cylinder is a 0.038 mm thick (tb) NiCrA1Y bond coat, 

and a variable thickness (tc) Zr028Y203 ceramic coating. Ceramic coating 

thicknesses from 0.0127 mm to 0.203 mm were investigated. The axial length 

(J) of the model was 0.229 mm, although a 50.8 mm length was used in some of 

the preliminary analyses. 

The finite element model was divided into nine equal elements in the axial 

direction, to accommodate axial variations in boundary conditions. In the 

radial direction, the ceramic and bond coats were divided into three element 

layers each, and the copper substrate was divided into six. This element 

configuration resulted in element aspect ratios (axial length to radial 

thickness ratios) from 0.375 to 6 in the ceramic coating, 2 in the bond coat 

and 0.17 in the copper substrate. Use of the long (50 mm) preliminary model 

resulted in element aspect ratios from 83 to 667 in the ceramic and 444 in 

the bond coat. It is generally desirable to maintain an aspect ratio between 

1 and 4, to ensure numeri cal accuracy. The effect of aspect ratio on the 

accuracy of these analyses was evaluated by comparing results using the long 

and short finite element models. 

The three coat; ng system mater; al s were assumed to be homogeneous and ; sotropic. Properties of all three materials were initially obtained from the 

literature. These properties are listed in Table 1. In the final analyses, 

measured material properties were substituted for literature values where 

applicable. The temperature dependency of these material properties created 

a non-linear problem. This required iterative analyses, with the solutions 

defined using a specified measure of convergence. 

A vari ety of thermal boundary conditi ons were sel ected for these coati ng 

systems analyses. These conditions were intended to simulate the range of 

anti ci pated operati ng conditi ons ina typi cal hi gh pressure rocket engi ne 

thrust chamber and to define an operating envelope based on coating material 

limitations. Thermal boundary conditions were established in two ways: (1) 

a hot surface temperature and a cold surface temperature were defined or (2) 

a heat flux and a cold surface temperature were defined. 

The outer surface of a typical rocket thrust chamber is cooled with liquid 

hydrogen, which has a boiling point of -240°C at ambient pressure. Test data 

indicate that the actual cold surface temperature of a rocket chamber during 

operation is between 0 and 200°C. For these analyses, three cold surface 

temperatures were evaluated: -240, 0, and 200°C. Heat fluxes in the range 

12 
-

w 

Material 

Zr02 8Y203 

Ceramic Coating 

Solid Zirconia 

(Stabilize d 

zr02 ) 

NiCrA1Y Bond 

Coat 

Oxygen Free 

Copper 

Substrate 

Tab 1 e 1 

Material Properties From the Literature (Refs. 10-16) 

Thennal 

Thennal Elastic \ Expansion 

Conductivity Modulus Specific Heat Coefficient 

(WinK) (GPa) (J /kg.K) (x 10-6 /C) 

2.2 x 10-4 T + 1.09 45.06 - 0.0248T 317 + 0.813T 7. 1844+0.003888T 

(400<T<22000K) (90 oC<T<12000C) -9.55 x 1 0-4~ (90 oC<T<12000C) 

+ 6.30 x 10-7~ 

(200 oK<T<15000K) 

2.708 - 5 x 10-4 T 205 400 to 640 8.6 to 15.2 

(70°C < T < 1100 0C) 

8.3 x 10-3T + 6.7 125.6-0.0149T 0.476 + 0.00086T 13.3182-0.0121T 

(400 oK<T<14000K) (90 oC<T<5500C) -0.0000015T2 + 5.8729E-5T2 

(90 oC<T<11000C) (90 oC<T<11000C) 

391 117 0.39 5.4 

Density Poisson's 

( g/cc) Ratio 

-- 0.23 

5.63 0.23 to 

to 0.35 

6.27 

6.98 0.23 

8.94 0.350 
typically encountered in rocket thrust chambers, 50 to 165 W/mm2, were also 

used for these analyses. Typically values of 49, 82 and 164 W/mm2 were used. 

Cerami c surface temperatures were evaluated in the range of 1000 to 2600°C 

(Zr028Y203 melting point). A maximum desirable ceramic operating temperature of 1650°C was established based on past experience, and this value was used for 

most of the analyses in whi ch a cerami c surface temperature was specifi ed. 

One structural boundary condition was imposed for these analyses: one radial 

edge of the coating system model was restrained from movement in the axial 

direction. This condition was held for all analyses. No external loads, 

pressures, or displacements were imposed. 

The thermal and structural boundary conditions are indicated graphically in 

Figure 4. Examples of the ANSYS finite element input and output are included 

in Appendix 2. The output from the thermal analyses consisted of heat flow 

rates and temperature distributions. The output from the structural analyses 

consisted of displacements, stresses, and reaction forces. 

2.2.2 Coating Structure Analysis 

The effect of coating microstructure on the thermal conductivity of a coating 

was evaluated using the ANSYS finite element code. A two-dimensional, planar 

model of a coating system was used for these analyses. This model, shown in 

Figure 5, used a biaxial plane thermal element. Each element was defined by 

four nodes having a single degree of freedom, temperature, at each node. 

The model shown in Figure 5 represents a simplified porous coating structure. 

This geometry was selected to facilitate the finite element modeling. The 

model was essentially composed of "blocks" of coating mat.erial. Pores were 

simulated by eliminating some of these blocks, in specified patterns. 

The model used in these analyses simulates the basic coating system being 

considered by NASA for rocket thrust chambers: a 0.889 mm thick copper substrate with a 0.038 mm NiCrAlY bond coat and a variable thickness Zr028Y03 

ceramic coat. The material properties that were used are presented in Table 1. 

The value of thermal conductivity for the ceramic in Table 1 was used for the 

solid ceramic elements in the finite element model. 

Boundary conditions for this model were established by specifying the hot 

cerami c surface temperature and the col d copper surface temperature. To 

maintain an accurate temperature gradient along the edges of the coating, the 

nodes along each edge were coupled to similar nodes within the coating. 

The effect of radiation across the individual pores was investigated using 

this model. Radiation links were specified between the upper and lower surfaces of the pores. Radiative heat transfer was calculated from the standard 

radiation function defined as follows: 

q = (J Z F A (Ti - T~) 

14 
Pores~ 

I 116-f~~~~ 118 119 120 121 

-/- 110 111 112 113 114 115 1.--1...,0....,..3-1---,1~04 105 106 107 108 109 I 

122 

97 98 99 100 101 102 

90 91 92 93 94 95 96 Ceramic 

...--__ -l-....!8~4~-I-~8~5_1--8~6~t_~;:;_+__;8~7_t-...::8:.::::8:....-t__i.89rt_ltrl (Variable 

Lj7U7~-l-_7l.!8L-I-~ ___ -+-~:9.79~t_~80!L..+_~8~1_t-""';i:-t_~82~t_--:;83=--...1 Thickness) 71 72 73 74 75 76 

64 65 66 67 68 69 70 I 

58 59 60 61 62 63 

51 52 53 54 55 56 57 I 

1~~1\7a~7a~/a~~O ~I~ 1 Bond 

/17~7-~ Y2~y2~1 O.038mm ___ -----;;-:-::-::--::--------------.J} Substrate _ C~=~~---~=_==~ - __ -_------..:C:.:o:.!p::.:::p:,:e,;,..r 

______ -=~ __________________ ----------~~---.....,l ~_ 0.89 mm ---- -

Figure 5. Finite Element Model of a Porous Coating 

where: 

a = Stefan-Boltzmann constant ~ = emissivity 

F = geometric form factor 

A = area 

q = heat flow rate 

Tl = temperature of hot surface 

T2 = temperature of cold surface 

The effect of coating microstructure on the thermal conductivity of ceramic 

coatings was investigated using this finite element model. Various coating 

structures were evaluated, having different pore sizes, pore shapes, pore 

densities, etc. Each coating structure was characterized by three geometric parameters: 

1. Pt - total porosity; the volume of pores divided by the total 

coating volume (% of total volume) 

2. Or - dimension ratio; the pore length (axial) divided by the pore thickness (radial) 

3. Pi - pore size; total porosity divided by the number of pores (% of total volume) 

The goal of these analyses was to establish a relationship between the three 

coating variables (Dr' Pt, Pi) and effective thermal conductivity. 

15 
The approach was to subj ect each coating structure to a fi xed temperature 

differential and calculate, using finite element analyses, the heat flow 

through the coating. The effective thermal conductivity of each coating 

structure was then calculated from 

where 

k 

Q 

A 

L\T 

t 

k = Q tjA L\T 

= 

= 

= 

= 

= 

effective thermal conductivity of the coating 

heat flow through the coating 

overall area of coating model perpendicular to heat flow 

temperature drop across coating 

coating thickness 

The effective conductivity of each coating structure was normalized to the 

maximum possible thermal conductivity. This maximum value was established by 

analyzing a coating with no porosity. Thus, each conductivity is expressed 

as a percent of maximum. 

This analytical model of the coating structure did not consider thermal contact resistance between layers of ceramic or between different materials. It 

was anticipated that the effect of thermal contact resistance, and any other 

unaccounted for phenomena, would be reflected in an empirical adjustment of 

the analytical model. Conflicting reports as to the significance of thermal 

contact resistance in TBC's have been observed in the literature (Ref. 3, 4). 

2.2.3 Failure Strain Analyses 

The four-point flexure test was initially evaluated using coating strength data from the 1 i terature (Ref. 5). Thi s reference quoted a compress i ve 

failure strain of 0.78% in a MgO-Zr02 coating, using a similar four-point flexure test. It was calculated that hardened copper could accommodate this 

strain level without plastic deformation. An elastic stress-strain analysis 

was thus derived, for a composite beam, to evaluate the failure strains in 

the coated specimens. 

Preliminary testing with the four point flexure test for measuring coating 

strengths revealed that plastic deformation of the copper substrate occurred 

before coating failure. The initial stress-strain analysis, derived for an 

elastic composite beam specimen, was then replaced by a more sophisticated 

analysis which accounted for plasticity in the copper substrate. This analy- sis is described in Appendix 1 with the computer program created to perform 

the numerical solution. 

16 
2.3 COATING OF SUBSCALE THRUST CHAMBER 

Initial development work on this program was accomplished using flat spec- imens and a standard plasma spray gun configuration. The test chambers were 

66 mm diameter cylinders and were not accessible to the standard gun configuration. A 61 cm small bore extension gun was therefore used to coat these 

chambers. The two gun configurations are shown in Figure 6. 

In the conventional configuration, the electrode, plasma gas flow and coating deposition were along a single axis normal to the coated surface. Gun standoff was varied over a range of 50-100 mm to aid in process control. The 

small bore coating equipment deflects the plasma gas stream at an angle of 

45° to the gun axis, and to the coated surface. Gun standoff was limited by the cylinder 10. Further coating development tests were required to produce 

equivalent coatings with changed equipment. 

The evaluation procedure consisted of first establishing three coating structures using conventional test specimens as described in Section 2.1.2. 

Metallurgical and material property evaluations were used to characterize each 

coating. Test rings 66 mm in diameter, formed from 0.81 mm thick copper 

strips, were then coated using the angle extension. Using the previously 

determined power levels, the gas flows and powder feed rates were varied until 

equivalent structures were obtained with the internal diameter equipment. 

In addition, the ring was periodically reversed relative to the gun entry end 

to compensate for the 45° impingement angle and offset the shadowing effects 

of previously deposited coating particles. 

To verify coating quality on the test cylinder, extensions of the same diameter were attached to each cyl inder and coated simultaneously using the 

settings established by the previous ring tests. After the coating operation 

was completed, these extensions were removed, sectioned and evaluated metallurgi cally. 

The use of thermocouples to measure substrate temperatures during deposit was 

not feasible due to the double wall construction of the test chamber. Temperature recorders* in the form of adhesive strips which indicated the maximum 

temperature by a permanent color change were located in selected locations on 

the thrust chambers. These provided i nformati on on the maximum temperature developed during the coating operation. 

The fixture used to coat the chamber is shown in Figure 7. The thrust chamber 

was located on a variable speed, rotating head. For control of substrate 

temperature, a controlled cooling air flow was introduced through the fuel 

1 i ne fitti ngs on one end. Exhaust ai r temperature was monitored to further 

insure part-to-part repeatability. 

* Manufactured by Tedatemp Corporation, Fullerton, CA 

17 
(A) Metco 7MB Direct Spray Gun 

(B) Metco 7MBT-12 45 Degree Angle Spray Gun 

Figure 6. Two-Gun Configurations 

The procedure consisted of mounting the thrust chamber as shown and abrasive 

blasting the internal diameter while rotating the assembly. The bond coat 

and ceramic overlay were then plasma sprayed using the angle extension. The 

operation was interupted during application of both the bond and ceramic 

coats after a thickness of 0.01 mm had been applied. The chamber was removed 

from the fi xture and remounted on the opposite end to compensate for the 

directionality of the spray and provide a more uniform coating structure. 

This reversal was repeated periodically throughout the coating cycle after 

each 0.01 mm, until full coating thickness was achieved. 

18 
Figure 7. Coating Fixture 

19 

3 

RESULTS, 

3.1 PLASMA DEPOSITION VARIABLES 

Results of tests to determine the effects of plasma deposition parameters on 

TBC's are reported in Sections 3.1 and 3.2. Section 3.1 reports the general 

effects of varying process parameters on the resulting coating. These tests 

were used to select three sets of deposition parameters for material property 

evaluation. The results of the property measurements for the three coatings 

are given in Section 3.2. 

3.1.1 Effect of Plasma Gas Flow on Coating Characteristics 

For these tests the basic parameters presented in Table 2 were used with the 

Metco 7M system for deposition of the NiCrA1Y bond coat. 

The first variable selected was plasma gas flow. The results of these tests 

are presented in Table 3 and Figures 8 through 16. At this time no attempt 

was made to optimize the coating but only to determine the effect of an 

individual parameter (gas flow) on coating structure. 

Based on thickness measurements alone, coating deposit rate was established 

as a function of plasma gas flow. The deposit rate decreased steadily with 

increasing gas flow. At a gas flow of 2.69 m3/hr no measurable deposit could 

be obtained. 

The photomicrographs shown in Figures 8 through 16, illustrate the effect of 

gas flow .on coating structure. At the lowest gas flow (1.70 m3/hr) coating 

buildup was irregular with large voids. The asperities visible in Figure 8 

show the peaks that give rise to the apparent rapid coating buildup. As flow 

was increased, the coating tended towards a denser structure with some large 

voids. At a gas flow of 2.12 m3/hr the spherical unmilled particles observed 

at lower gas flow tended to disappear although some porous regions existed in 

the coating as shown in Figure 11. The most uniform coating obtained with 

these parameters was developed at a gas flow of 2.26 m3/hr. As the gas flow 

was further increased, the coating tended to deteriorate with only the larger particles having deposited. At the maximum flow of 2.69 m3/hr only an occasional deposited particle was observed and the substrate was significantly oxidized. 

Similar tests were performed to determine the effect of plasma gas flow on 

the ceramic coating structure. For these test the parameters in Table 4 were 

used as a starting point. 

21 
Table 2 

Basic NiCrA1Y Test Parameters 

Current 

Voltage 

Argon Flowrate 

Hydrogen Flowrate 

Standoff Distance 

Powder Port No. 

Carrier Argon Flowrate 

Powder Feedrate 

Cooling Air Jet Pressure 

Nozzle Model No. 

Traverse Rate 

Substrate Material 

500 amperes 

60 volts 

Variable 

0.28 m3/hr 

3.1 cm 

#2 

0.28 m3/hr 

5.7 gm/min. 

550 kPa 

713 

2.5 cm/sec 

0.81 mm copper 

Table 3 

Effect of Plasma Gas Flow on Deposit Rate of NiCrA1Y 

Gas/Flow Deposit Rate 

(m3/hr) (mm/traverse) 

1.70 0.016 

1.84 0.012 

1. 98 0.0074 

2.12 0.0064 

2.26 0.0031 

2.40 0.0021 

2.55 0.0010 

2.69 * 

*Too small to measure. 

Monitoring of the substrate temperature during coating deposition was accomp1 i shed usi ng a 1.6 mm Inconel sheathed thermocoupl e inserted into a hol e 

drilled through the specimen. It was located so that its tip was flush with 

the substrate surface. 

A photomicrograph of the baseline ceramic coating ;s shown in Figure 17. The 

average thickness of the ceramic is 0.09 mm. The individual coating particles 

can be distinguished at 1000X magnification. The majority of the particles have been flattened by impact to an average thickness of 0.005 mm. 

22 
Figure 8. 

Figure 9. 

Magnification: 100X 

NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 1.70 m3/Hour 

Ni CrAl Y 

Magnification: 100X 

NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 1.84 m3/Hour 

23 
NiCrAlY 

;._-Substrate 

Magnification: 100X 

Figure 10. NiCrA1Y Bond Coat Applied for 12 Cycles at Gas 

Flow of 1.98 m3/Hour 

+-_ .... ,i CrAl Y 

Magnification: 100X 

Figure 11. NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.12 m3/Hour 

24 
Figure 12. 

Figure 13. 

Magnification: 100X 

NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.15 m3/Hour 

Magnification: 100X 

NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.26 m3/Hour 

25 
Bond Coat 

_--oubst rate 

Magnification: 100X 

Figure 14. NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.40 m3/Hour 

_~ ___ Bond Coat 

.... --Substrate 

Magnification: 100X 

Figure 15. NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.40 m3/Hour 

26 
.___- Bond Coat 

-- Substrate 

Magnification: 100X 

Figure 16. NiCrA1Y Bond Coat Applied for 12 Cycles at a Gas 

Flow of 2.69 m3/Hour 

Table 4 

Baseline Ceramic Coating Spray Parameters 

Durrent 

Voltage 

Argon Flowrate 

Hydrogen Flowrate 

Standoff Distance 

Cooling Air Pressure 

Deposit Rate (avg) 

Traverse Rate 

Powder Feedrate 

Substrate Temperature 

Coating Thickness 

400 amps 

50 volts 

1 m3/hr 

0.23 m3/hr 

10.2 cm 

551 kPa 

0.008 mm 

2.5 cm 

3.0 gms/min. 

97-119°C 

0.10 mm 

Some voids and numerous microcracks are also visible. Increasing the plasma 

gas flow to 1.15 m3/hr and then to 1.43 m3/hr did not produce a visible change 

in microstructure. However, it was observed that this gas flow rate increase 

decreased the rate of deposition by 42 percent. The substrate temperature 

also increased with increasing gas flow, by 80°C in the first case and by 

13°C in the second. 

27 
Magnification: 200X 

Magnification: 1000X 

Figure 17. Baseline Thermal Barrier Coating (Mount No. 5263) 

3.1.2 Standoff Distance Effects 

The effects of standoff distance (distance between gun and substrate) are 

tabulated below. This data was generated using the baseline parameters given 

in Table 4. 

28 
Standoff 

10.2 cm* 

9.0 cm 

7.6 cm 

11.4 cm 

Substrate 

Temperature 

97-119°C 

108-128°C 

144-167°C 

129-153°C 

*Baseline coating 

Deposit Rate 

(per cycle) 

0.008 mm 

0.008 mm 

0.008 mm 

0.003 mm 

Effect on Coating Structure 

Figure 17 

Slight porosity decrease 

Larger interlaminar structure 

Large particles entrapped 

Decreasing the standoff distance to 9.0 cm increased the coating density slightly as shown in Figure 18. The substrate temperature also increased to 

a maximum of 132°C. Fu rther reduci ng the standoff to 7.6 cm increased the 

size of the agglomerated particles and caused wider laminar separations as 

shown in Fi gure 19. Substrate temperature also increased to a maximum of 

167°C during coating. 

Increasing gun standoff to 11.4 cm produced a coating with poor integrity as 

shown in Figure 20. During normal plasma spraying, with the selected spray 

parameters, the cooling jet was adjusted to impinge at the point where the 

plasma hit the substrate. The jet served to cool the substrate and to remove 

unmelted nonadherent particles. 

Magnification: 1000X 

Figure 18. Thermal Barrier Coating With Gun Stand-off Distance of 

9.0 cm 

29 
Magnification: 1000X 

Figure 19. Thermal Barrier Coating With Gun Stand-off Distance 

of 7.6 em 

Magnification: 1000X 

Figure 20. Thermal Barrier Coating With Gun Stand-off Distance 

of 11.4 em 

30 
At 11.4 cm the effectiveness of the jet was greatly diminished. Large unmelted 

parti cl es were entrapped in the coati ng and the substrate temperature i ncreased to a maximum of 153°C, approximately mi dway between the temperature observed at 7.6 cm and 9.0 cm. Also, particle temperatures decreased to the 

point where interparticle bonding was decreased. Oeposit efficiency also 

dropped to 40 percent of that recorded at the baseline standoff. 

Attempts to spray at standoffs greater than 11.4 cm were unsuccessful. No 

measurable deposit was obtained. 

3.1.3 Arc Voltage Effects on Coating Structures 

In the plasma spray system, arc voltage (a dependent variable) is a function 

of the plasma gas composition. As the amount of hydrogen in the argon/hydrogen plasma gas mixture is raised, gas enthalpy, heat transfer and arc voltage all 

increase. For practical reasons, this effect is generally controlled by adding sufficient hydrogen to produce the desired voltage. Data is therefore 

reported on the basis of voltage. 

The baseline parameters in this study are those reported in Table 2 except 

that hydrogen flow was adjusted to vary the voltage. Results are summarized 

below: 

Voltage 

60* 

66 

72 

78 

Substrate 

Temperature (OC) 

97-119 

108-130 

149-171 

143-152 

*Baseline coating 

Oeposit Rate 

(mm per Cycl e) 

0.008 

0.008 

0.007 

0.006 

Effect on Coating Structure 

Figure 17 

None definable 

Slight densification 

Laminar separation 

Increasing the voltage from the baseline of 60 volts to 66 volts had no 

visible effect on coating quality. A substrate temperature increase of 11°C 

was observed. The coating is shown in Figure 21. When the voltage was 

raised to 72 volts (Fig. 22) interlaminar bonding appeared to improve and the 

Size of the large voids was reduced. Also, less evidence of ceramic bond 

coat voids was evidenced. The substrate temperature also increase by 52°C. 

Further increasing the voltage to 78 volts tended to produce laminar separation in the coating. The deposit rate decreased by 29 percent and substrate 

temperatures were 5°C lower than at 72 volts, due to the lower density of the 

coating. Voltage could not be increased beyond this point due to equipment 

1 imitat ions. 

The effect of voltage on the bond coat was established using a 61 cm angle 

extension gun. This unit was selected for these tests since it was capable of 

31 
Magnification: 1000X 

Figure 21. Thermal Barrier Coating With Plasma Arc Voltage of 66 VOlts 

Magnification: 1000X 

Figure 22. Thermal Barrier Coating With Plasma Arc Voltage of 72 Volts 

32 
coating the 6.6 cm base of a test thrust chamber. The parameters in Table 5 

were used as a baseline for these tests. 

Table 5 

Baseline, Angle Extension Spray Parameters 

Current 

Voltage 

Argon Flowrate 

Hydrogen Flowrate 

Standoff Distance 

Powder Port Type 

Carrier Gas Flowrate 

lJPOwder Feedrate 

Anode Model 

Traverse Rate 

Deposit Rate 

500 amperes 

45 volts 

1. 7 m3/hr 

0.8 m3/hr 

3.2 em 

#2 

0.28 m3/hr 

0.098 gm/sec 

#713 

2.5 em/sec 

0.0025 mm/traverse 

Tabl e 6 

Effect of Voltage on Bond Coat Structure 

Hydrogen 

Flowrate Deposit Rate Depos it Wi dth 

Voltage (m3 /hr) (mm/cycle) (mm) 

45 0.085 0.0025 not definable 

50 0.20 0.0051 3.2 

55 0.33 0.010 4.8 

60 0.51 0.020 6.4 

65 0.65 0.028 6.4 

70 0.76 0.033 7.9 

75 0.82 0.038 7.9 

Table 6 summarizes the measured results of these tests. The relationship between arc voltage and hydrogen flow proved to be 1 i near. The wi dth over 

which the deposit occurred also increased linearly with increasing hydrogen. 

The width measurement was made directly using a machinists scale. Thus, the 

accuracy of this measurement is 1 imited. The increase over a 5 volt range 

Was less than the measured accuracy (+0.5 mm) making this relationship less 

definitive. The deposit rate increasea rapidly over the 55 to 65 volt range 

indicating that this range is especially sensitive to voltage variations. 

33 
The effect of voltage variations on bond coat deposition is shown in Figure 

23 to Figure 29. In each case the plasma spray gun was traversed across the 

specimen twelve times. 

In Figure 23 the arc voltage was 45 volts. There is no visual evidence of 

bond coat deposition. Oxidation of the substrate is apparent. At 50 volts, 

coating deposition has begun. In Figure 24 small local areas evidence individual particle deposition. The coating is sporadic and again the copper 

substrate is oxidized. The effect of increasing the voltage to 55 volts is 

evident in Figure 25, where significant bond coat deposition has occurred. 

The oxidation of the substrate observed in the two previous specimens is 

absent and the bond-substrate interface is relatively clean. (In viewing these photomicrographs the main deposit is in the central third. Material at 

each end is overspray.) 

At 60 volts the coating thickness is nearly twice that of the coating deposited at 55 volts. The substrate is just beginning to evidence overheating at 

the edges. The effect of the 45 degree plasma jet impingement angle is also 

evident. The coating buildup slopes to the left with elongated voids that 

are created by the masking effects of previously deposited particles. With 

the voltage increased to 65 volts, the coating substrate bond shows evidence 

of significant oxidation. Wide variations in coating density were also 

observed, ranging from large porous areas to high density regions. 

Magnification: 100X 

Figure 23. Substrate After 12 Traverses at 45 Volts 

34 
Magnification: 100X 

Figure 24. Substrate After 12 Traverses at 50 Volts 

Magnification: 100X 

Figure 25. NiCrA1Y Deposit After 12 Traverses at 55 Volts 

35 
Magnification: 100X 

Figure 26. NiCrA1Y Deposit After 12 Traverses at 60 Volts 

Magnification: 100X 

Figure 27. NiCrA1Y Deposit After 12 Traverses at 65 Volts 

36 
Magnification: 100X 

Figure 28. NiCrA1Y Deposit After 12 Traverses at 70 Volts 

Magni fi cat ion: lOOX 

Figure 29. NiCrA1Y Deposit Ater 12 Traverses at 75 Volts 

37 
At 70 volts, Figure 28, the coating exhibited some densification with large 

voids. The substrate-coating interface is severely oxidized indicating poor adherence. The slope of the voids and deposit to the right is evidence of 

the 45 degree spray angle. 

The deposit at 75 volts (Fig. 29) is similar to that at 70 volts. Void size 

has increased but the coating is more uniform, particularly at the bond line. 

3.1.4 Effect of Arc Current on Structure 

Varying the arc current independently within normal operating parameters had 

little effect on coating structure. With the equipment configuration used 

for this test, operating current is limited to the range of 350-500 amperes. 

The ceramic baseline coating sprayed at 400 amperes was shown in Section 

3.1.1. The current was reduced to 350 amperes for the specimen shown in 

Figure 30. Little, if any, differentiation can be made between it and the 

baseline coating that is attributable to the lower current. A coating sprayed 

at 475 amperes (Fig. 31) showed little change in structure. A slight tendency 

of the voids towards a spherical shape was observed but this could not be 

established on a reproducible basis. Additional tests with higher currents 

Magnification: 100X 

Figure 30. Thermal Barrier Coating With Plasma Arc 

Current of 350 Amperes 

38 
Magnification: 1000X 

Figure 31. Thermal Barrier Coating With Plasma Arc Current 

of 475 Amperes 

were performed as described separately in Section 3.1.6. These entailed the 

use of a different anode to allow operation at 80 kW where other modifications 

to the operating parameters masked the effect of arc current alone. 

3.1.5 Anode Selection Effects 

The water cooled copper anode used in a plasma spray gun also functions as a 

nozzle for controlling and ducting the plasma gas flow. Design of this 

component is one of the primary factors in determining the exit gas velocity and maximum arc current that can be used. The data in Table 7 was supplied 

by Metco Inc. as a guide for evaluating the performance characteristics of 

the various anodes available for the 7M system. 

Trial coatings were sprayed with two of these anodes using the parameters given in Table 8. Manufacturer recommended parameters were used in each 

case. A baseline coating deposited with the GH anode is shown in Figure 32. 

This structure is typical of the thin coatings developed during this study. 

The type 700 anode is similar to the GH but is operated at higher power levels. 

A bond coat deposited using this anode is shown in Figure 33. Coating thickness developed more rapidly using the 700 anode than with t~e type GH anode. 

The deposit also evidenced greater porosity and oxidation than the baseline 

(GH anode) coating. No obvious benefit in coating quality was observed with 

39 
Table 7 

Plasma Spray Anode Performance Characteristics 

Maximum Electrical 

Parameters Pl asma Nozzle 

Gas Exit Li fe 

Anode System kW Amps Volts Velocity (hrs) 

GH A/H2 37 500 75 1829 25 

700 A/H2 42 600 70 1768 20 

703 A/H2 45 600 75 2621 15 

704 A 58 1000 58 2835 30 

Tabl e 8 

Effect of Anode Selection 

Current Voltage Power Standoff Thickness 

Coating Anode (amperes) (volts) kW (cm) (mm) 

Bond GH 500 65 32 10.0 0.025 

Ceramic GH 450 50 22 7.6 0.025 

Bond GH 400 50 20 10.0 0.051 

Cerami c 704 1000 54 54 7.6 0.025 

Bond 704 1000 52 52 10.0 0.130 

_. 

this anode. Operating the 700 anode at lower power levels produced coatings 

similar to those obtained with the type GH anode. 

The type 703 anode operated at 45 kW produced excessive substrate oxidation 

as shown in Figure 34. Severe oxidation pits are visible in the copper substrate, and unmelted particles are evident where particle melting and 

vacuum fusion of the NiCrA1Y produced entrapped voids. The laminar structure 

normally seen in these structures is not evident. 

Two examples of coatings applied with the type 704 anode are shown in Figures 

35 and 36. A zirconia coating applied over a standard bond coat is shown in 

Figure 35, and Figure 36 shows just a NiCrA1Y bond coat applied with the type 

704 anode. In both tests the deposit rate was low, and thin dense coatings 

resul ted. Severe oxi dat i on occurred at the bond coat-substrate interface, 

indicating overheating. The bond coat applied under these conditions was 

also rich in oxides. No indications of coating improvment were obtained with 

these special anodes and it was decided to continue tests with the standard 

GH anode. 

40 
Magnification: 500X 

Figure 32. Baseline Coating Applied with a GH Anode 

Magnification: 500X 

Figure 33. Bond Coat Deposited at 42 kW With A Type 700 Anode 

41 
Magnification: 500X 

Figure 34. Bond Coat Applied at 45 kw With a Type 703 Anode 

- Bond 

4 Copper 

Magnification: 500X 

Figure 35. Zirconia Applied to the Baseline Bond Coat With a 

704 Anode at 58 kW 

Figure 36. Bond Coat Applied With a Type 704 Anode at 58 kw 

42 
3.1.6 Effect of Traverse Rate 

The baseline traverse rate used in these tests was 2.5 cm/sec. This was 

reduced incrementally to 1.2 cm/s. No visible change was found in the coating microstructure. However, as traverse rate was decreased, substrate temperature increased. At 1.2 cm/s this increase amounted to 12°C. This slight 

increase in substrate temperature did not visibly affect bond quality. 

3.1.7 Powder Feed Rate 

Zirconia feed rate was varied over the range of 0.04 to 0.09 gm/s. The basel ine value was 0.05 gm/sec. Reproducible feed rates lower than 0.04 gm/s could not be achieved. The reduction in feed rate below baseline (Fig. 37) 

did not affect the coating microstructure. As feed rates were increased, 

laminar voids developed as shown in Figures 38 and 39. At feed rates higher 

than 0.09 gm/s deposit efficiency decreased, as evidenced by visible, unmelted 

particles bouncing off the fixture and substrate. Generally, it was found 

that the lower the feed rate the more uniform the coating, with a correspond- ing reduction in unmelted particles in the coating. 

Magnification: 1000X 

Figure 37. Thermal Barrier Coating With Powder Feed Rate 

of 0.04 gms/sec 

43 
Magnification: lOOOX 

Figure 38. Thermal Barrier Coating With Powder Feed RAte of 0.062 gms/sec 

Magnification: 1000X 

Figure 39. Thermal Barrier Coating With Powder Feed RAte of 0.09 gms/sec 

44 
3.2 PROPERTY MEASUREMENT 

Based on the laboratory data from the studies performed in Section 2, three 

sets of ceramics and one set of bond coat deposition parameters were selected 

for further study. The parameters are given in Table 9. Examples of coated 

specimens used for property measurement testing are shown in Appendix 3. 

Table 9 

Plasma Spray Parameters Selected for Coating Property Measurements ' 

Ceramic Ceramic Ceramic 

Coating Coating Coating :l 

#1 #2 #3 Coat -

Current (amps) 400 500 600 500 

Vo ltage (volts) 50 65 70 65 

Power (kW) 20 32 42 32 

Argon (m3/hr) 2.97 2.97 4.25 5.66 

Hydrogen (m3/hr) 0.085 0.14 0.20 0.11 

Stand-Off (cm) 6.4 6.4 6.4 

Plasma Velocity (m/s) 25.4 30.5 43.7 32.5 

Powder Port No.2 No. 2 No. 3 No. 2 

Cooling Air (kPa) 551 551 551 551 

Nozzl e GH GH 703 GH 

,---._. 

The same bond coat parameters were used in all tests to apply a nominal 0.025 

mm coating to the copper substrate. Substrate temperatures were held within 

the range of 97 to 120°C during coating application. Higher substrate temperatures caused excessive oxide formation at the bond coat-copper interface. 

The fi rst system (Fi g. 40) used the standard bond coat and the zi rconi a-8% 

yttria was applied at 20 kW. The 32 kW coating is shown is Figure 41 and the 

40 kW coating is shown in Figure 42. 

As the power level increased the coatings exhibited a reduction in pore size 

and an increase in porosity. It was found that the coatings produced at 20 

and 40 kW could be dupl icated readily. However, those produced at 32 kW 

varied in density and other properties as discussed in Section 3.5. Generally, 

as anode wear increased the coating tended to become less dense. This effect 

was only evident on the 32 kW coatings and no quantitative relationship could 

be established. Anode usage was limited to four hours to minimize anode wear 

effects but they could not be eliminated from the 32 kW coatings. Because of 

this variation, the property data reported in this section is limited to values obtained from specimens produced with new anodes unless otherwise noted. 

45 
Bond coat 

Magnification: 100X 

Figure 40. Thermal Barrier Coating Applied at 20 kW (System 1) 

Zr02 

Bond Coat 

Magnification: 100X 

Figure 41. Thermal Barrier Coating Applied at 32 kw 

46 
Magnification: 100X 

Figure 42. Thermal Barrier Coating Applied at 40 kw 

3.2.1 Density 

Coating density was determined by three methods: 

Visual inspection of a 100X photomicrograph 

Photometric evaluation of SOX photomicrographs 

Weighing and measuring a sample of the coating 

The following average values were obtained: 

~sual Photometric Weight 

Coat; ng System 1 

Coating System 2 

Coating System 3 

% Theory) 

85 

92 

95 

47 

(% Theory) (glee) 

80 I 4.8 

90 5.2 

95 5.3 
3.2.2 Thickness Control 

Because of the extreme thinness of the coatings required in this program (0.013 to 0.10 mm) control of the application parameters was considered an 

important factor. Conventional coatings vary by +0.02 mm or more from the 

nominal thickness. For the coatings evaluated fn this study, this could 

amount to +12% of the thickest coatings while producing completely uncoated 

areas withlthe thinner coatings. To assure a precise and uniform coating, it 

was found that low powder feedrates, in the range of 0.05 to 0.10 gm/s, were 

required. Under these conditions typical build-up rates were 0.0025 to 0.008 

mm per traverse. The low powder feed al so provided a more uniform ceramic 

deposit, free of entrapped particles. Under these conditions coatings 0.025 

mm thick could be produced to an accuracy of +0.005 mm. 

3.2.3 Thermal Diffusivity and Conductivity 

Thermal conductivity and diffusivity of the coatings were measured as descri bed previously. The val ues obtai ned are reported in Tabl e 10 and in 

Figures 43 and 44. 

The coatings applied at 20 and 40 kW show an increase in thermal conductivity 

with increased gun power. The 32 kW coating is inconsistent in that it does 

not lie between the other two values as would be anticipated. In other tests 

the 32 kW coating responded as if it were less dense than the 20 kW coating. 

The source of thi s erratic performance at 32 kW has not been i dent ifi ed. 

Apparently at this power level the plasma flame is sensitive to minor variations in anode wear and gas flow, at levels below the sensitiity of the 

process controls. Because of this fluctuation, the 20 and 40 kW coatings 

were selected for coating the experimental thrust chambers. 

3.2.4 Surface Finish 

Surface finish did not appear to be a function of plasma spray gun parameters. When surface finish values of the as-sprayed coatings were checked, arithmetic 

average val ues of 5-10 m; crometers were obtai ned • Generally these values 

appeared to be the result of initial particle size rather than power levels, 

once a uniform coating was obtained. 

3.2.5 Crystal Structure 

X-ray diffraction measurements of the coatings gave nominal values of 94-98% 

cubic and tetragonal with the remainder being monoclinic. The initial powder 

was reported as 10% monoclinic with the balance cubic. This transformation 

of the monoclinic zirconia to cubic/tetragonal was similar in all cases. 

There appeared to be less monoclinic zirconia (about 3%) in the 20 kW coatings 

48 
~ 

~ 

Sample 

20 kW 

32 kW 

40 kW 

Temperature 

( OC) 

23 

100 

200 

300 

400 

500 

23 

100 

200 

300 

400 

500 

23 

100 

200 

300 

400 

500 

Table 10 

Thermal Diffusivity and Conductivity of Zr02Y203 Layers 

Thickness Density Specific Heat Half-Time Diffusivity Conductivity Conductivity 

(cm) (gm cm-3 ) (W s gm-1K-1) (sec) (cm2 sec-1 ) (W cm- 1K-l) (Btu in hr- 1ft-2F-l) 

0.0254 4.8 0.457 0.0775 0.00298 0.00654 4.53 

0.0254 4.8 0.491 0.0891 0.00250 0.00590 4.09 

0.0254 4.8 0.522 0.0924 0.00238 0.00596 4.13 

0.0254 4.8 0.541 0.0965 0.00229 0.00593 4.11 

0.0254 4.8 0.552 0.0101 0.00219 0.00581 4.03 

0.0254 4.8 0.567 0.1009 0.00218 0.00591 4.10 

0.0178 5.2 0.457 0.0337 0.00413 0.00981 6.80 

0.0178 5.2 0.491 0.0338 0.00397 0.01010 7.00 

0.0178 5.2 0.522 0.0354 0.00372 0.01010 7.00 

0.0178 5.2 0.541 0.0365 0.00363 0.01020 7.07 

0.0178 5.2 0.552 0.0380 0.00345 0.00991 6.81 

0.0178 5.2 0.567 0.0383 0.00342 0.01910 7.00 

0.0178 5.2 0.457 0.0339 0.00379 0.00901 6.25 

0.0178 5.2 0.491 0.0353 0.00356 0.00909 6.30 

0.0178 5.2 0.522 0.0384 0.00321 0.00871 6.04 

0.0178 5.2 0.541 0.0397 0.00311 0.00875 6.07 

0.0178 5.2 0.552 0.0408 0.00302 0.00866 6.00 

0.0178 5.2 0.557 0.0421 0.00292 0.00860 5.96 

- .. -.. ~-.. - --
0 

" J! 

E 

~ 

>- I-

;; 

1ii 

:0 

u. 

u. 

15 

...J 

« 

::; 

a: 

w 

:r: 

ITEMPERATURE ('F) 

200 400 600 800 1000 0.008 r------.------,-C------..:...------=.;..:....-----...:..:;.-=---~ 

0.006 

0.004 

0.002 

o 20kW 

6. 32 kW 

o 40kW 

A ____ ~ 

~-----~-----~-------L.> ------l\. 

o :I-----r::::r---_: 

0.000 L..... ____ -'--____ -'-____ -L ____ --L. ____ --'-____ ---l 

100 200 300 400 500 

TEMPERATURE ('C) 

Figure 43. Thermal Diffusivity of Zr02Y203 Layers 

TEMPERATURE ('F) 

0,016 r-___ -=:;2~::..O ____ -=;40::...0 ___ -"i6~r-0 ___ ~800:'C----.!C:l0rOO~___. 

0.012 

o 20kW 

1\ 32kW 

o 4QkW 

0.010 l:r ___ -6-- ---------"-- - - - -rr- - --_-6 

o~--~Q~----~o~----~o------~o __ -----o 

0.008 

0.... 0.006 ------~Or_----~o~----~o~------ao-------O 

0.004 

0.002 "-___ -'--___ -'-___ .....L ___ -L ___ .--L ___ ---' 

10 

o ~ B B a B B 

TEMPERATURE ('c) 

[: 

oj" 

~ 

~ '" 

~ 

:; 

§ 

" 0 

z 

0 

<> 

... 

.. 

~ UJ 

1: 

.... 

Figure 44. Thermal Conductivity of Zr02Y203 Layers 

600 

than in the 40 kW coating, but this variation was within the error limits of 

the analysis and it could not verified. 

50 
3.2.6 Calorimeter Tests 

The 1.27 cm calorimeter probe is shown in Figure 45 and the test surface 

installed in a guard is shown in Figure 46. Some loss of insulation between 

the probe and guard is visible in the second figure. When this loss occurred 

the change in outlet temperature was less than 0.5°C and the results remained 

within experimental accuracy. 

In operation the plasma flame was directed at the 1.27 cm diameter face of 

the calorimeter. Water flow and the T between the inlet and outlet water 

temperatures were recorded for calculating the total heat input to the cold 

face calorimeter. Both coated and uncoated calorimeters were tested. 

The appearance of a zirconia coated calorimeter is shown in Figure 47. The 

coating was applied to a thickness of 0.10 mm at 40 kW. In this test, the 

calorimeter was exposed to a 32 kW plasma flame. The relationship between 

power level and gun distance to heat flux is shown in Figure 48. Input to 

the calorimeter was measured for three sets of gun parameters as shown, and 

plotted as a function of distance. The gun parameters were given previously 

in Table 9. 

Figure 45. Water Cooled Calorimeter Probe 

51 
Figure 47. 

Figure 46. Calorimeter Face in Guard Shield 

Calorimeter Face With 40 kW Zr02Y203 Coating 0.10 mm 

Thick Exposed to a Heat Flux of 9.15 watts/mm2 

Heating appears to be more sensitivite to voltage (hydrogen flow) than to 

current. The current increased by 100 amperes for each increase in power. 

The 15 vol t increase between 20 and 32 kW shows a far greater ; ncrease in 

heat input than the 5 volt increase between 32 and 40 kW. 

52 
12r----r,-----------------------------------------~ 

10 

42kW 

- ______ :.:::::::: 32kW 

22 kW 

1.0 2.0 3.0 4.0 5.0 

DISTANCE (In.) 

Figure 48. Heat Flux From a Plasma Torch as a Function of 

Distance at Selected Power Levels 

The effect of the zi rconi a coat i ngs on heat input to the calorimeter is 

reported in Figure 49 through 51. The 20 kW coating provided the greatest 

reduction in heat input to the calorimeter. The coating remained intact 

until the heat flux through the coating reached 8.2 watts/mm2 at which pOint 

surface melting occurred. When the 32 kW coating was tested failure occurred 

at a heat input of 9.0 watts/mm2 through spallation of the coating. The 

third coating, 40 kW zirconia, survived until surface melting occurred at an 

input of approximately 16.3 watts/mm2 or twi ce that of the 20 kW coati ng. 

Failure of this coating also occurred as a result of melting. Some cracking normal to the surface occurred that may al so have contributed to the low 

insulative value obtained with this coating. 

Generally, the tests confirmed the thermal conductivity values previously 

measured and showed that the coatings could be retained up to the point of 

surface melting. The poor performance of the 32 kW coating supported the 

selection of the 20 and 40 kW for the coating of test thrust chambers. 

3.2.7 Residual Stress and Stress Free Temperature 

Several flat copper strip specimens were coated for residual stress and stress 

free temperature measurement. A copper thickness of 0.8128 mm and a bond coat 

thickness of 0.0508 mm were used for these tests. Density, residual stress 

53 
18.0 

16.0 

14.0 

12.0 

W 

E 10.0 ~ x 

:3 

u. 

f- 8.0 

~ J: 

6.0 

4.0 

2.0 

FAILED BY 

MELTING 

\ 

\ 

\ 

, 

~ 

COATED-- " 

............ .......... ....... ....... ....... ..... -.......... _-

10 

DISTANCE (em) 

12 

------------

14 16 18 20 

Figure 49. Heat Flux Transmitted to a Copper Calorimeter at 32 kW 

Through a 0.1 mm 20 kW Zirconia Coating 

18.0 

16.0 

14.0 

12.0 

'E 

E 10.0 

~ x 

:> 

..J 

u. 8.0 f-

'" w 

J: 

6.0 

4.0 

2.0 

------- L-____ ~ ____ -L ____ _L. ____ ~ ______ L_ _____ ~ ____ ~ ____ ~ ____ ~ ____ ~ 

6 10 

DISTANCE (em) 

12 14 16 18 20 

Figure 50. Heat Flux Transmitted to a Copper Caloriometer at 32 kW 

Through a 0.10 mm, 32 kW Zirconia Coating 

54 
18.0 

16.0 

14.0 

12.0 

N 

E 

E 10.0 

~ x 

:::> 

-' "- 8.0 I-« 

w 

:x: 

6.0 

4.0 

2.0 

FAIL BY 

MELTING 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ 

\ , 

"- COATED-- " , , 

........ ........ ...... '--- 10 

DISTANCE (em) 

12 

---- -:-::-:::-::-=--~--------

14 16 18 20 

Figure 51. Heat Flux Transmitted to a Copper Calorimeter at 32 kW 

Through a 0.10 mm, 40 kW Zirconia Coating 

and stress free temperature (SFT) were measured as described in the previous section. The results are shown in Table 1. The spray parameters for these 

coatings are given in Table 9. 

These data show an increase in density and a rather constant residual stress 

and SFT with an increase in gun power from 20 kW to 32 kW. A substantial 

increase in residual stress and SFT was observed when the gun power was 

increased further, to 42 kW, although the desnsity remained constant. Increas i ng the coat i ng thi ckness from 0.0965 mm to 0.254 mm a 1 so caused a 

significant increase in residual stress and SFT. No density measurement was 

made on this latter coating. 

It is interesting to note that these data show the coating to incur a compressive residual stress during plasma spraying. This could effectively 

weaken the coating in service since compressive spallation is the predominant 

coating failure mode. However, the data generated thus far is rather limited 

and does not warrant conclusive interpretation. 

The coatings produced in this study were strength tested in their as-sprayed 

condition (e.g., no annealing). Thus, the strength data generated reflects 

any effect of residual stress on coating strength. Reduction of residual 

stresses may present a practical means of increasing coating strength. This 

concept, although not pursued on this program, may be worthy of future 

coating development studies. 

55 
Table 11 

Residual Stress and Stress Free Temperature Results From Coated 

Flat Specimens (0.8128 mm copper and 0.0508 mm NiCrA1Y bond) 

Deposit 

Rate Weight Coating (mil s/ Gai n 0 a1 L S TSF Density 

Type* pass) (gms) Imml ImmJ lmm) (MPa) (C) (%) 

I 0.1 0.3990 0.1778 0.09652 50.8 -7.2 58 88.6 

0.1 0.3778 0.3048 0.09652 50.8 -12.3 84 84.0 

II 0.6 0.4341 0.2540 0.09652 50.8 -10.3 74 96.5 

0.6 0.4315 0.2794 0.09652 50.8 -11.3 79 96.0 

I 

- I -- I 5.4864 0.02540 55.12 

;1447 \ --

III I 0.4 

I 0.5~~.9812 0.1016 LO.8 -75.6 ~~ 93.~ 

3.3 ANALYTICAL MODELING 

3.3.1 Thrust Chamber Model Development 

Initial finite element performance analyses of the thrust chamber coating 

system were subcontracted to Control Data Corporation. This initial work 

included the model construction, preliminary test runs, and several coating 

analyses. All later finite element analyses were performed at Solar. 

The original thrust chamber coating system model was 50.8 mm long and was 

divided into six elements axially. This resulted in an element aspect ratio 

(length/width) of 667 for the bond coat elements, which are the most extreme. 

The model length was later changed to 0.229 mm, which reduced the bond coat 

element aspect ratio to 3. This change caused no difference in the thermal 

analyses and less than a three percent variation in the calculated stresses. 

The shorter model was used for all of the final coating analyses. 

The finite element model of the thrust chamber coating system was used to 

evaluate typical coating performance under selected operating conditions. 

Material properties from the literature were used in these analyses (Table 

1). The selected operating conditions included a hot ceramic surface temperature of 1650°C (the maximum desirable ceramic operating temperature), cold 

copper surface temperatures of -240, -18 and 204°C, and a specified heat flux 

of 18 W/mm2. These temperature selections were explained earlier but this 

heat flux was selected inadvertently due to an error in the original analyses 

performed by CDC. The heat flux was originally intended to be 49 W/mm2 but 

the units were miSinterpreted to be W rather than W/mm2. Nonetheless, these 

56 
analyses were useful in developing the coating model and observing the general 

behavior of the coating system under potential operating conditions. 

These initial coating systems analyses used a constant ceramic coating thickness of 0.10 mm. Steady state analyses were performed to evaluate the temperature di stri buti ons, heat fl uxes and stress profi 1 es through the coatings. 

One series of coating system analyses used a constant hot surface temperature 

of 1650°C and three cold side temperatures: -240, -18 and 204°C. The resulting temperature gradients are shown in Figure 52. This figure shows that 

the bul k of the temperature drop occurs throug h the thi n cerami c 1 ayer, subjecting the copper to a relatively low temperature gradient. The temperature gradients through the coating and the copper do not appear to be signi- ficantly influenced by the cold side temperature, although the overall temperature difference was increased by 31% from the smallest value to the largest. 

The stress profiles resulting from the temperature gradients shown in Figure 

52 are presented next. The hoop (or circumferential) stresses are equal to 

the axial stresses (due to the radial temperature gradient) and are shown in 

Figure 53. The radial stress profiles are shown in Figure 54. Step discontinuities in these stress profiles occur at the interfaces of different 

materials in the coating system. These discontinuities result from the 

changes in elastic moduli between materials, since the analyses assume continuous strain across the interfaces. 

00 c:: 

1750 -z w ~ :EO 0- <C m 0-

c:: 0 1500 to- w 0 0 

-

0 1250 t-

-

° 

Q) 

1000 ... ~ :::I 

-f! 

Q) 750 to. 

E 

Q) 

t- 500 -

250 - ~ TO _ 204°C 

0 - ~ TO =-18°C 

-250 \. I I 

TO = -240°C 

~ 

3.28 3.30 3.32 3.34 3.36 3.38 3.40 

Radius (em) 

Figure 52. Temperature Radient for 1650°C Wall Temperature 

57 
100 

o 

o 

-::!E C 

<t z 

a: 0 

W en 

o 

a:: 

w 

0.. 

0.. 

o 

o 

-<'CS 

0.. 

:E 

-I/) - 100 

I/) 

! 

CJ) 

--200 

-300 

-400~----~----~----~----~--~~--~ 

3.28 3.30 3.32 3.34 3.36 3.38 3.40 

Radius (em) 

Figure 53. Axial Stress Gradient for 1650°F Wall Temperature 

200 

o 

-<'CS 

D-

~ -200 

I/) 

I/) 

~ 

-

o 

:E 

« 

a: 

C 

Z 

o 

en 

a: 

W 

0.. 

DO 

o 

CJ) _ 400 W 

() 

TO = 204°C 

TO = -18°C 

TO = - 240°C 

-600 

3.28 3.30 3.32 3.34 3.36 3.38 

Radius (em) 

Figure 54. Thrust Chamber Radial Stress Gradient for 

1650°C Inner Wall Temperature 

58 
The hoop and axial stresses in the coating are shown to be compressive; the 

maximum values occurring at the coating surface with a linear reduction 

through the thickness. The stresses in the bond coat are again compressive, 

whi 1 e the stresses in the copper change from compress i ve to tensi 1 e. The 

coating stresses are of significant magnitude, with the maximum stresses on 

the order of -240 to -400 MPa. Stresses in the copper wall range from -31 to 

+110 MPa. The maximum tensile stress in the copper falls between the yield strengths of annealed and hardened oxygen-free copper. 

The radial stresses through the coating system are much less severe than the 

hoop and axial stresses. Thus, the coating system is essentially in a biaxial 

stress state during operation. Radial stresses in the ceramic are approximately 0.5 MPa maximum, while the radial stresses in the copper are on the 

order of -0.9 to -1.6 MPa, maximum. 

The cold surface temperature appears to have a significant effect on the hoop and axial stresses in the coating. A cold temperature change from 204 to 

-240°C (a 31% increase in overall temperature differential) results in a 60% 

increase in the maximum compressive coating stress. The effect on the hoop 

and axial stresses in the copper is less pronounced, with approximately a 45% 

increase in stress observed for the same change in cold side temperature. 

Cold side temperature also had an effect on the radial stresses, but the 

stress magnitude remained relatively insignificant. 

A second set of coating system analyses used a constant radial heat flux of 

18 W/mm2 through the coating and the same three cold surface temperatures as 

before: -240, -18 and 204°C. The resulting temperature gradients are shown 

in Figure 55. The heat flux specified for these analyses was similar to the 

resulting heat fluxes in the previous analyses. In fact, the two cases having 

a 204°C cold side temperature are nearly identical. 

As expected, the hot surface temperature decreases with decreasing cold 

surface temperature, since the heat flux is constant. It is interesting to 

note that the decrease in hot surface temperature (333°C) is approximately 25% less than the decrease in cold surface temperature (444°C), over the 

range eval uated. Thi sis attri buted to temperature affects on the materi al 

properties. As in the previous analyses, most of the temperature drop occurs 

through the ceramic coating, subjecting the bond and substrate to to relatively small thermal gradients. 

The hoop and axial stress profiles resulting from the temperature gradients in Figure 55 are shown in Figure 56. These stress profiles are quite similar 

to those of the previous analyses, again showing the ceramic coating under 

very high compressive stress. The copper is subjected to slight compressive 

stresses at the bond interface and moderate tensi 1 e stresses at the outer 

surface. 

These first two series of analyses were conducted to exercise the thrust 

chamber coating system model, and to evaluate the general behavior of the 

coati ng system under typi ca 1 rocket engi ne operati ng conditions. The model 

was found to operate acceptably and the coating system behavior was found to 

be reasonable. 

59 
c a: 

z w 

1750 ~ , 0 D.. 

D.. al 0 

1500 - 0 

-

0 1250 ~ ° 

-CD 

-1000 I- :::s 

-co -

CD 750 le. 

E 

CD 500 ~ 

f- TO = 204°C 

0 ,," 250 I- -:E l\. TO =-18°c « 0 I"- a: 

w TO =-240°c 

-250 t- 0 1 .1 I 

3.28 3.30 3.32 3.34 3.36 3.38 3.40 

Radius (em) 

Figure 55. Thrust Chamber Temperature Gradient for 1650°C 

Inner Wall Temperature 

-CIS 

Q. 

E 

-UI 

In 

Q) 

... 

en 

-

~ c a: 

:E z w 

0 Q. 

« Q. 

a: £Xl 0 W 0 0 

200°C 8l O°C \ 

" 120 

-200°C 

3.28 3.30 3.32 3.34 3.36 3.38 3.40 

Radius (em) 

Figure 56. Thrust Chamber Axial Stress Gradient for 1650°C 

Inner Wall Temperature 

60 
This model was then used to evaluate the effect of thickness on the predicted 

coati ng temperatures and stresses in the thrust chamber envi ronment. These 

analyses again used literature values for material properties. The purpose 

of these analyses was to observe the general effect of coating thickness on 

coating stress and temperature, and to establish a probable range of the 

coating thickness required for rocket thrust chamber application. 

Figure 57 shows both surface temperature (left scale) and hoop (or axial) 

stress (right scale) in the coating for a constant heat flux (20 w/mm2). 

This heat flux is at the lower end of a typical thrust chamber operating 

range. Both coating temperature and stress increase linearly with coating 

thi ckness. 

This analysis indicates that a coating thickness of less than 0.13 mm may be 

required to maintain an acceptable coating temperature (2000°C) at this heat 

flux condition. The compressive coating stresses in this thickness range 

woul d be 1 ess than 400 MPa. The s i gnifi cance of thi s was not known at the 

time of these analyses. 

Figure 58 shows the effect of coating thickness on heat flux for a fixed 

temperature gradi ent. The hot surface temperature (1650°C) used for thi s 

analysis is a desirable upper limit for the coating. The cold temperature 

(-18°C) is a reasonable anticipated operating value. This curve shows the 

heat flux is inversely proportional to coating thickness. At very low thicknesses, the heat flux is very sensitive to coating thickness. Again, the 

-() 

° 

-Q) 

-. 

::;, 

-a:s 

-. 

Q) 

Q. 

E 

Q) 

IQ) 

(,) 

a:s 

--. 

::;, 

en 

4000 r-----~------...,.....--___. 300 

-a:s 

400 Il. 

E 

3000 450 

Tcold =-18°C 500 

Q = 20 w/mm2 

2000 .... 550 

600 

1000 650 

0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 

Ceramic Coating Thickness (mm) 

-(I) 

(I) 

Q) 

-. 

en -

Q. 

0 

0 

l: 

'tJ 

c 

a:s 

-

.~ 

>< « 

Figure 57. Relationship of Coating Surface Temperature and Stress 

to Coating Thickness to Constant Heat Flux 

61 
-

C\I 

... 

... 

E 

E 

-3 

-

>< 

:::J 

LL 

-<0 

Cl) 

J: 

30.00 

28.00 

26.00 

24.00 

22.00 

20.00 

18.00 

16.00 

14.00 

12.00 

10.00 

0.06 0.08 0.10 0.12 0.14 

THot = 16500 C 

TCoid = - 20°C 

0.16 0.18 0.20 0.22 

Coating Thickness (mm) 

Figure 58. Initial Predicted Heat Flux Versus Coating Thickness at Constant ~T 

range of heat fl uxes covered by thi s graph is in the moderate operating 

regime. This analysis indicates that coating thicknesses of less than 0.075 

mm may be required to accommodate the high heat flux conditions without 

excessive temperature. The maximum compressive coating stress for this case 

is constant with varying coating thickness and is 350 MPa. 

3.3.2 Coating Structure Model Development 

Finite Element Model of Porous Coating 

A variety of coating structures were evaluated using the finite element model 

shown in Figure 5. These structures incl uded a variety of pore dimension 

ratios (Dr), total porosities (Pt), and pore sizes (Pi). An attempt was made 

to establish a relationship between these three coating parameters (Dr, Pt, Pi) 

and the effective thermal conductivity of the coating system (k). 

An evaluation of the significance of radiation to the overall heat flow 

through a TBC was conducted. This was accomplished by comparing the heat 

flux through one coating model, with and without radiation across the pores. A negligible effect was observed. Thus, radiation was omitted from further 

analyses, to reduce the required computer time. 

Figure 59 shows plots of effective thermal conductivity vs. pore dimension 

ratio for various combinations of total porosity and individual pore size. 

62 
>- 100 ls;: 

j:: 90 

0 

::::) 

c_ 80 

z:E 70 0::::) 

0:E 60 ...J - <e>< 

~<e 50 a::E 

WLL 40 :7:0 

I- 0 

w~ 30 

>-

I- 20 

° w 10 LL 

LL 

W 0 

0 

~~ _________________ F 

------E 

::---------0 

~--~~~--------C 

B 

------A 

10 15 20 

DIMENSION RATIO (Dr) 

A 

B 

C 

o 

E 

F 

25 

Pt 

26.70 

13.35 

26.70 

13.35 

6.675 

13.35 

30 

Pi 

0.74 

1.48 

1.48 

0.74 

0.74 

0.37 

Figure 59. Influence of Microstructure on Effective Thermal Conductivity 

Inspection of these curves shows a consistent trend towards higher thermal 

conductivity with decreasing dimension ratio (Dr)' This trend is most pronounced for small values of Dr and tends to level off at large values of Dr' 

Also evident from this figure are inconsistencies in the effects of both 

porosity and pore size on effective conductivity. For example, curves A, 0 

and E show a consistent pattern of increasing conductivity with decreasing 

porosity. (This trend is intuitively understandable.) However, curves B 

and C depict the opposite effect; that is, a decreasing conductivity with 

decreasing porosity. Similarly, curves B, 0 and F indicate an increase in 

conductivity with decreasing pore size, while curves A and C show the opposite 

effect. 

Numerical regression analyses were performed to establish mathematical expressions for effective thermal conductivity in terms of one and two coating 

parameters. Due to the inconsistencies just described, it was not possible 

at this point to establish expressions containing all three coating parameters. It is believed that different coating parameters must be established, 

to replace or compliment the three parameters initially selected, in order 

to adequately and uniquely characterize a coating structure. Limited regression analyses were performed however, to illustrate the approach and methodology employed to establish an analytical thermal model of a porous coating 

structure. 

All of the curves in Fi gure 59 appear to have an approximately exponenti al 

form. One-variable regression analyses were performed to establish mathematical expressions for curves A, 0 and E. Curves A and E were found to best 

fit an exponential relationship, while curve 0 was found to best fit an 

63 
inverse Dr (Dr being the independent variable) relationship. Curve D can 

also be expressed in exponential form with slightly less accuracy. These 

expressions and their respective correlation coefficients (r) are presented 

here: 

Curve A k = 56.94 - 13.82 in (Dr) (r = 0.996) (10) 

Curve E k = 94.13 13.26 in (Dr) (r = 1.000) (11) 

Curve ° k = 67.46 7.74 in (Dr) (r = 0.975) (12 ) 

k = (42.59 Dr + 59.35)/Dr (r = 0.999) (13 ) 

Two-variable regression analyses were performed on two sets of curves in 

Figure 59. Each set used dimension ratio as one variable. The first set 

showed a cons; stent decrease ; n conducti vity with ; ncreas; ng porosity and 

contained curves A, ° and E. The second set showed a consistent decrease in 

conductivity with increasing pore size, and contained curves B, D and F. The 

following expressions were established for k: 

k = 526.8 pt - 0•70 0r- 0•32 

for a pore size of Pi = 0.74 and 

k = 65.4 Pi

- 0•31 0r- 0•13 

for a porosity of Pt = 13.35; 

coefficient. 

(rm = 0.906) (14) 

(r m = 0.767) (15 ) 

rm is the multiple regression correlation 

Some of the analytical "data" generated using the finite element porous 

coating model were used to create the plot of effective thermal conductivity 

versus porosity shown in Fi gure 60. All of the curves presented show the 

same general trend of decreasing thermal conductivity with increasing porosity, except for curve G, which shows a slight rise in conductivity when the 

porosity is increased from 15 to 25 percent. It should be noted that curve G 

still 1 i es in the range of "scatter" of the other curves. 

Curiously, curve G also most closely matches the experimental data, plotted 

as curve H, for porosities 1 ess than 15 percent. The experimental data is 

derived from measured values of coating conductivities and porosities (reported earlier) and an assumed value for the thermal conductivity of the solid 

ceramic of 2.06 W/mC, which ;s one of the commonly reported literature values. 

It appears from Figure 60 that the finite element model of a porous coating 

generally predicts the same effect of porosity on thermal conductivity as 

observed experimentally. Although the experimental data falls outside of the 

analytical results, the shape of the curve is baSically consistent. The 

objective of the analytical model was to simulate the effects of coating 

structure parameters on thermal conductivity, and then to empirically adjust 

the model to accurately calculate the actual value of a coating's thermal 

conductivity. 

64 
> ls: 

j:: 

() 

:::> c_ 

CURVE Pi Dr 

A 0.74 20 

B 0.74 10 

C 0.74 7.5 

o 0.74 3.75 

E 1.48 7.5 

F 0.37 3.75 

G 1.48. 15 

z~ 

0:::> 

()~ 

.... - <2:>< 

~<2: 

a:~ 

w LL 40 

::I: 0 

-::-.....;;;;::==--........ -E H EXPERIMENTAL 

o I-~ 

w~ 

> 

I-

() 

W 

LL 

LL 

W 

C 

20 B 

A 

10 

O~ __ ~ ____ ~ ____ ~ ____ ~ ____ ~ ____ -L __________ ~ 

10 15 20 25 30 

POROSITY (%) 

Figure 60. Finite Element Analysis of Porosity Effect on Effective 

Thermal Conductivity of Coatings 

Due to the inconsistencies in the analytical model results described earlier, 

no attempt to empirically adjust the model was made. The results thus far 

support the contention that such a model can be empirically modified to produce accurate results. However, the basic model must first accurately represent all of the critical structural parameters of a porous coating. Therefore, additional work is required to establish more appropriate coating characterization parameters before a useful analytical model can be developed. 

Theoretical Thermal Conductivity Predictions for Porous Coatings 

Much work has been done to theoretically predict the properties of mixtures 

of materials (Refs. 6, 7). This work encompasses mixtures of all numbers and 

types of phases. All binary mixtures can be divided into three basic classes: 

(1) one-phase miscible mixtures, (2) two-phase systems with one continuous 

phase and one dispersed phase and (3) two-phase systems with two continuous 

phases. There are general mixture rules for each of these classes of mixtures 

for predicting material properties. It is most often necessary to combine 

empirical data with theoretical equations to accurately predict specific 

properties of a particular mixture. 

A porous ceramic coating may be characterized as a mixture with one solid 

continuous phase (the ceramic) and one gaseous dispersed phase (the pores). 

In this case, the continuous phase is considered to be the "hard" phase and 

the dispersed phase is the "soft" phase. In addition, the continuous phase 

in a porous ceramic is most likely the major phase; that is, the phase com65 
prising the largest volume fraction. Figure 61 illustrates this type of 

mixture. 

The thermal conductivity of mixtures such as that depicted in Figure 61 have 

been evaluated by Maxwell and applied by Eucken (Ref. 6). Eucken suggests an 

expression for the resultant thermal conductivity of a mixture (km) as 

km = kc (16) [

1 + 2 vd (1 - kc/kd)/(2kc/kd + 1)] 

1 - vd (1 - kc/kd)/(kc/kd + 1) 

where k is the thermal conductivity, v is the volume fraction and the subscripts c and d stand for the continuous and dispersed phases, respectively. 

When kc » kd, the expression for resultant conductivity can be simplified to 

and if kd » kc the expression simplifies to 

km ~ kc [( 1 + 2v d ) / (1 - v d )] 

(17) 

(18) 

Note that Equation (18) expresses the resultant (or effective) thermal conductivity of a mixture solely in terms of the conductivities and volume fractions 

of each component in the mixture. There is evidence to suggest that other 

factors are of equal importance in determining effective conductivity (Ref. 

7). For example, some two-phase systems, such as A1203-Zr03 and MgO-MgA1204' 

can exhibit thermal conductivities lower than either single phase. ThlS 

results from flat microcracks which open along grain boundaries during thermal 

cycling due to differences in the thermal expansion coefficient within the 

material. The small fraction of porosity which results from this cracking may 

Figure 61. Illustration of Mixture Having a Continuous Major 

Phase and Dispersed Minor Phase 

66 
cause a significant reduction in thermal conductivity when the porosity 

occurs as a continuous phase perpendicular to the heat flow path. Similarly, 

the thermal conductivities of other oxides can be significantly altered by heat treating, with little change in porosity or grain size. 

Thermal cycling of some two-phase solid mixtures has been observed to continuously change the effective thermal conductivity due to crack growth during 

cooling and crack annealing during heating. The opening and closing of such 

cracks have caused hysteresis in conductivity measurements of some materials 

(e.g., aluminum titanate) (Ref. 7). This type of effect from thermal cycling 

may be pertinent to the behavior of thermal barrier coatings, especially 

those subjected to high heat fluxes. No reference of this behavior has been 

found for TBC's in particular, but the phenomena is worthy of note. 

Temperature may also influence the thermal conductivity of a mixture by 

affecting the conductivity of each phase. In the case of pores containing 

air, not only does the true thermal conductivity of air increase sharply with 

temperature, but so does the effective thermal conductivity due to radiation 

heat transfer. This is illustrated in Figure 62 where the effective thermal 

conductivity, due to radiation, of a pore having a thickness dp, ;s shown as 

a function of temperature. The thermal conductivity of solid Zr02 is shown 

also, for comparison. It is interesting to note that a pore having a thickness on the same order of magnitude as the coating thicknesses being investigated in this study (0.10 mm), even at high temperatures (1600°C), has an 

effective thermal conductivity of only 3 percent that of solid Zr02. This 

agrees with the previous conclusion that radiation across the individual pores makes a negl i gi bl e contri but i on to the overall heat fl ux through a porous coating. 

> I-

-

> 

Io 

:::l 

cz E O~ 

o 0 

...J ~ «N 

:E E 

a: ~ w () 

J: ~ I- :::: 

w ~ >-

Io 

W 

u.. 

u.. 

w 

10-2~----------~--------------~ RADIATION 

dp = S.Omm 

~ RADIATION 

dp = 1.0 mm. 

RADIATION 

---dp = 0.10 mm 

~RADIATION 

dp = O.OSmm 

10-S __ ~ __ ~ __ ~ __ ~ __ ~ ________ ~ 

o 400 800 12001600 2000 

TEMPERATURE (OC) 

Figure 62. Radiation 

67 
The nature of the interface between two phases ina mixure may al so affect 

the properties of the mixture. In porous ceramic coatings this effect may 

influence the resultant thermal conductivity, since heat transfer across the 

interface must occur by natural convection. 

Other factors that have been considered in their effect on the properties of 

two-phase mixtures having one dispersed phase are the maximum packing fraction 

and the shape, agglomeration, orientation and distribution of the dispersed particles. The packing fraction accounts for the fact that the dispersed 

phase cannot pack in such a manner as to completely fill a space. Nielsen 

(Ref. 7) has described an analytical approach for calculating the properties 

of mixtures which takes into account these latter factors. 

The general mixture rule proposed by Nielsen is: 

Pc 1 + A B vd = 

Pm 1 - B !/Iv d 

(19) 

where A = 1/(KE 1) (20) 

Pc/Pd - 1 

B = (21) 

Pc/Pd + A 

= 1 + C -2 m) cf>d 

cf>m 

(22) 

and Pc = property of continuous phase 

Pd = property of dispersed phase 

Pm = property of mixture 

KE = Einstein coefficient 

vd = volume fraction of dispersed phase 

cf>m = maximum packing fraction 

The constant A depends upon the shape of the dispersed particles, their state 

of agglomeration, their orientation and the nature of the interface. "A" can 

vary from zero to infinity and is related to the Einstein coefficient which 

can be calculated theoretically or derived empirically. The factor !/I takes 

into account the maximum packing fraction which is defined as: 

cf>m = true particle volume 

volume occupied by particles 

Values of m can be estimated or measured in a variety of ways. 

Table 12 presents some values of cf>m which were derived largely from theory. 

Table 13 presents some values of KE derived for mechanical properties of 

mixtures. In general, val ues of KE are usually s1 ightly greater for thermal 

properties than for mechanical properties, but often only small errors will 

result from using the mechanical values for thermal calculations. 

68 
Table 12 

Values of ¢m for Various Mixtures 

Particles Type of Packing ¢m 1 

Spheres Hexagonal close packing 0.7405 

Spheres Face centered cubic 0.7405 

Spheres Body centered cubic 0.60 

Spheres Simpl e cubi c 0.5236 

Spheres Random close packing 0.637 

Spheres Random loose packing 0.601 

Fibers Parall el hexagonal 0.907 

Fi bers Pa ra 11 e 1 cubic 0.785 

Fibers Parall el random 0.82 

Cubes Random 0.70 

Rods LID = 4, random three 0.625 

dimensional (approx.) 

Rods LID = 8, random three 0.48 

dimensional (approx. ) 

Rods LID = 16, random three 0.30 

dimens i onal (approx.) 

Rods LID = 40, random three 0.13 

dimensional (approx.) 

Rods LID = 70, random three 0.065 

dimensional (approx. ) 

The original Einstein equation, derived for the viscosity of suspensions of 

rigid spheres, can be generally expressed as: 

(23 ) 

for low concentrations of the dispersed phase when Pd/P~ is very large. This 

form of the equation may be applied to porous coatings lf P is used to represent thermal resistivity, since the resistivity of air is much greater than 

that of the ceramic. This expression provides a simple method of determining 

KE from experimental data. Equation (23) can thus be transformed to: 

km / kc = 1/ (1 + KE v d ) ( 2 4 ) 

for low values of vd and large values of kc/kd. This is referred to here as 

the simplified Einstein equation. 

69 
Table 13 

Values of KE for Various Mixtures 

;-__________ ~I ----~------~I--=1K - Type of Dispersed Orientation of Particles 

Phase and Type of Stress 

Dispersed spheres 

Dispersed spheres 

Spherical aggregates 

of spheres 

Cubes 

Uniaxially oriented 

fibers 

Uni ax i ally oriented 

fi bers 

Uniaxially oriented 

fibers 

Uniaxially oriented 

fibers 

Uniaxially oriented 

fibers 

Fibers randomly 

oriented in three 

dimensi on 

Any. No slippage 2.50 I 

Any. Slippage 1.0 

Any. <Pa = <Pm of spheres 2.5/ a 

within aggregate 

Random (approximate) 3.1 

Fibers parallel to 2L/D 

tensile stress component 

Fibers perpendicular to 1.50 

tensile stress component 

Longitudinal-transverse 2.0 

shear 

Traverse-transverse 1.5 

shear 

Bul k modul us 1.0 

Shear. (Approximate) L/2D 

Comparison of Theoretical Conductivity Calculations and Experimental Data 

Three methods of calculating the theoretical thermal conductivity of a porous 

coating have been suggested: (1) the Maxwell theory. (2) the Einstein equation and (3) the simplified Einstein equation. Each of these methods was applied 

to a porous zirconium oxide coating to predict effective thermal conductivity 

as a function of porosity. Thermal conductivity values from the literature 

were used for solid Zr028Y203 and air. both at 1200oK. The results of these 

analyses are presented in Figure 63 and are compared to the experimental data 

presented earlier. 

Theoretical curve A in Figure 63 was calculated from equation (16) based on 

Maxwell's mixture theory. This equation expresses the effective thermal 

70 
1.00~-------------------------------------------------------

0.90 

i: 6 0.80 

>:= 

~ <t 0.70 

()a: 5 ~ 0.60 

~ ~ 0.50 

()...J 

~ ~ 0.40 

~ ~ 0.30 

><~ := ~ 0.20 

0.10 

Kc = 2.31 W/mC 

Kd = 5.66 (T = 12000 'K) 

A MAXWELL EQUATION 

B EINSTEIN EQUATION 

C SIMPLIFIED EINSTEIN EQUATION 

D EXPERIMENTAL 

E EMPIRICALLY ADJUSTED 

EINSTEIN EQUATION 

0.00 t--.-..JI------L_--'-_-'-_-'--_..I...--..JI------L_--'-_...;;::L. _____ ----' 

0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00 

VOLUME FRACTION POROSITY 

Figure 63. Theoretical Thermal Conductivity of 

Porous Coating 

conductivity of the mixture in terms of the conductivities and volume fractions of the individual phases~ It does not account for pore size, distribution, agglomeration, orientation, or interface effects. 

Theoretical curve B was established using equation (19) (the Einstein equation). It was necessary to estimate values of KE and ¢~ to apply this equa- tion. A value of KE = 3.5 was selected based on the value of KE for random 

cubes presented in Table 13. The selected value of KE is slightly greater than the value in the table for mechanical properties, to compensate for the 

difference between KE values for mechanical and thermal properties. A value 

of ¢m = 0.70 was selected based on the value for random cubes presented in 

Table 12. 

The properties, P, in equation (19) represent thermal conductivities for this 

analysis. This equation presumably accounts for the distribution, agglomeration, orientation and interface effects of the dispersed pores. It is recognized that the assumed values of K~ and ¢m are educated guesses and should be 

investigated in more detail if thlS analytical approach is pursued further. 

Theoretical curve C was established using equation (24) (the simplified 

Einstein equation). Only a value for KE was needed for this analysiS. 

Again, this value was selected as KE = 3.5. 

Curve 0 in Figure 63 represents the experimental data generated earlier. The 

value of solid ceramic conductivity used to generate this curve is a commonly 

quoted literature value, kc = 2.06 W/mC. It should be noted that besides the 

possible error in the value used for kc (~ 9% based on reported data) this 

71 
experimental data curve also represents two measured values, km and Pt, which 

are subject to experimental error. 

It is obvi ous from Fi gure63 that the experimental data does not agree well 

with the theoretical analyses. The three theoretical models all agree fairly 

well with each other in the low porosity region; the largest discrepancy (about 15% difference) being between the two versions of the Einstein equation. The Maxwell conductivity curve lies midway between these two. 

The fifth curve in Figure 63 represents an empirically adjusted Einstein 

equation. In this case two values of KE were calculated from the experimental 

data using equation (24). These values are: 

= 

11.74 @ 10% porosity 

16.32 @ 15% porosity 

= 14 average 

The average value of KE was used in equation (24) to construct curve E, which 

matches the experimental data reasonably well (within 9.5%). Although this 

value of KE is larger than values quoted in the literature, it is acceptable 

since there is no theoretical upper limit to the value of KE' The equation 

for curve E is: 

(25 ) 

Equation (19) (the Einstein equation) was also used to generate an empirically 

calibrated expression for effective thennal conductivity. In this case a 

val ue of ¢m = 0.70 was assumed and val ues of KE were cal cul ated usi ng the 

experimental thermal conductivity data in equations (19) through (22). This 

analysis yielded an average value of KE = 1.06, which was used with the 

assumed value of ¢m to calculate a curve of thennal conductivity vs. porosity. The resulting curve fell almost exactly on the empi rically adjusted curve 

created previously using the simplified Einstein equation. . 

No attempt to empirically calibrate equation (18) was made. This expression 

has no inherent variables which can be manipulated. It appears from Figure 63 that the difference between thi s equation (curve A) and the experimena 1 

data is more than just a constant, and therefore, would probably involve an 

exponential factor of some form. 

Overall, equation (25) appears to provide the simplest, reasonably accurate 

expression for calculating the effective thermal conductivities of the ceramic 

coatings in this study. It is recognized that the experimental data base 

should be expanded to provide further verification of this relationship. It 

is likely that a more accurate value of KE may be determined through additional testing, but that the form of equation (25) will still apply. 

72 
3.3.3 Discussion of Model Development 

A practical method for analytically modelling TBC's has been developed. This 

method basically consists of two parts. First, the material properties of the 

TBC are determined, either experimentally or analytically. Second, these 

material properties are used in a finite element model of the coating system 

to predict the thermo-mechanical response of the coating to specific operating 

conditions. This model can also be used to predict the effects of coating 

process parameters on coating performance. This requires that the effect of 

the process parameters on the coating properties first be determined. Then, 

these materi al property changes can be incorporated into the fi nite el ement 

model, to evaluate the effects on coating performance. 

The material properties used in this study were either measured or taken from 

the literature. A more complete analytical model could be constructed if the 

material properties were analytically derived. This could be accomplished by 

establishing empirical relationships between coating properties and coating 

deposition parameters. Coating properties could then be predicted by monitoring the coating deposition parameters. An alternative approach might be 

to relate all coating properties to easily measured coating characteristics, 

such as porosity. 

The objective of the porous coating model was to establ ish an analytical method of predicting the thermal conducitivity of a coating based on porosity. 

Thermal conductivity was selected for this investigation since it is one of 

the most critical properties in these analyses. It was intended to uniquely 

defi ne the porous structure of a coati ng in terms of geometri c parameters, and relate these parameters to thermal conductivity. These relationships 

would then be empirically adjusted. A real coating could then be defined in 

terms of these geometric parameters and it's thermal conductivity calculated. 

The porous coating model was intended to be developed one step further. A 

relationship between coating deposition parameters and the geometric coating 

parameters was to be established empirically. This model could then be used 

to predict the thermal conductivity of a coating based on its deposition 

parameters. 

However, the porous coating model was not sufficiently developed to achieve 

these results. The geometric parameters selected to characterize a porous 

coating structure proved inadequate. Despite the obvious inconsistencies in 

the model, however, much of the analysis was consistent and produced logical 

results. It is believed that the basic model concept is sound, but requires further development. The geometric coating parameters need to be re-evaluated 

and perhaps redefined or expanded. 

The theoretical models for coating thermal conductivity appear to have merit. 

Empirical adjustment of the Einstein equation produced a reasonably accurate 

expression for thermal conductivity in terms of porosity. However, additional 

empirical data is required to fully evaluate these models. 

73 
3.4 COMPUTER ANALYSES 

3.4.1 Coating Strength Data Analyses 

Coating strength data were generated using the four-point flexure test, for 

the three coating types previously described: those applied at the 20, 30 

and 40 kW power 1 evel s. These are referred to as coati ng types A, Band C, 

respectively. Two sets of coating strength data were generated at different 

points in time. Each set tested three thicknesses of each coating type. Set 

number 1 tested coatings of thicknesses 0.0508, 0.1016 and 0.1524 mm; set 

number 2 tested coatings of thicknesses 0.0254, 0.0508 and 0.1016 mm. Three 

specimens were tested for each coating type and thickness combination. The 

strength data is presented in Table 14a and 14b, along with the measured loads 

and deflections at failure and calculated radii of the neutral axes. Coating 

strength is defined here as the maximum compressive strain in the coating at 

the time of failure (em)' Also reported is the strain at the bond interface 

of the coating (eb), and at the coating mid-thickness (ec)' 

The coating strength data is plotted in Figure 64 as maximum coating strain 

vs. coating thickness. Although coating thickness was not known to affect 

coating strength, it did provide a convenient graphical means for comparison 

of coating strengths. Curves A, Band C represent the original data set and 

curves 0, E and F represent the second data set. Each data point represents 

the average of three strength tests. 

It is difficult to make intuitive conclusions about the relative strengths of the three coati ng types, based on the data presented in Fi gure 64. The 

strengths of each coating type overlap each other, within each data set, and 

exhibit inconsistent trends as functions of thickness. It is apparent, 

however, that the strengths of the coatings in the second set were significantly hi gher than those in the fi rst. The reason for thi s has not been 

determined. Although the coating appl ication parameters were the same for 

both coating specimen sets, there may have been a subtle improvement in the 

appl ication process for set #2 due to the experience of coating set #I. 

The apparently inconsistent results from the coating strength tests are not 

unreasonable. Rarely do ceramic materials of any type exhibit deterministic 

strengths (Ref. 8). It is often convenient to employ statistics in evaluating the strength of ceramic materials. Thus, statistical analyses of the 

coating strength data were used to compare the strengths of the three coating 

types. 

Because of the obvious difference in strengths between the two sets of coating 

tests, each set was evaluated individually. A matched-pair t test (Ref. 9) 

was used to compare the strengths of the three coatings in each set. This 

statistical method is used to make inference about the mean of the difference 

between two matched groups. In this case the strengths of coatings having the same thi ckness were matched to form pai rs of data. Thi s mi nim; zed the 

possible influence of thickness on the strength comparison. 

74 
Table 14a 

Four-Point Bend Specimen Results - Data Set 1 

t P Y Rl Eb Ec Em R 

Inner Interface Coating surface Neutral 

Coating Thickness Load Deflection Radius Strain Strain Strain Radius 

Number Type (mm) (kg) (mm) (mm) x 10-2 x 10-2 x 10-2 (mm) 

1 A 0.0508 15.5 5.18 33.27 -0.801 -0.877 -0.952 33.59 

2 16.0 5.59 31.24 -0.840 -0.921 -1.001 31. 56 

3 16.5 6.10 28.96 -0.889 -0.976 -1.063 38.66 

4 0.1016 21.0 6.58 27.18 -0.703 -0.888 -1.073 27.47 

5 22.5 6.58 27.18 -0.703 -0.888 -1.073 27.47 

6 20.5 6.38 27.94 -0.691 -0.871 -1.051 28.24 

7 0.1524 27.0 8.18 23.11 -0.575 -0.900 -1.226 23.40 

8 28.7 8.18 23.11 -0.575 -0.900 -1.226 23.40 

9 27.5 7.19 25.40 -0.550 -0.847 -1.143 25.69 

10 B 0.0508 17.0 5.79 30.23 -0.861 -0.944 -1.027 30.54 

11 18.0 5.99 29.46 -0.870 -0.963 -1.048 29.78 

12 18.0 5.79 30.23 -0.861 -0.944 -1.143 25.69 

13 0.1016 29.0 7.39 24.89 -0.742 -0.943 -1.145 25.18 

14 24.0 8.18 23.11 -0.775 -0.992 -1.209 23.40 

15 28.7 7.87 23.62 -0.765 -0.978 -1.190 23.91 

16 0.1524 26.8 7.19 25.40 -0.550 -0.847 -1.143 25.69 

17 25.0 6.20 28.70 -0.519 -0.782 -1.044 29.00 

18 27.1 7.77 23.88 -0.566 -0.882 -1.197 24.17 

19 C 0.0508 8.3 8.69 22.10 -1.082 -1.196 -1.309 22.39 

20 10.6 6.99 25.91 -0.965 -1.062 -1.159 26.21 

21 -- -- -- -- -- -- -- 22 0.1016 23.0 8.99 21.59 -0.807 -1.039 -1.271 21.87 

23 22.5 6.58 27.18 -0.703 -0.888 -1.073 27.47 

24 21.5 7.29 25.15 -0.737 -0.937 -1.137 25.44 

25 0.1524 35.0 7.49 24.64 -0.558 -0.864 -1.169 24.93 

26 22.0 11.18 18.80 -0.627 -1.027 -1.427 19.07 

27 21.0 8.38 22.61 -0.580 -0.913 -1.246 22.89 I 

Table 14b 

Four-Point Bend Specimen Results - Data Set 2 

t P 'l Rl tb tc Em R 

Inner Interface Coating Surface Neutral 

coating Thickness Load Deflection Radius Strain Strain Strain Radius 

Number Type (mm) (kg) (mm) (mm) x 10-2 x 10-2 x 10-2 (mm) 

1 A 0.0254 15.3 8.13 23.18 1.252 1.306 1.360 23.50 

2 14.6 7.87 23.73 1.229 1.281 1.334 24.05 

3 13.8 8.13 23.18 1.252 1.306 1. 360 23.50 

4 0.0508 13.7 7.62 24.32 1.204 1.256 1.308 24.64 

5 13.8 7.37 24.95 1.180 1.230 1.280 25.27 

6 13.7 6.86 26.38 1.127 1.175 1.222 26.70 

7 0.1016 14.8 7.62 24.32 1.204 1.256 1.308 24.64 

8 17.3 8.13 23.18 1.252 1.306 1.360 23.50 

9 16.0 7.37 24.95 1.180 1.230 1.280 25.27 

10 B 0.0254 14.0 8.13 23.18 1.046 1.154 1. 262 23.47 

11 13.6 8.13 23.18 1.046 1.154 1. 262 23.47 

12 -- -- -- -- -- -- -- 13 0.0508 17.5 8.38 22.67 1.063 1.173 1.284 22.96 

14 17.1 8.13 23.18 1.046 1.154 1.262 23.47 

15 17.7 8.13 23.18 1.046 1.154 1. 262 23.47 

16 0.1016 20.0 10.67 19.31 1.190 1.319 1.449 19.60 

17 20.1 10.67 19.31 1.190 1.319 1.449 19.60 

18 19.8 9.65 20.58 1.138 1.260 1.382 20.85 

19 C 0.0254 20.0 9.91 20.23 0.837 1.085 1.333 20.50 

20 19.5 9.14 21.33 0.812 1.047 1.282 21.61 

21 17.0 8.64 22.19 0.794 1.020 1.246 22.48 

22 0.0508 22.2 9.65 20.58 0.829 1.073 1.317 20.85 

23 22.0 7.37 24.95 0.741 0.942 1.143 25.25 

24 21.0 9.65 20.58 0.829 1.073 1.317 20.85 

25 0.1016 23.3 9.40 20.94 0.821 1.060 1. 300 21.21 

26 19.3 6.86 26.38 0.716 0.907 1.097 26.67 

27 20.7 6.86 26.38 0.716 0.907 1.097 26.67 

75 
1.5 

DATA COATING 

-

~ CURVE SYMBOL SET TYPE 0 

-

z 

« 

a:: 

ICJ) 

w 

a:: 

::J 

...J 

« 

LL. 

w 

" « 

a:: 

w 

> « 

A 1:::.. 1 A 1.4 B 0 1 B 

C 0 1 C 

0 \1 2 A 

1.3 E @ 2 B 

F 0 2 C 

1.2 

1.1 

1.0~ __ ~~~ __ ~~ __ ~ __ ~ __ ~ ____ ~ ____________________ ~ 

0.02 0.04 0.06 0.08 0.10 0.12 0.14 0.16 

COATING THICKNESS (mm) 

Figure 64. Coating strength Data From Four-Point Flexure Tests 

The Student's t distributions may be used for inference about the mean of a 

sample group when (1) the population distribution is normal (or at least 

symmetrical and unimodal), (2) the population variance is unknown and estimated by the sampl e vari ance and (3) the sampl e is random. The 1 atter two 

conditions are inherently met by the nature of the coating strength tests. 

However, the strengths of ceramics do not necessarily exhibit a normal distribution. Thus, condition (1) may not be satisfied. To determine how well the 

coating strength data generated in this study matched a normal distribution, 

a Chi-squared goodness-of-fit test (Ref. 9) was employed. A 95 percent 

confidence level was selected to evaluate the hypothesis that the coating strength samples were from a normal distribution. These tests indicated that 

the hypothesis could not be rejected for data set 2 and could be rejected for 

data set 1. The calculated value of Chi-squared for data set 2 was actually 

quite close to the rejection limit. Thus, little evidence of a normal distribution was detected through these analyses. 

Despite this result, condition (1) may be reasonably satisfied by applying 

the central limit theorem. This theorem states that even if a population is 

not normal, the sampling distribution of averages for that population is 

approximately normal for a large number of samples (>30). The sampling 

distribution of averages is the probability distribution associated with the 

sample average (.1). This distribution consists of all possible values of 

'1, for a fixed sample size, and the probabilities associated with these 

values of the random variable. 

In the case of the coating strength data generated for this program, only 

nine strength values per coating type, per data set, were obtained. Although 

76 
this does not meet the minimum sample size recommended to approximate a normal 

distribution, it was used for this analysis and this factor should be considered when evaluating the results. The three coating strengths measured for 

each coating thickness/type combination were averaged to create the data base 

for this analysis. 

The null hypothesis used in these analyses was Ho:Md = 0 and the alternative 

hypothesis was Ha:Md f 0 in which Md is the population mean for the difference in strength between two coatings. This hypothesis was tested to compare 

pairs of coating strength data; three tests were required for each data set: 

A vs. B, A vs. C and B vs. C. The Student's t values were calculated as 

follows: 

t 

where 

Yd/n 

e ya - (e y d) 2 / n 

n - 1 

n = number of pairs 

Yd = strength difference 

Jd

d 

= average strength difference 

S = sample variance 

Mdo = mean strength difference assumed by the null hypothesis 

A confidence limit of 95 percent was used for these analyses. Thus, a = 0.05 

where a corresponds to the probability that t exceeds a corresponding value 

in the Student's t tables. The degree of freedom (Y = n - 1) for these tests 

was always 2. The t val ues are therefore to.025 2 = 4.303 and to 975 2 = 

-4.303. In summary, the probability that -4.303 <'t < 4.303 if Y= 2·is 95 

percent, assuming the null hypothesis is correct. 

The coating strength data used in these analyses are shown in Table 15. The 

t statistics, and some of the intermediate calculated values, for these data 

are presented in Table 16. 

Inspection of the calculated t statistics for data set 1 suggests the following: 

The strength of coating A is not significantly different from the 

strength of coating B. 

The strength of coating A is significantly different from the 

strength of coating C. --

The strength of coating B is not significantly different from the 

strength of coating C. 

77 
Table 15 

Coating Strength Data for Statistical t-Test Analysis 

Data Average Compressive Failure Strain (%) 

Set Thickness 

Number (mm) Coati n9 A Coati ng B Coating C 

1 0.0508 1.005 1.073 1. 234 

0.1016 1.066 1.181 1.160 

0.1524 1.198 1.128 1.281 

2 0.0254 1.350 1.270 1.320 

0.0508 1.260 1.270 1.430 

0.1016 1.290 1.260 1.170 

Table 16a 

Matched Pair t-Test Calculations - Data Set 1 

A vs. B A vs. e B vs. e 

Thickness Difference (Difference)2 Difference (Difference )2 Difference (Difference)2 

(rom) Yd1 

2 

Yd2 Yd/ 2 Yd1 Yd3 Yd3 

0.0508 -0.068 0.004624 -0.229 0.052441 -0.161 0.025921 

0.1016 -0.115 0.013225 -0.094 0.008836 0.021 0.000441 

0.1524 0.07 0.0049 -0.083 0.006889 -0.153 0.023409 

/--

2: -0.113 0.022749 -0.406 0.068166 -0.293 0.049771 

Average 

difference -0.0377 -0.1356 -0.09767 

Yd 

(Sample 

variance)2 -0.007779 0.00661 0.01058 

S 2 d 

test 

statistic 2.219 -8.649 -1.645 

t 

78 
Table 16b 

Matched Pair t-Test Calculations - Data Set 2 

A vs. B A vs. C B vs. C 

Thickness Difference (Difference)2 Difference (Difference)2 Difference (Difference2 ) 2 2 2 (rom) Yd1 Yd1 Yd2 Yd2 Yd3 Yd3 

0.0254 0.08 0.0064 0.03 0.0009 -0.05 0.0025 

0.0508 -0.01 0.0001 -0.17 0.0289 -0.16 0.0256 

0.1016 0.03 0.0009 0.12 0.0144 0.09 0.0081 

2: 0.10 0.0074 -0.02 0.0442 -0.12 0.0362 

Average 

difference 0.03333 -0.00667 -0.04 

Yd 

(Sample 

variance)2 0.00203 0.02203 0.00577 

·2 Sd 

test 

statistic 1.281 -0.0778 -0.912 

t 

These statistical conclusions can be interpreted to indicate that there is a 

strength difference among the three coating types. Since A and C are probably 

different, and B cannot be distinguished from either A or C, it can be assumed 

that B has a strength somewhere between A and C. Thus, these results suggest 

that there was an increase in coating strength associated with an increase in 

plasma spray power level for the coatings in data set 1. 

Assuming that the conclusions made from the analysis of data set 1 are valid, 

a plot of average failure strain vs. power level was constructed; see Figure 

65. Table 16 shows the pertinent data for this curve. This figure serves to 

illustrate the results of the data set 1 strength analysis and is not intended 

to be taken conclusively. 

Inspection of the calculated t statistics for data set 2 suggest the following: 

The strength of coating A is not significantly different from the 

strength of coating B or C. 

The strength of coating B is not significantly different from the 

strength of coating A or C. 

79 
-1.31======~ - DATASET2 ~ o 

-

z 

~ 

a:: 

~ -1.2 en 

w 

a:: 

~ 

-1 

-

~ 

u.. 

w -1.1 

" ~ 

a:: 

w 

> 

~ 

- 1.0 L.-___ ..L.-. ___ -' 

20 30 40 

POWER LEVEL (kW) 

Figure 65. 

Average Coating Failure Strain Versus 

Plasma Spray Power Level 

These statistical conclusions indicate that there was no difference in 

strength among the three coating types in data set 2. This analysis suggests, 

therefore, that there was no effect of power 1 evel on the strengths of the 

coatings in data set 2. The average of all coating failure strains in data 

set 2 was calculated as -1.29 percent. This result was added to the data set 

1 results in Table 17 and Figure 65. 

The results of the strength data analyses for data sets 1 and 2 are basically inconsistent. Data set 1 indicates an effect of power level on coating 

strength whereas data set 2 does not. In addition. data set 2 shows a signi- ficantly greater coating strength than data set 1. These inconsistencies 

suggest that additi onal testi ng is requi red before any concl usi ons can be 

made about the true strength of these coatings or the effect of power level 

on coating strength. 

If we assume that power 1 evel does not have an effect on coati ng strength, 

then each data set represents a characteristic strength distribution for the 

coating system tested. Although the difference in strength between each data 

set is obvious, it is currently unexplained. Thus, although each data set 

theoretically represents the same coating system, they wi 11 be eval uated 

separately, without assuming that either one is incorrect. 

The strength data from data sets 1 and 2 were used to evaluate the probability 

of coating survival vs. coating strain. The coating strength data from each 

set was ranked according to value. The probability of survival (Ps) for each 

data point was then calculated from: 

Ps = 1 - R/(N + 1) (26 ) 

80 
Table 17 

Average Coating Strength as a Function of Plasma Spray Power Level 

Power Average Rang=l Data Coating Level Fail ure 

Set Type JkWJ Strai nl%J 

1 A 20 -1.09 -0.952 to -1.226 

B 30 -1.13 -1.027 to -1.209 

C 40 -1.23 -1.073 to -1.427 

2 A,B,C 20,30,40 -1.29 -1.170 to -1.430 

where R is the rank of the strength value and N is the number of specimens 

in each set. This data is shown plotted in Figure 66. The difference in 

strengths between the two data sets is illustrated clearly in this figure. 

Computeri zed regression anal yses were used to fit fourth order polynomi nal 

equations to this data. These resulting equations are: 

data set 1 : (r = 0.991) 

Ps = -83.2 + 272.1e 2 3 4 319.0 e + 160.4e - 29.3e (27) 

data set 2: (r = 0.989) 

Ps = 693.1 2331.1e + 2926.ge2 - 1622.4e3 + 334.7e4 (28 ) 

where e is the strain in percent, and r is the correlation coefficient. 

Other probability distributions were also evaluated for their applicability 

to the coating strength data. For example, the Weibull relationship (Ref. 

8), which is often found applicable to ceramic strength data, was investigated. The Weibull relationship is: 

inin [1/(1 - F)] = m in (e/eo) (29) 

where F is the probability of failure, m and eo are constants, and e is the 

failure strain. Thus, a plot of 1n1n [1/(1 - F)] vs. 1n(e) should yield a 

straight line if this relationship is valid. Figure 67 shows the strength 

data from data sets 1 and 2 plotted in this manner. Linear regression analyses were used to mathematically model the data. The resulting equations for 

the two data sets are: 

81 
100 

-

~ 90 0 

-..J 80 

< 

> - 70 > a: DATASET2 

::::l 60 

en 

LL. 50 0 

DATA SET 1 

> 40 .... 

..J 

30 £Xl 

< £Xl 20 

0 

a: 

D10 

O 

0.90 1.00 1.10 1.20 1.30 1.40 1.50 1.60 

MAXIMUM COATING STRAIN (%) 

Figure 66. Survival Probablility Versus Coating Strain 

2.00 ----------------....-----. 

1.00 

~ 0.00 . 

~ 

--:!:- -1.00 

z 

..J 

3 -2.00 

-3.00 

b.. DATA SET 1 

o DATA SET 2 

-4.00L-__ ~ __ ~ _____ _L __ ~ __ ~ ____ L-__ L_ __ ~ __ ..J 

- 0.05 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 

LN (STRAIN) 

Figure 67. Coating strength Weibull Plots 

82 
data set 1 (r = 0.965 m·= 12) 

£n1n(l/(l- F) = 12.12 in(e) - 2.075 (30) 

data set 2 (r = 0.970 m = 16) 

tntn (1 /(1 - F) = 16.16 £n(e) 4.577 (31) 

where r is the correlation coefficient, and m is the slope of the curve (or 

the Weibull modulus). 

Inspecti on of Fi gure 67 shows the marked difference in strength observed 

between data sets 1 and 2. There is also a slight difference in the slopes 

of the two curves. The slopes of these curves are referred to as the Weibull 

moduli ["m" in equation (29)]. A large value of m indicates a small spread 

in strength values. Therefore, Figure 67 indicates that data set 2 was not 

only stronger but also had a narrower strength distribution. This again may 

be attributed to subtle improvements in the coating application process between the two specimen sets. 

The Weibull probability of failure plots have quite acceptable correlation 

coefficients. Also, the Weibull moduli calculated from this data are in the 

range of values quoted in the literature for many monolithic ceramics (m = 4 

to 20). The significance of this has not been determined. 

Either of the probabil ity of fai 1 ure di stri buti ons presented for coati ng 

strain can be used (graphically or mathematically) to predict coating reliability under specific operating conditions. Alternatively, the limiting 

operating conditions for a coating can be established for a desired survival 

probability. These analyses first require a thermomechanical model of the 

coating system, to calculate the strain in the coating. This strain can then 

be used to predict the probability of coating survival, using the probability distributions just presented. 

3.4.2 Performance Analyses of Actual TBCs 

Performance predi cti ons for the thermal barri er coati ngs produced in thi s 

study were made using the thrust chamber model described earl ier in this 

report. A range of steady state operating conditions was analyzed which 

included a variety of heat fluxes, temperature gradients and coating thicknesses. Measured material properties of the coatings were used; each coating 

type was unique in its value of thermal conductivity. An average thermal 

conductivity was used in some analyses, to predict the effects of other 

parameters. 

Material properties other than the thermal conductivity of the coating were 

obtained from the literature. These properties are listed in Table 1. Since 

thermal conductivity is the only coating property of consequence for steady 

state thermal analyses, the thermal analyses of these coatings are expected 

to be quite accurate. 

83 
The structural analyses depend upon values for the coefficient of thermal 

expansion, Poisson's ratio and the elastic modulus. Although literature 

values were used for these properties, little discrepancy between the structural analytical results and the true behavior of these coatings is expected. 

The coefficients of thermal expansion measured in this study agreed well with 

the literature values, thus giving confidence in the use of the literature 

values in the analyses. The literature values were used because they were 

defined as functions of temperature whereas the measured values were obtained 

at a specific temperature. Furthermore, the use of the elastic modulus was 

often eliminated by evaluating coating strength in terms of strain. Thus, 

Poisson's ratio, which is of secondary importance, remains as the only material property of any uncertainty in the structural analyses. Future studies 

may focus on obtaining accurate measurements of this property, as well as 

better measurements of the other properties, to improve the overall coating 

model. 

Three values of thermal conductivity were used for the coatings in these 

analyses: 0.60, 0.935 and 1.0 W/mC. These represent the highest and lowest 

measured values of the three coating types tested, and an average of the 30 

and 40 kW coatings. This average was taken from the most consistent experimental data that was obtained. 

The cold surface temperature for these analyses was 200°C, unless otherwise 

stated. This temperature was thought to be most representative of a steady 

state rocket engi ne ope rat ion. Some anal yses were performed with a -20°C 

cold surface temperature. 

Parametric coating systems analyses were conducted using four specific heat 

fl uxes:· 16, 49, 82 and 164 W/mm2. These heat fl uxes cover the range of 

values typically encountered in uncoated copper rocket thrust chamber walls, 

the highest value representing an extreme condition at the throat section. 

Applying a TBC will implement a reduction in heat flux through the wall, by 

i ncreas i ng the operati ng wa 11 surface temperature and thus decreasi ng the 

temperature difference between the gas and the wall. 

A simple analysis indicates that this reduction in heat flux may be significant. For example, let us assume that the heat flux to the hot surface of 

the thrust chamber is due only to convection of the hot gases, and that the 

heat transfer coefficient and surface area are constant. Then, any change in 

the temperature difference between the gas and the hot surface will result in 

a proportional change in heat flux. 

Let us assume that the hot gas temperature is 3000°C. Data from NASA indicates that the heat flux through a bare copper thrust chamber wall may reach 

164 W/mm2. Let us further assume that the surface of the copper wall reaches 

its melting point of HOO°C at this condition (extreme case). We can now 

calculate the expected heat flux through a coated thrust chamber wall operating with an acceptable surface temperature of 2000°C from 

= (32) 

84 
where Q = heat fl ux 

Tg = gas temperature 

Ts = surface temperature 

and the subscripts c and b refer to the coated and bare thrust chamber walls, 

respectively. This calculation indicates that the heat flux through the 

coated wall may be only 53 percent of the heat flux through an uncoated wall, 

or 86 W/mm2 at the maximum condition. 

Based on this calculation, it is assumed that the heat flux value of 82 W/mm2 

used in these analyses is representative of the maximum actual heat flux in 

the throat section of a coated high pressure rocket thrust chamber. The lower 

heat fl ux val ues are expected to represent the non-throat secti ons of the 

coated wall, or fart-load operating conditions. The analyses using a heat 

flux of 164 W/mm are useful in establishing trends from the parametric study 

results. 

Figure 68 shows the temperature profiles through a coating system having a 

0.0254 mm TBC with average thermal conductivity, for two heat fluxes: 49 and 

82 W/mm2. These are medium and high heat flux values for a coated high pressure rocket thrust chamber. It is obvi ous that most of the temperature drop 

occurs across the TBC, subjecting the bond and substrate to a relatively 

small thermal gradient, as intended. As also expected, the thermal gradient through the coating is greater, and the surface temperature much higher, with 

the hi gher heat fl ux. The lower heat fl ux requi res a surface temperature approximately equal to the maximum desi rable ceramic temperature ("'1900°C). 

3000~----------------------------------, 

2700 

2400 

6 2100 

° -; 1800 ... 

::s 

as 1500 ... 

~ 1200 

E 

Q) 900 I600 

300 

Measured - K = 1.000 W/moC 

t = 0.0254 mm 

O~ ____ -L ____ ~~ ____ ~ ____ ~ ______ ~ ____ _ 

32.80 33.00 33.20 33.40 33.60 33.80 34.00 

Radius (mm) 

Figure 68. Predicted Coating Temperature Profile 

85 
Figure 69 shows the hoop (or axial) stress profile that results in this 

coating system from the imposed thermal gradients. These curves show the TBC 

and bond coat in compression and the substrate primarily in tension. The 

higher heat flux induces higher compressive stresses in the ceramic and higher 

tensile stresses in the substrate, than the lower heat flux. These differences 

are approximately proportional to the difference in temperature drop between 

the two cases. The magnitudes of the coating stresses are relatively high, 

on the order of 500 MPa. The significance of these stresses will be discussed 

1 ater. 

Figure 70 shows the radial stress profil e through the same coating system. 

The magnitude of these stresses are quite small, (0-3 MPa), but are greatest 

in the copper substrate. Also, the effect of heat flux is most significant in the substrate, whereas the temperatures, hoop and axial stresses were most 

affected in the coating. 

Fi gure 71 represents the temperature profil es through two coati ng systems 

having a 0.0254 mm TBC with different thermal conductivities: 0.6 and 1.0 

W/mC. The heat flux through these coatings is 82 W/mm2. The thermal gradient 

and surface temperature of the ceramic coating are much greater with the 

lower value of k. No difference was calculated in the thermal gradients of 

the bond and substrate between the two cases, and again, this thermal gradient 

is relatively small. In this case, both of the ceramic surface temperatures 

exceed the maximum desirable ceramic operating temperature. 

Figure 72 shows the hoop (or axial) stress profiles through the same coating 

system. Again, the TBCis shown under significant compressive stress (500-800 

500 

0 

- A 

« 

rl.. -500 

:liE 

-tn 

tn 

! -1000 Measured - K 

t 

- 1.000 W/m °C 

- - 0.0254 mm C/) 

-1500 A: Q - 49 W/mm2 

B: a - 82 W/mm2 

-2000~ ____ ~ ____ ~ ____ ~ ____ ~ ____ ~ ____ ~ 

32.80 33.00 33.20 33.40 33.60 33.80 34.00 

Radius (mm) 

Figure 69. Predicted Hoop and Axial Coating Stress Profile 

86 
o.o----------------------------------~~ 

-1.0 

~ -2.0 

Q) 

... 

en -

-3.0 Measured 

K = 1.000 W/m °c 

t = 0.0254 mm 

-4.0~ ____ ~ ____ ~ ____ ~ ____ ~ ______ ~ __ ~ 

32.80 33.00 33.20 33.40 33.60 33.80 34.00 

Radius (mm) 

Figure 70. Predicted Radial Stress Profile Through Coating Using Average Measured Thermal Conductivity 

-() 

0 

-Q) 

... 

::s 

-co 

... 

Q) 

5000 

4000 

3000 

B 

Q = 82 W/mm2 

t = 0.0254 mm 

A: K = 1.0 W/m °c 

B: K = 0.6 W/m °c 

E 2000 

Q) 

t1000 

O~ ____ ~ ____ -L ______ ~ ____ ~ ____ ~ ____ __ 

32.80 33.00 33.20 33.40 33.60 33.80 34.00 

Radius (mm) 

Figure 71. Coating Temperature Profile 

87 
500~------------------------------~----

0 

-

« A 

a.. -500 

:E B -rn 

rn 

! -1000 Q = 82 W/mm2 

- t = 0.0254 mm en 

-1500 A: K = 1.0 W/m °C 

B: K - 0.6 W/m °C 

-2000~ ____ ~ ____ ~ ____ ~ ____ ~ ____ ~ ____ ~ 

32.80 33.00 33.20 33.40 33.60 33.80 34.00 

Radius (mm) 

Figure 72. Hoop and Axial Coating Stress 

MPa). The bond and some of the substrate are also in compression, while most 

of the substrate is in tension. The coating with higher thermal conductivity 

is seen to incur significantly less stress than the other coating, while the 

remainder of the coating systems stresses are identical. 

These initial analyses indicated that a coating thickness of 0.0254 mm may be 

inadequate for very high heat flux conditions. Thus, a series of analyses 

were conducted to determine the effect of coating thickness on coating surface 

temperature at the maximum heat fl ux condit; on expected ina typi cal hi gh 

pressure rocket thrust chamber: 82 W/mm2. The results of these analyses are 

presented in Figure 73, for two values of coating thermal conductivity: 0.6 

and 1.0 W/mC. A desirable maximum ceramic operating temperature is shown 

here as 1900°C. 

Inspection of Figure 73 is very informative. This analysis demonstrates the 

importance of accurately knowing the thermal conductivity of a TBC and of the 

ability to accurately, and precisely, control coating thickness. For example, 

suppose that a 0.015 mm coating with k = 1.0 W/mC is specified for this high 

heat flux application, to operate at the limiting surface tempera,ture. If 

the conductivity of the coating is actually 0.6 W/mC, the ceramic surface 

temperature must be increased to approximately 2600°C to accommodate the same 

heat fl ux. 

Similarly, assume that a 0.009 mm coating with k = 0.6 W/mC is specified for 

this same application, again to operate at the limiting surface temperature. 

If a portion of this coating turns out to be 0.015 mm thick, the ceramic 

temperature may exceed the desi red 1 imit in that area if the heat fl ux is 

88 
-(.) 

-

° 

(1) 

'-

:::J 

-cu 

'-

(1) 

Co 

E 

(1) 

I-

(1) 

() 

«S 

-'-

:::J 

(J) 

C) 

c 

-«S 

0 

(.) 

3000 

2500 

2000 

1500 

1000 

500 

0 

Operating Limit 19000 C 

---------

a = 82 W/mm2 

To = 200°C 

0.000 0.004 0.008 0.012 

Coating Thickness (mm) 

0.016 0.020 

Figure 73. Effect of Coating Thickness on Surface Temperature for Fixed Heat Flux 

maintained. The effect of thickness is seen to be more pronounced for the 

lower thermal conductivity coating, as its slope in Figure 73 is steeper than 

the other. 

Figure 74 shows the effect of coating thickness on the maximum compressive 

coating strain in the TBC, for two coating thermal conductivities (0.6 and 

1.0 W/mC) and a heat flux of 82 W/mm2. Coating strain is seen to increase 

with coating thickness, and at a faster rate for the lower thermal conductivity coating. Again, the significance of thermal conductivity is demonstrated 

in this figure. For example, a 0.010 mm coating with k = 1.0 W/mC operating 

at this condition would experience a compressive strain of approximately 

0.95%. If the coating actually had a conductivity of 0.6 W/mC, the coating strain would increase to 1.50%, for the same heat flux. Thus, coating strain 

may be significantly affected by thermal conductivity, with the effect being 

more pronounced at higher thicknesses. 

Figures 73 and 74 demonstrate the sensitivity of coating performance to 

thickness and material property variations. Thickness variations are signi- ficant due to the thin values required for high heat flux applications. Thus, although the variation required to cause failure may amount to a 100% 

increase in coating thickness, the absolute value may only be on the order of 

0.01 mm. Control of coating thickness to this degree may be very challenging. Also, thermal conductivity measurements made earlier were found to vary by up 

to 15%, for one particular coating type. This variation may have a significant effect on coating performance. The point to be made is that the accep- tabl e operati ng condit; ons for a particul ar coati ng system may be si gnifi89 
2.00 

-. 

~ 

-

° 

c 1.60 

co oV ... 

CJ) 

- ~\~ 

c 

C) ~~ 1.20 .- ~ - *'- co 

0 

(.J 

CD 

> 0.80 

.-

m 

m 

CD 

... Q = 82 W/mm2 

c. 

E 0.40 To = 200°C 

0 

(.J 

0.00 "'"----'---'----'-_-'--_""'--_"-----11...---&._--'-_-' 

0.000 0.004 0.008 0.012 0.016 0.020 

Coating Thickness (mm) 

Figure 74. Effect of Coating Thickness on Maximum Compressive Coat i ng Strai n 

cantly limited by the accuracy and precision of the coating application 

process. 

The curves in Figures 73 and 74 were constructed for a specific heat flux. 

These curves are useful for preliminary coating system evaluations, if the 

heat flux is approximately correct. However, the true effect of TBC variations on coating performance requires an extended heat transfer analysis. For 

example, an increase in coating thickness will most likely cause a decrease 

in heat flux due to an increase in surface temperature. Thus, more detailed 

analyses are required to accurately evaluate a particular coating system for 

a specific application. 

For example, a simple evaluation of the relationship between heat flux and 

coating thickness can be made as follows. Let us assume that the following 

parameters are constant: gas temperature (Tg), col d surface temperature (Tc), heat flow area (A), hot surface heat transfer coefficient (h), and the 

coati ng thermal conductivity (k). For a steady state thrust chamber condition' we can write: 

Q/A = h(Tg - Ts) = k(Ts - Tc)/t (33) 

where t is the coating thickness. 

Solving for t: 

t = k(Ts - tc)/h(Tg - Ts) (34 ) 

90 
If we let ta represent the actual coating thickness and td represent the 

design coating thickness, we can evaluate the effect of thickness variations 

on the coating surface temperature: 

ta/td = (Tsa - Tc)(Tg - TSd/(Tsd - Tc)(Tg - Tsa ) (35 ) 

where the subscripts a and d represent actual and design conditioni respec- tively. Assuming reasonable values for Tg (3000°C) and Tc (200°C) we can 

plot coati ng surface temperature versus normal i zed thi ckness (ta/td) for 

various design surface temperatures (see Fig. 75). Note that at a normalized 

thickness of 1.0, the actual coating surface temperature equals the design 

surface temperature. Actual surface temperature then increases non-linearly 

with coating thickness, rapidly at first, and gradually approaches a maximum 

value. It is interesting to note that coating surface temperature is especially sensitive to thickness variations around the design point. 

Curves like those in Figure 75 can be used to select a nominal coating design 

thickness, based on a maximum allowable coating temperature and a probable tnickness variation. By locating the point defined by (1) a normalized 

thickness based on the maximum possible coating thickness and (2) a maximum 

allowable coating temperature, one can define the coating design temperature 

and thus the nominal coating thickness. For example, if we select a maximum 

allowable coating temperature of 2000°C, and suspect the coating may be 100% 

thicker than the design thickness in some areas (Ta/Td = 2), then the coating should be designed for a surface temperature of approximately 1450°C. 

2800r-----~c~e-r-a-m~ic~M~e~lt~in-g-----------===~----==~ 

26001--~ 

IT 

o 

-; 2400 ... 

::s 

-f! 2200 c)cu 

co. 

=E 

co cu 2000 

OtU CU 

-(,) 

co co 1800 ::s_ -... (,) ::s 

<I: CJ) 1600 

Tempe .... ra:;.t.:..u_r~e~:::.....~~-.:::::;;_-:::::.~..-= __ ~ ... 

1400L--LL-__ J----L--~~~~~~~~~ 

0.00 1.00 2.00 8.00 

Normalized Thickness (T actual/T design) 

Figure 75. Effect of Thickness Variations on Surface Temperature 

91 
The analytical results shown in Figures 73 and 74 indicate that the TBC's 

under study must be very thin to survive the high heat fluxes anticipated in 

high pressure rocket engine thrust chambers. Figure 73 indicates that a 

coating thickness less than 0.015 mm is required to prevent surface melting 

at the highest heat flux evaluated. These analyses also indicate that the 

coatings under study are likely to fail structurally prior to melting. For 

example, Figure 73 shows that a 0.015 mm thick coating with k = 1.0 W/mC 

would operate at the maximum desirable surface temperature of 1900°C. However, Figure 74 shows that this coating would incur a compressive strain of 

approximately 1.30 percent at this condition. This strain corresponds to a 

10 to 50 percent survival probability, depending on which strength data set 

is used in Figure 66. 

Figure 76 shows the effect of coating thickness on surface temperature for a 

coating with an average thermal conductivity of 0.935 W/mC, over a range of 

heat fluxes. This effect is linear for a constant heat flux. This analysis 

demonstrates that surface temperature is much more sensitive to coating 

thi ckness changes at hi gh heat fl uxes, than at low heat fl uxes. The heat 

fluxes shown in this figure cover the range of anticipated operating conditions 

for a typical high pressure thrust chamber application. Again, it can be 

seen that a very thin coating (less than 0.02 mm) is required to accommodate 

-() 

° 

-0) 

-:::s 

-

cu 

-

0) 

Co 

E 

0) 

I0) 

0 

cu 

--:::s 

CJ) 

2500--~----~-----r-----------------------' 

2000 

1500 

Measured - K 

To 

_ 0.935 W/m °C 

= 200°C 

OL-__ L-__ ~ __ ~ __ ~ __ ~ __ ~ __ ~ __ -L __ ~ __ ~ 

o 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09 0.10 

Coating Thickness (mm) 

Figure 76. Coating Surface Temperature Versus Coating Thickness 

for Various Heat Fluxes Using Average Measured 

Thermal Conductivity 

92 
the high anticipated rocket engine heat fluxes (82 W/mm2) without excessive 

surface temperatures «2000°C) for this coating. 

Figure 77 shows the effect of coating thickness on maximum compressive coating 

strain for the same conditions as just described. This effect is linear for 

a constant heat flux. Coating strain is most sensitive to coating thickness 

at the high heat flux conditions. At the highest anticipated rocket engine 

heat flux condition, 82 W/mm2, the predicted coating strain is quite high 

compared to the survival probabil ity curves shown in Fi gure 66. Coati ng thicknesses of this coating type which incur reasonable strains (say <1.0%)at 

this heat flux are quite thin (-0.01 mm). 

Figure 78 shows the effect of coating thickness and thermal conductivity on 

heat fl ux for a specified ceramic surface temperature of 1927°C and a cold 

surface temperature of -18°C. This hot surface temperature is a desi rabl e 

operating limit for the ceramic coating. Heat flux is inversely proportional 

to coating thickness and directly proportional to thermal conductivity. The 

heat flux values shown in this figure cover the range of heat fluxes expected in most rocket thrust chambers. The coating thicknesses required to maintain 

an acceptable surface temperature at the most severe rocket engine conditions 

are again quite small «0.02 mm), even for the coating with the highest 

measured thermal conductivity. 

2.5~---r----~------~--~----------------' 

-

#. 

....... 2.0 

c 

~ 

en -

~ 1.5 

.--«S 

o 

() 

Q) 1.0 

> 

U) 

U) 

Q) 

'-E 0.5 

o 

() 

Measured - K = 0.935 W/m °C 

To = 200°C 

O~ _____ ~ __ ~ __ ~ __ -L __ ~ __ ~ ____ ~ __ ~ __ ~ __ ~ 

o 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09 0.10 

Coating Thickness (mm) 

Figure 77. Coating Strain Versus Coating Thickness for 

Various Heat Fluxes 

93 
C\I -

E 

E 

-

3: -

>< 

:::l 

LL 

-a:s 

Q) 

z: 

110 

100 

90 

80 

70 

60 K2 

50 

40 

30 

20 

10 

0.000 0.020 

K1 = 1.00 W/M °C 

K2 = 0.60 W/M °C . 

TH =T 1927°C 

To = -20°C 

0.040 0.060 0.080 

Coating Thickness (mm) 

0.100 0.120 

Fi gure 78. Heat Fl ux Versus Coati ng at .Constant L\T 

Figure 79 shows the effect of coating thickness and temperature differential 

on the heat flux through a coating having an average measured thermal conductivityof 0.935 W/mc. The cold surface temperature for these analyses is 

-l8°C. The cerami c surface temperature associ ated with the upper curve is 

2200°C, which is greater than the desired maximum coating temperature, but 

less than the ceramic melting point. The lower curve represents a moderate 

ceramic surface temperature. Heat flux is inversely proportional to coating thickness and directly proportional to the temperature differential. The 

heat fluxes in this figure cover the range of values expected in the rocket 

engine thrust chambers. Again, the coating thicknesses required to accommodate the higher heat fluxes, even with an excessive ceramic temperature, 

under these conditions, is very small. 

The performance curves presented in this section can be used to evaluate the 

general behavior of the TBC's produced in this study, under anticipated high 

pressure rocket thrust chamber operating conditions. These curves are also 

useful for demonstrating the effects of various parameters on coating performance. The analytical thrust chamber model can easily be used to evaluate a 

particular coating system under more specific operating conditions, and can 

be readi ly refi ned to incorporate more accurate materi al propert i es and/or 

boundary conditions as they become available. 

3.4.3 Coating Design Technique 

To illustrate an analytical coating design technique, the models described in 

this report were used to establish design criteria for the coatings produced 

94 
120~----------------------------------~ 

100 

N -

E 80 

E 

-

~ 60 

>< 

:::s 

i:i: 

c; 40 

CD 

:I: 

20 

Measured - K = 0.935 W/m °c 

To=-18°C 

o~ ____ ~ ____ ~ ______ ~ ____ ~ ____ ~ ____ ~ 

0.000 0.020 0.040 0.060 0.080 0.100 0.120 

Coating Thickness (mm) 

Figure 79. Heat Flux Versus Coating Thickness for Two ~Ts 

in this study for rocket thrust chambers. These design criteria are essentially operating limitations and coating thickness limitations which must not 

be exceeded, to assure successful operation of the TBC. The physical limitations of the coating dictate the design criteria. The phYSical limitations 

considered for these anal syses were maximum compressive coati ng strai nand 

maximum coating temperature. 

The maximum allowable coating strain was established using Figure 66. 

Strength data from the first, and weaker, specimen set were used to evaluate 

the "worst" case. A survival probabil ity of 95 percent was arbitrarily selected as a design criterion for the coating. Using the curve in Figure 66 for 

data set 1, a maximum allowable coating strain of one percent was established. 

A coating temperature limitation of 2000°C was also selected, based on information from the literature and past experience at Solar. 

The coating system model of a rocket thrust chamber wall was used to evaluate 

the performance of the TBC's under steady state operating conditions. Various 

operating conditions, coating thicknesses and coating types were evaluated to 

define those combinations that cause excessive strain or temperature in the 

coati ng. A constant col d surface temperature of 200°C was rna; ntai ned for 

these analyses and a range of heat fluxes was investigated. 

Two coating types, A and C, were evaluated. These represent the 20 and 40 kW 

coatings respectively. For these analyses, the coating types were differentiated only by the thermal conductivity of the coating. It was assumed that 

both coating types reflect the same strength distribution. 

95 
It was found that coating strain, rather than temperature, was always the 

limiting factor in these analyses. The results are presented in Figure 80 

which shows maximum allowable heat flux through the coating vs. coating thickness for the two coating types. These results corroborate past experience which indicates that thin coatings can tolerate greater heat fluxes 

than thick coatings, without failing. These analyses show the inverse relationship between maximum allowable heat flux and coating thickness. 

It is again noted that the high heat fluxes anticipated in high presure rocket 

engines (~82 W/mm2) require an extremely thin coating «0.02 mm) for structural integrity. Also of interest is the significant effect that coating properties and coating thickness can have on the allowable heat flux for thin 

coatings. 

3.4.4 Discussion of Analytical Results 

Coating systems analyses were conducted to establish the performance characteristics and operating limitations of the Tacs produced in this study for 

high heat flux applications. The finite element model that was used in these 

analyses simulated the coated throat section of a thrust chamber, which 

experiences an extremely high heat load. Previous calibration of the finite 

element model, including an evaluation of the significance of element aspect 

70.00 

60.00 

N 

E 50.00 

E 

3: 40.00 -

>< 

:::s ~ K = 1.38 W/mC u: 30.00 

-a:s 

(1) 

J: 20.00 K = 2.32W/mC 

10.00 

0.00 ~_-'-_---L __ -'-_--'-__ L.-.._-'-_---L_----' 

0.000 0.020 0.040 0.060 0.080 0.100 0.120 0.140 0.160 

Coating Thickness (mm) 

Figure 80. Maximum Heat Flux Versus Coating Thickness at 1% Strain 

96 
ratio, provides confidence that the model works properly. However, the 

accuracy of these results ultimately depends upon how closely the model 

simulates the actual coating system. This is determined largely by the 

accuracy of the geometry, boundary conditions and material properties. 

The geometry of the finite element model is that of a cylinder. Although the 

actual thrust chamber throat is a curved surface, it is expected that the 

model geometry accurately represents a very small length at the middle of the 

throat, where the slope of the surface is parallel to the gas flow. The 

boundary conditions selected for these analyses covered ranges of anticipated 

operating conditions, based on the best information available from NASA. The 

most accurate material property data available were used in this model, and 

were specified as functions of temperature. The results of these analyses 

are therefore thought to be reasonably accurate. 

These results suggest that the coatings evaluated in this study must be very 

thin to be practical for rocket thrust chamber applications. The thicknesses 

required to assure structural integrity at the high heat flux operating 

conditions are on the order of 0.010 mm. (This is based on the assumption that coating failure results from compressive thermal stress.) 

Furthermore, it is obvious from these analyses that variations in coating 

thickness and material properties may significantly affect the performance of 

very thin coatings. Precise control over coating thickness and accurate 

knowledge of the coating properties may be essential to the successful operation of TBCs in high heat flux environments. Because of the thinness of 

these coati ngs, adequate control of the coati ng thi ckness may present a 

formidable challenge. Methods of accurately predicting coating material 

properties and precisely controlling coating thickness may each require a 

Significant development effort. 

Two basic approaches for making these coatings more compatible with the high 

pressure rocket thrust chamber requirements have been defined: 

1. Develop a stronger coating to insure that surface temperature, rather than coating strain, is the primary limitation for heat 

fl ux. Then, by mai ntai ni ng the lowest col d surface temperature practical, the heat flux through the coating can be maximized for a 

given coating thickness. 

2. Develop' a technique for controlling coating thermal conductivity. Then heat flux can be regulated within the material limitations of 

the coating system. 

Both of these approaches require additional coating development work. 

3.5 TEST THRUST CHAMBER COATING 

A test thrust chamber is shown in Fi gure 81. Thi s chamber has an inner 

di ameter of 6.6 cm. Coati ng of thi sinner di ameter requi red the use of 

different equipment than had been used in the previous program development 

97 
Figure 81. NASA Thrust Chamber for Test Coating 

steps which employed a standard 7M gun configuration. The gun was changed to 

use a 7MBT-12 angle extension. The anode for this unit is designed to deflect 

the plasma gas at an angle of 45 degrees to the gun axis for coating the 

internal surfaces of small cylinders. 

The equipment changeover required minor adjustments in the deposition process 

to achieve coatings equivalent to those initially established. Copper rings, 

6.6 cm in diameter by 1.25 cm long, were fabricated from 0.81 mm copper sheet. 

The ri ngs were mounted ina rotati ng fi xture and the inner surfaces were 

coated. Forty-two rings were coated initially. Arbitrary adjustments to the 

gun parameters were made until coating appearance and deposit rates appeared 

similar to those obtained with the standard equipment. Initially, wide variations in coating quality were observed. These were traced to rapid anode 

wear. Normally these variations would not be significant when conventional 

coating thickness were being applied. However, the very thin coatings required for this application were sensitive to minor changes in gun characteristics. Anode usage was then limited to three hours. 

Twenty-eight additional rings were then coated and metallurgically sectioned 

to further refine the coating process. The parametric variations are given in Table 18. These tests included minor adjustments to the spray parameters and thickness variations to verify consistent coating application. 

98 
Table 18 

Specimen Definition in Final Ring Coating Tests 

Thi ck (mil) Gas flow (cfh) Ring 

Specimen Vo 1 ts-amJ~s Volts-amps Bond Cerami c Bond Ceramic 

43 65-500 70-600 1 4 60 80 

44 65-500 55-400 1 4 60 50 

46 65-500 55-400 1 4 60 50 

47 70-600 70-600 1 4 80 SO 

4S 70-600 70-600 1 4 100 80 

49 70-600 55-400 2 4 80 50 

50 70-600 65-500 2 4 SO 60 

51 70-600 65-500 2 4 SO 60 

52 70-600 65-500 1 0.5 80 60 " 53 70-600 65-500 1 0.5 80 60 

54 70-600 65-500 1 0.5 SO 60 

55 70-600 65-500 1 1 SO 60 

56 70-600 65-500 1 1 80 60 

57 70-600 65-500 1 1 80 60 

5S 70-600 65-500 1 2 SO 60 

59 70-600 65-500 1 2 80 60 

60 70-600 65-500 1 2 80 60 

61 70-600 N/A 1.5 N/A 70 N/A 

62 70-600 70-600 1.5 1 70 SO 

63 70-600 70-600 1.5 0.5 70 80 

64 70-600 70-600 1.5 2 70 SO 

65 70-600 70-600 1.5 4 70 SO 

66 60-500 N/A 1.5 N/A 70 N/A 

67 70-600 70-600 1.5 0.5 70 80 

68 70-600 70-600 1.5 1 70 SO 

69 70-600 70-600 1.5 2 70 SO 

70 70-600 70-600 1.5 4 70 80 

The bond coat selected for these test chambers is shown in Figure S2. This 

bond coat was applied at 42 kW as indicated for ring #49 in Table 18. The 

ceramic applied to this ring was not considered to be suitable, so the ceramic 

coating parameters of ring 47, Figure 83, were selected for the 42 kW ceramic 

coating. Table 19 shows the parameters used to spray the cylindrical thrust 

chambers and the original flat test specimens. 

The effect of i ncreasi ng pl asma gas flow and the bond coat can be seen by 

comparing Figure S2 and 83, both were applied at the same parameters except 

for plasma gas flow. The higher flow used for ring #47 caused overheating 

and oxidation of the substrate and poor deposition of the NiCrA1Y. However, 

for the zirconia coating it was found necessary to use a higher power level 

and gas flow to achieve the desired coating structure. The coating applied 

at 32 kW with a lower gas flow had a nonuniform structure with numerous large voi ds. 

99 
Ceramic 

Bond Coat 

Substrate 

Mag: 100X 

Ceramic 

Substrate 

Mag: 500X 

Figure 82. Photomicrographs of 42 kW Bond Coat (Ring 49) 

100 
Ceramic 

Bond Coat 

substrate 

Mag: 100X 

Ceramic 

Bond Coat 

substrate 

Mag: SOOX 

Fi gure '83. Photomi crograph of 42 kW Zi rconi a-8% Yttri a Coati ng 

for Ri ng 47 

101 
Table 19 

Plasma Spray Parameters for Flat and Ring Specimens 

j;lat I Fl at I Flat l--r I I 

Ceramic Ceramic Ce rami c Flat Ring I Ri ng 

Coating Coating Coating Bond Ce rami c Bond 

#1 #2 #3 Coat Coating Coating 

Current (amps) 400 500 600 500 600 600 

Voltage (volts) 50 65 70 65 70 70 

Power (kW) 20 32 42 32 42 42 

Argon (m3/hr) 2.97 2.97 4.25 5.66 2.27 2.27 

Hydrogen (m3/hr) 0.085 0.14 0.20 0.11 0.425 0.453 

Stand-Off (cm) 6.4 6.4 6.4 6.4 3.18 3.18 

Plasma Velocity (m/s) 25.4 30.5 43.7 32.5 -- -- Power Port No. 2 No.2 No. 3 No. 2 No. 2 No. 2 

Cooling Air (kPa) 551 551 551 551 538 538 

Nozzl e GH GH 703 GH 713 713 

Having selected the specific parameters for coating the thrust chamber, a 

fixture was constructed for rotating and cooling the chamber during coating application. The set-up is shown in Figures 84 and 85. A standoff support 

was constructed to allow the plasma gun to exit from the chamber at each end 

of every traverse. Thi s pl atform was then mounted on the end of a holl ow 

shaft. Cooling air for the chamber was introduced into the hollow shaft 

through radial holes and a concentric manifold. This air was then introduced 

into the chamber fuel passages through tubi ng connected to the fuel 1 i nes, 

and used to cool the substrate during coating. A collector ring, not shown 

in the photographs, was added to catch the cooling air and direct it over a 

thermocouple for monitoring cooling air temperature during coating. The 

entire unit was then coupled to a variable speed drive for rotation during 

spraying. Rotational speed was 73.5 rpm which produced a surface speed of 

15.2 mimi n. 

The first cylinder was coated to a nominal ceramic coating thickness of 0.025 

mm. Telatemp Recorders (Telatemp, Fullertin, CA) were usd to determine the 

maximum substrate temperature during coating. A concentric flange with a 6.3 

mm wall was attached to the end of the cylinder for metallurgical sectioning 

after coating. The cylinder was coated and a section made of the test specimen. It was found that the coating shown in Figure 86, was unacceptable. The directional effects of the 45 degree extension gun are evident. The 

coating section shows a shadowing effect from the initial particles much like 

a snow fence with inconsistent voids and uncoated areas. To prevent this, it 

was decided to reverse the cylinder with respect to gun traverse to provide a 

more uniform coating buildup. The effect of this on the ceramic phase is 

evident in Figure 87 with a uniform deposit of zirconia on each side of the 

bond coat projections. Backside cooling of the substrate was inadvertently reduced duri ng thi s test and substrate overheat; ng resulted in excess i ve 

oxidation and poor adherence. 

102 
Figure 84. Overall Thrust Chamber Coating Set-Up 

Figure 85. Coating Thrust Chamber 

103 
(Magnification: 500X) 

Figure 86. First Coating Applied to a Thrust Chamber With a 7MBT 

Extension 

(Mgnification: 500X) 

Figure 87. Coating Applied With Periodic Reversing of Cylinder 

104 
The initial cyl inder was stripped and recoated with cycl ic reversal. The 

maximum temperature of the unit was found to be 121°C using the data from the 

Telatemp recorders. A test section showed the coating thickness to be between 

0.019 and 0.025 mm which is within the error expected for a nominal 0.025 mm 

coating. 

The second cylinder was coated in the same manner to a nominal ceramic coating 

thickness of 0.013 mm. This coating thickness was found to range between 

0.013 mm and 0.019 mm with the majority of the coating being close to the 

lower limit. This coating closely followed the contours of the bond coating 

without the leveling effects of the thicker coatings (see Fig. 88). It is 

expected that this might create a heat transfer problem in the anticipated 

service environment. 

The third cylinder was coated to a thickness of 0.076 mm in the same manner 

as the previous two. No significant change in coating characteristics was 

observed. The fourth cylinder was coated to a thickness of 0.025 mm. This 

coating appeared to be slightly more porous than the previous specimens, but 

in the thin coating no quantitative differentiation could be made. 

The change from flat specimens and rings, used for initial studies, to a small 

diameter cylinder introduced variations in the coating procedure. The limited 

access prevented using gun standoff as a process parameter and requi red 

adopting previously developed parameters in a new gun configuration. Also, 

parameters developed for gun power levels at significantly higher values than 

40 kW could not be used. 

(Magnification: 500X) 

Figure 88. Coating Applied to a Planned Thickness of 0.013 mm 

105 
In practice, the coating power levels developed for the ceramic coating proved 

to be effective. It was necessary to adjust powder and plasma gas flow to 

provide a duplicate zirconia coating. Applying a reliable NiCrA1Y bond coat 

was less successful and this process needs to be improved. The gases entrapped in the cylinder during spraying reacted to minor disturbances and unmelted 

oxidized particles were entrapped in the bond coat, preventing proper adherance. Bonding of the NiCrA1Y to the copper substrate was not fully reliable 

and no nondestructive test was found. Throughout the cylinder coating tests 

this bond was found to be the weak point of the system and generated several 

problems in reliability that have not yet been resolved. 

Four additional test chambers were coated at a later date. These coatings 

were applied with several process modifications to reduce substrate oxidation. 

The modifications consisted of improved substrate cooling, inert gas injection 

at the plasma gun tip, increased powder feed to reduce the time required to 

apply the bond coat, and improved substrate surface preparation using a larger 

grit si ze. 

All of the coated test chambers are more fully documented in a separate lab 

report issued to the NASA program manager: Solar Report No. SR85-R-2151-00 

(01/31/85). 

106 
4 

CONCLUSIONS, 

1. A practical analytical model of a TBC has been developed. This model 

combines empirical material property data with finite element analyses 

to predict the performance of the coating in service. This model can be 

used to evaluate the effect of coating deposition parameters on coating performance by fi rst determi ni ng the effect of these parameters on the 

coating properties. 

2. An analytical model of a porous coating structure was partially developed. This model requires additional development but appears to have 

potential for ultimately correlating coating thermal conductivity with 

deposition parameters. 

3. Theoretical models of mixtures appear to have potential for predicting the thermal conductivities of TBC's. These models require empirical 

adjustment but may provide an accurate and practical method of calculating coating thermal conductivity as a function of porosity. 

4. A practical method of testing coating strength has been developed. The 

four point flexure tests provide a simple and economical method of 

measuring coating strengths. Plastic deformation of the copper substrate 

causes complications in the strain analyses and contributes to the 

experimental error of the measurements. Due to the statistical nature 

of ceramic strengths, sets of 30 specimens are recommended per test to 

obtain reliable results. 

5. A convenient method of utilizing the coating strength data is to establish a failure probability distribution for each coating type. Comparison of strength distributions can reveal differences between coating 

types. These failure probabilities can be used to statistically predict 

coating failure. 

6. The coating stregth data generated in this study produced conflicting results. Two sets of data were generated at different times. The second 

set showed a significant increase in strength and a narrower strength distribution. Also, the first data set indicated an effect of plasma 

gun power level on coating strength whereas the second set did not. No 

conclusions are drawn as to the effect of power level or about the true 

strength of these coatings. 

7. Performance analyses of the coatings produced in this study indicate 

that extremely thin «0.02 mm) coatings may be required to operate 

reliably in the high heat flux environment of a high pressure rocket 

thrust chamber. 

107 
8. These performance analyses indicate that structural failure is likely to 

occur before the coating melts, for the coatings evaluated. 

9. Two methods of improving the applicability of these coatings to rocket 

thrust chamber applications are {l) increase coating strength so that 

surface temperature becomes the limiting factor rather than strain and 

(2) increase the thermal conductivity of the coating to accommodate 

higher heat fluxes. 

10. These analyses have demonstrated that coating performance is extremely 

sensitive to coating thickness and thermal conductivity at the very high 

heat f1 ux conditions. The need for very accurate knowl edge of the 

coating properties and precise control of the coating thickness is 

apparent at these high heat fluxes. 

11. A practical coating deSign technique has been demonstrated. This technique uses the coating performance model to calculate the maximum allowable heat flux through the coating, as a function of thickness, based on 

physical coating limitations (strain, temperature). This provides a 

convenient plot of the recommended operating envelope for a specified 

coating reliability. 

12. Coating tests indicate that the weak point in these coatings may be the 

bond-to-substrate interface. 

108 
5 

RECOMMENDATION, 

1. Additional strength data should be generated to resolve the conflicting 

results obtained to date, and to establish a reliable failure probability 

distribution for these coatings. A method of eliminating plastic deformation of the substrate should be investigated to eliminate a source of 

experimental error. Testing specimens made with a stronger substrate 

(e.g., steel) should be tried. 

2. A method of measuring Poisson's ratio and the elastic modulus of these 

coatings should be explored, to improve the basic coating model. 

3. The porous coating model should be developed further. Alternative methods of characterizing a coating structure to accommodate analytical 

modeling should be explored. 

4. Experimental coating development should be conducted to improve the 

strength of these coatings and perhaps to control their thermal conductivities. The analytical model of the porous coating may be useful in 

developing a technique for controlling the thermal conductivity of the 

coating. 

5. Performance analyses of these coatings should be conducted for transient 

thermal conditi ons. These conditi ons may induce hi gher strai ns than 

those calculated for steady state conditions. 

6. Coating strength should be measured as a function of time and temperature. This could be done by subjecting groups of specimens to high 

temperature environments for specified lengths of time and testing their 

strength with the four-point flexure test. These results could then be 

incorporated into the coating model to predict coating life during a 

specified operating cycle. 

109 

REFERENCES, 

1. Taylor, Raymond E., "High Temperatures - High Pressures", 1979, Volume 

I I, pp. 43-58. 

2. Andrews, Andrew I., "Porcelain Enamel s" , 2nd Ed., The Garrard Press, 

Champaign, IL, 1961. 

3. Price, H. G., et al., "Reliability and Effective Thermal Conductivity of 

Three Metallic-Ceramic Composite Insulating Coatings on Cooled HydrogenOxygen Rockets", NASA Technical Note, NASA TN 0-7392, November 1973. 

4. Liebert, C. H. and Gaugler, R. L, "Significance of Thermal Contact 

Resistance in Two-Layer Thermal-Barrier-Coated Turbine Vanes", NASA 

Technical Memorandum 81483, 1980. 

5. NASAjPWA - 4-Point Testing ••• 

6. Kingery, Bowen, Uhlmann, Introduction to Ceramics, 2nd Ed., John Wiley 

and Sons, 1976. 

7. Nielson, L. E., Predicting the Properties of Mixtures: Mixture Rules in 

Science and Engineering, Marcel Dekker, Inc., New York, 1978. 

8. Ritter, J. R., "Engineering Design and Fatigue Failure of Brittle Materials", Fracture Mechanics of Ceramics, Vol. 4, Plenum Publishing Corp., 

1978. 

9. Dowdy, S. and Wearden, S., Statistics for Research, John Wiley & Sons, 

1983. 

10. Stepka, Francis S., "NASA Thermal Barrier Coatings - Summary and Update", 

NASA Technical Memorandum 790, September 1978. 

11. Amos, D. J., "Analytical Investigation of Thermal Barrier Coatings on 

Advanced Power Generation Gas Turbines", Westinghouse Electric Corp., 

NASA CR-135146, July 1977. 

12. Andress, D. E., "Analytical Study of Thermal Barrier Coated Fi rst Stage 

Blades in an FlOO Engine," NASA CR-135359, February 1978. 

13. Liebert, C. H. and Gaugler, R. Eo, "Significance of Thermal Coating 

Resistance in Two-Layer Thermal Barrier Coated Turbine Vanes", NASA 

Technical Memorandum, 81483. 

111 
14. Esposito, John J and Zabora, Ronald F., "Thrust Chamber Life Prediction - Vol. 1 - Mechanical and Physical Properties of High Performance Rocket 

Nozzle Materials", NASA-CR-134805, Boeing Aerospace Company, March 1975. 

15. Wikes, K. E., et al., "Thermophysical Properties of Plasma Sprayed Coatings", Battelle Col umbus Laboratories NASA CR-121144, March 1973. 

112 
APPENDIX 1. 

COATING STRESS/STRAIN ANALYSIS FOR FOUR POINT FLEXURE TEST 

113 

Plastic deformation of the copper substrate during strength testing required 

that a sophisticated stress analysis be developed to calculate the true 

strains in the coating at failure. The premise of this analysis is that the 

neutral axis of the coated beam specimen shifts towards the ceramic coating 

as the copper plastically yields. The location of the neutral axis must be 

determined to evaluate the strain in the coating. The neutral axis is defined 

as the plane which experiences no axial dimensional change, thus the sum of 

the bending moments about this axis is zero. This provides a method of 

locating the neutral axis by using a moment balance. 

It is assumed for this analysis that the copper substrate behaves as a perfectly plastic material (strain hardening is neglected) and possesses equal 

properties in tension and compression. Material property values for the 

copper and the ceramic were obtained from the 1 iterature (Refs. 10-16). 

These properties are: 

Yield strength, copper 

Elastic modulus, copper 

Elastic modulus, ceramic 

YScu = 70 MPa 

Ecu = 1.12 x 105

4

MPa 

Ecer = 4.48 x 10 MPa 

The bond coat was neglected in these analyses. 

The stress distribution in the coated flexure specimen during testing is 

shown in Figure 89. The stress in the copper substrate increases linearly 

with distance from the neutral axis, until the yield stress is reached. The 

stress beyond this point remains constant and equal to the yield strength of 

the copper. The strain in the ceramic coating is assumed to be elastic until 

the pOint of failure. Thus, the stress in the coating is assumed to remain 

directly proportional to the distance from the neutral axis. 

Three dimensions are required for this analysis: Rl (or R2), tl and t2 which 

are defined in Figure 89. An arbitrary approximation for the radlUS of 

curvature of the neutral axis (R) is used to begin an iterative solution 

procedure. Each iteration involves calculating the extent of the elastic 

zone (e) and the net moment about the neutral axis. If the calculated net 

moment is not sufficiently close to zero, a new value of R is selected and 

the next iteration performed. The true neutral axis is found by iterating 

until the net moment is essentially zero. 

The computer program created for this analysis uses the midpoint of the beam 

specimen, (tl + t2)/2, as the initial "guess" for R. An iteration procedure 

is then used to converge on the true R, subtracting increments of DR 

(typically 2.54 x 10-6 mm) from R for each iteration. Convergence on a 

solution is determined when the sense of the net moment changes sign. The 

true neutral axis is calculated as the average of the last two radii evaluated. The strain in the ceramic coating (EC) is then calculated from 

EC = c/R 

where c is the distance from the neutral axis to the point of interest. 

115 
Substrate 

\ 

Neutral 

Axis Coating 

f 

--T2-+---1 T1 - 1 

Inner Tension 

I 

R2 

X1 X1 

I 

R 

Span 

- YSCIA 

Stress 

- YSCIA 

Compression 

I 

R1 Radius of Curvature 

Figure 89. Stress/Strain Model for Plastically Deforming 

Four-Point Bend Specimen 

116 
APPENDIX 2. 

EXAMPLES OF FINITE ELEMENT THRUST CHAMBER ANALYSES 

INPUT AND OUTPUT 

(THERMAL AND STRESS) 

117 

INTRODUCTION 

This appendix documents the basic thermo-mechanical coating system computer analyses used to evaluate thrust chamber coating performance on this program. 

Representative input and output files for use with the ANSYS computer code 

are presented. A thermal analysis input file is listed on page 121. The 

associated output for this file is listed on pages 122 through 136. 

A stress analysis input file is listed on page 138. The associated output 

for this file is listed on pages 139 through 154. 

Post processing input and output files are listed on pages 155 and 156. This 

analysis calculates stresses at the nodes in the finite element mode. 

119 
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SwANSON ANALYSIS SYSTEMS. INC. HUUSTON. PFNNSYLVANIA 15342 PHONE (412}7'4&-3301l TV.X 51 

TITLE 10.OQ86 SI 8/M4 C 

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123 

7 

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NIJMBER, DF ELEMENTS: 72 

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I TO'NCDE 

? TO NODE 

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ANALYSIS DATA WRITTE~ ON FILF?7 ( ??q LI~JFSl 

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b IN STEPS OF 

A r. :, Y::i - f 1\ G lr;i': F to( 1/'; r; Ai'~ A L Y t> 1 ~ :"l,.:-i r f '" ,,~ v I ~ T (J~; ll. I: f? S (: L A f.! r u h' h I I: f ::; A lJ [: 1, 1 'II' 1 

~ \\ A N ~ lJ 1\ A 1\ A L Y ::; I:: .:; Y S T H'~ 5. JI'; r • H (,lJ:-, T lJ I'.. P F ,,!~ oS Y L V A 1\1 A I:' 311 2 PH lJ" F, I III ? ) 7 LI b - :~U (I 4 i •. i< 'i 1 I) • 

THRUST CHA~ri~h MUC~L 

** *"*I\GIICE ** .. ** Tf-;JS Tii l"r A,,:SYS GFNFkAL PLJkPfJ::iE 

F H, I IF' E L E ~ E to T' CO tv P II T F t-< P R (J t~ lolA I~ • N F T T H FRS wAN oS 0 r~ A 1\ A L Y SIS 

S~STE:vlS, II\C. I\Uk THE r.ORPORATJOh StIPPLYlNr, THE CU~PiJlER 

F A elL I r 1 E S FUR T H J S A 1\ A L Y :; I SA;, S I J "') F A i~ Y f< F 5 P () ~J SIB I L 11 Y FUR .. ____ ,_"_, ____ THE VALIDITY; ACCURACY, OR APPLTCARTI ITY OF ANY RFSULTS 08rAIr~ED FMCrv' ThE-AI·,S'I:r iiySTFr.;. THF USFf< HU1'>l VEhIFV HI':; G~I\ RESiJLTS. . 

-,- .... '--,-- S~A NSCJN'-'AII,-A[Y5TS SYSTFM~-;-INi. ~ T S FNnF A VnRTNr, TO /;1AK E i HE 

ANSYS PROGRA~ AS CO~PLETF.. Ar.r.URATE. ANn EASY TO USE AS 

POSSIBLE. SUGGE'S'TJONS AI';O r.O~1~IFJ.JTS ARE w!Ol r.OMEo. ANY 

ERRORS ENCGUNTEREo l~ EITHER lH!o DOCUMENTATION uR THE --.. ------"l1t SUL T S--Sf. C u[IT-'8t:--I r;VEIlI-HEI"Y KRliU(;rlT i(J OIJR A TTEI. T H).'~. 

- .-.... ______ ,~ *,~*_A "'AU_SJ L,lJ£.IJ(j NS * ,,* ~~ 

VALUF 

~ ~H: ~F; c~,~ Hil Nr·T~t'LT·"-:-:--:--'- - L . N 0 I ) A L HE" T FLU I~ KEy • • 1 

~ASTFR DUF READ KEY. • 1 

/J A IF~ 1 A!-.. TA RLE" Ei'!,.T.Ii 1 E$o .. 0_0_ ,., . _, ? 

""'JFORM TEI".PERATURE •• TEMPERATURE ZE~O SHIFT 

0.00 

4nO.OIl 

.. " .. .... ELi: ~ t. '~T 1 Y P E ~ '" "'" "' .. 

ISO P A i'c ,,_l,!:!~R.fII,A..L S U L Tn. ? - 0 

~.'lfi~fiFR UF- ELEMEI';T TYFE5= 

KFY OPTlG~.S 

? ~ /.I <; 

;:> II fJ 0 tI 

***** TABLE OF ELEtJEI'JT Rf'AI r.uNSTAIHS *'**** NO. 

1'i11r>lHFR OF REAL CtJNSTAr\T SETS= n 124 

IO.Hil7 

t.J " IN(JTPR 

7 

() 0 0 o c 
FLEMENT 

·**TI.ELEtJENT OEFINrHONS u"'** 

NODES MA T TYPE CLASS 

1 i q 8 1 -i--- 1-.---b---------.-

~ :3 10 .'1 2 1 1 0 

3 II 11 10 3 1 1 0 

---~------.-~q.-+~ -H ~ ---------------i : g -------.. - ;, 7 U 13 6 1 I 0 

7 q 16 15 8 1 I 0 

--~---t~·-H·-_t~--16 t---- : -·-8------·--

10 12 1'1 18 11 1 1 0 

II 13 20 1'1 12 1 .1 0 

----l·~--·-~~----H---~g--~§ . ---t --: 2-------

111 17 211 23 16 1 1 0 1~ 18 25 24 17 1 1 0 __ Lh_._ .. J.'t _ _ 2b ._.....2.5 .. __ .1..13 ._ .. _._ ... _ .. _.1 1 0 .. ____ ._ 17 20 27 '::6 1'1 1 1 0 

18 ?1 28 '::7 20· 1 1 0 

19 23 30 '::9 22 2 1 0 

~~ --.-.~~ ---.g----~~ --_1-~- ----------.--. -.-- ~ : g 

?? ?h 33 32 25 ~ 1 0 ?~ 21 311 33 26 2 1 0 

?II 28 35 34 27 2 1 0 ;':I'i 3() ~7 30··.,('1----------·--··--·---·-· .~ 1 (l 

?" 31 38 ~1 30 ~ 1 0 

?7 3? 39 38 31 2 1 0 

?A ~3 40 39 32 ,:: 1 0 ? q ~ II '41 (J Q --. -3 ~- .. --------.--.- -.... - ? 1 0 

~o 3':) 42 41 34 ,:: 1 0 

31 ~7~<J n ~6 ? 1 0 ~? 3H 45 <J4 31 ,:: 1 0 

~~ 3~ 4b uS 36 2 1 0 ~4 ~O 41 ue 39 ,:: 1 0 

:~c; 41 Ui' 47 uO ,:: 1 0 

3h 42 (J9 uK 41 ? 1 0 17 Ihl C; 1 ·-50 ---zl3-- ----- ---- . - .. ---- 3' 0 

~k uS 52 SI 4u 3 1 0 

39 Ub 53 S2 LIS :5 1 0 

40 <+1 '54 53 'Ii! 3 1 0 

U 1 4il 'i"i- 54 -·-·u7 :5 1 Ii 

U? U9 5~ 55 ~K 3 1 0 II~ 51 ~A 57 50 3 1 0 

/14 52 ~G 58 51 1 1 n --.- -'-ll~ '51 bO'---S-9 --'-52 -----········3 1 0 

IIh 54 01 60 53 5 1 0 

47 55 6~ 61 54 3 , 0 

UP. 50 63 62 55 3 1 0 ---U 9 5 M . --65 .-. 6 /I -·--ST ----.---.--.. -... 3 I 0 

50 5~ ~e 65 5d 3 1 0 

'il bil Ii 7 otl 5'1 ~ 1 !) . 

'i~ 61 6d 67 60 3 1 0 ----'i~ ··-62"· '-6"--68 . --61-·-------·------------- 3 1 0 

511 6~ 70 6q 62 3 1 0 

55 65 72 71 bll 3 1 0 

'ih 6h 73 72 6S 3 1 0 ----'i7 "-6T--T4 -73"------06 -3 1 0 

5A 6n 7~ 74 67 3 1 0 

5'1 he; 76 75 68 3 1 0 

bO 71) 77 10 69 :5 , 0 

-6f 7?-·-7Cj -,1l---,-r---------------3 1 0 

f,;; 73 80 ]q 72 ~ 1 0 b~ 74 81 eo 73 3 1 0 

b /j 7"; 8,1 8 1 I II 3 1 0 '---n" 7f,- 113-82--7'5 3 1 0 b{o, 77 84 A3 76 :5 1 0 

h7 79 db as 78 3 1 0 nn 80 87 86 7S 3 1 0 

he; '81 "-'-1\il"'8T-/lO 3 1 0 

70 82 8Y dd 81 ~ I 0 

71 d3 gU P.9 d2 ~ 1 0 7~ 84 91 SO &~ 3 , 0 

NUMAER OF ELEMENTS = 7? MAX I1·jtJM NODE NIJ~:BFR USFD = 91 

125 
----. ""tX,..-------.,y,.------""'"TFfXy--'------" 

(OR R) COl? THETA) (OR Rn 

I 1.24"S O.90000E-0~ o~OOOOnF.OO ----r-----T:-2IJ" S O~-(f{ffit-~O; nMoon·Oo -- 3 1.2945 0.60000E-02 O~OOOOOF+OO 

a 1.2945 o.aSOOOE-O~ o.nnOOOF+OO 

5 ,1.2945 0 JOOQJl.E-OL_~!)_n'<)(IOF+OU ---n---·--r-..-~'n('3 0.1500 OE -0 2 (). 011 () 0 of + 0 0 

7 1 • 2945 0 • 0 0 0 0 0 E • 0 0 n • 0 0 (J 0 II F ... 0 0 

A 1.2958 Q.qOOOOE-O~ 0.000001' ... 00 

9 1 • 295 eo. 7 'i 0 0 0 E - 0':- 0 ~o Ii 0 n 0 F + () II _ -----; 0 -- --------f~ 2 (r58----0-~-~1'l 0 O'O-E'';O ~ -'--0.0011 n OF + 0 0 

11 1.2Q58 0.45000E-0~ o.nooonF+O" 

I? 1.2958 O.3000IlF-O? o.noOOOF+IJO ___ ---'_ 3 ______ ,_L...2,9.5c ·0 .1.5.0_0_0E..~O,2 __ -0, .. JI_"-O{JtlF+O I; 

14 1.2QSa 0.00000£+00 O.tlOO()OF+uO Ie; 1.2472 O.'HlOOOI::-O? O.1l0000FTOO 

1 h 1 .2972 0 • 7 'i 0 0 0 F - 0 ~ 0,. (I 0 0 I)() F'" 0 0 

17 1.2972 O.hOOOOE-n? 1).onOI)I,F"OIi 1'.'1 ------T: 2'1'7 2-----0-:-1.1 c;'b-OOE-;:O;: ---0 ~ i) () 0 0 of -+ 1I 0 

19 1.2972 0.30000E-0':- o.ooonoF+un 

?O 1.297? O.I"iOOOE-I)? tl.nIlOOOF+OO 

? t 1 .;>, Y 7 2 0 .000 () 0 E: + {J 0 11.0 n 0 () II'" TOO -?,,-T. ,>g/jS-------;r::q 0 OOO-F-':'O? -'0. flllll 0 Ill .. T(1 (1 

~ ~ 1 .240 SO. 7 ":l 0 0 0 E -O? 0 • ,) n () 0 II F + I) (/ 

?4 1.298'i O.hOOOOF-O;> O.i'OOOOF+OO , e; 1 • "9 ~ SO. 4 'i 0 I) 0 F -I)? () • 1,0 Ii 0 II F + 0 0 

;>n . - - ---T~ ;n1'l\ '3 --,,-- --if~-3-0 0 InrE;;n?-- -0.1i 0 (l(1l)F + (j C. 

;>7 1.2985 O.15000E-0; o~nOOOOF+OO 

;>H l.29d5 O.OOOOOF+o" o_0oono~+uo 

,9 t • 2990 0 • q 0 0 0 0 f:: - O? 0 • :1 r. ill: () f- + U (, 

~ 0 1 • 2 '1 9 0 --\) • 75!l 0 (, E;'; 0) ? 0 • ,tC) 0 0 0 F + U (, 

'l; 1 1 • .? 9 9 II 0 • h (l () 0 (If- - 0 ? 0 • I. Ii () n II F + (i(, 

~,l 1 • d g 9 0 () • <.ISO 0 G E - (J , n • IHI 0 0 II F + I) n 'l;~ 1.2990 O.~O()OOF-O? O.00000F+no 

14 1 • 29 9 0 -- ---- 0 • 1 'i 0 (T 0 F - (P- 0 . 111) 0 0 (, F .00 ~e; 1.<''190 O.()OOOOE ... OO O.tiOIl\;OF.\"-1 ~n 1.2995 O.qCO()O~-O? o.nnonOF+~O 

~7 1.2995 O.7'i000E-O? (l.OOt)liOF+<iO '~i! - 1'.-2995" -------O-.hlllHdi E'';;IJ~ n __ I) () 0 0 (,F +\1;, 

~4 1.2995 O.4~OOOE-O? 1I.00000F+OO 

40 1.299S 0.30000E-0? o.nOnOoF+OO 

4t 1.2995 0.15000E-O? O.OOOOOF+OO 

--4;>----------- r:-2 '19 'i" 0 • OllllOTlF';:-ii-O ----0: 0 0 0 0 0 F + Ii 0 

a~ 1.3000 o.gOeOOE-O? O.OOOOOF+oO 

4" 1.3000 0.7S000E-0? O.OOOOOF+on 

"5 1.3000 O.hOOOOE-O? O.OOOOOFTOO 

---zrf, 1 • 3000 0 • at; 0 oOr-o;> --(Y:-Ci ((0" of +11 0 

47 1.3000 O.30000E-0? o~oounoF+GI) 

1.18 1.3000 0.lS000F-0? O.OOOO(IF+I;O 4 q 1 .3000 0 • 00000 E + I) 0 IL fI 0 0 0 0 F. () I) .50 1.3039 O~-qTIlrOlJE-;;-('f?'--IT~nOO(jOF+OO 

5t ~1.3039 O.7S000E-O~ 0.00000 .... 00 5~ 1.3039 O.nOOOOE-O~ o~nOOOtlF+OO 

"'53 1.303'1 O.I.I'i000E-02 O.OOOfJ(lF.OO ., 4 I • 3 0 3 C; 0 • 3 ()(mrTt'-1l ?--o-~-n orr (T 11 F-+ 0 0--

.,., 1.303<; O.tSoOOF-O? o.oonOTIF+OO 

Sh 1.3039 O.OOOOOF+OO o.OOOOOFTOO 

.,7 1.3086 O.90000E-O~ OpOOOOOF.OO -<;/l-------T:3(J86------------u-:tSO'CTOF..;-0? --0 _ 0 00 OOF"'OO 

.,9 1.30a~ O.hOOOOE-O? o.oouOnF+OO 

hO 1.308b O.GSOOOF-O? O.OOOOOFTOO 

f, t 1 • 30 B b 0 • 30000 E - O? O. 0 I) 0 0 0 F + 00 ___ _ --h? "I .308., -.----(J ~ I 50 00 F -O? 0.000 O-OF'fOIT h~ 1.30tH, O.OOOOOE+OO O.uOOOOF+OO 

nJ l.3140 0.90000E-02 0.000001'+00 

he; 1.31aO 0.75000F-0? O.OOOOOF+OO 

-ftft ---------r:-314('f------o.-IiUU(f(Jf=-.-n?" ---n',.; ITTnr-o-I1F.O 11 

ft7 1.3140 o.a50DOF-O? O.OOOOOFTOO 

68 1.3140 0.30nOOE-0? O.OtlOOOF+OO 

h9 1.311.10 O.t"OOOE-O~ O.OOOOOF+OO 

----,0 1 • 3 I<iu O. (J (l'!JlHrn-rr0 ---O~ 11 11 11 Ii M: TOO 

11 1-"-,,-2-02--- "(J.-<}-(}-O·U-ttF--ft-;>- .-- -'i}-.'OHO 11111' +(1 0 

7? 1 • .3;:> 0;:> O. 7., 0 DOE -Ii ~ 0.000 1/ OF;. tJ n 7~ 1.3202 O.~OOOOF-O? O.OOOOHF+uO 

71.1 1.3202 O.<.I5000F-O~ O~OOCOOF.OO 

7e; 1.3202 O.30000E-0? O.oOOOOF+OO 

7ft 1.3202 O.ISOOOE-of) O.OOOOOFTOtl ---H----}:-H .. ~t-----;,o8-"-: ~ g R 8 g~~---g~~-~ g-g g-~+-gg------- -------.-----,----, 

79 1.3272 O.7SUOOE-O~ O.OOOOOF+Oo 

80 1.3272 O.bOOOOE-O? O.OOuOOF+OO 

---~~- I ~iffi g:~mgtg~---2~-g-g8,gRHgg--- ---- A~ 1.3272 O.15000E-O? O.OOOOOF+OO 

A4 1.·3272 O.OGOOOE+OO O.OOOO()F+Otl A., 1.33S0 0.900DOE-O? O.OOQOOF;.OU ---g6-------r.:-33'S0 O. 7smr~-___n_::-rro OO'OF+ao---- .. ------

A7 1.3350 O.nOOOOE-O? OpOOOOOF+OO BA 1.3350 O.45000E-O? O.OOOOOF+OO 

A9 1.3350 O.30000E-0? O.OOOOOF+OO 

---q D ---T~33'S,,(i 0 • I 5000 f - 0;:> ('I • O-irn-n OF" +-0'1) - -------------------------

91 1.33S0 O.OOOOOE.OO O.OOOOOF+OO 
---------***** 

MATERIAL 

KXX PROPERTY TAbLt (LINEAR INTERPOLATION) Tf~P KXX TE~P KXX 

---0 :-O--O-.:T2"501fF~1I1j'----2rl3"'Oo.O 0 .1 ?'5"iYol~~lr2f-­

MA TER JAL 2 

KXX PROPERTY TABLE (LINEAR rNTE~POLATTONl 

TEMP KXX TEM~ KXX 

----O--:.-0---ll:-rr&-uoE.;.-oJ--2 !OO~-O--O--:;-f,-oTaF-;;;O ~ -

MATERIAL 3 

KXX PROPEiHr TABU, (LINEAR It-.TERPOLArTUNl T~~~ KXX Tt~P KXX 

----0 ~ 0- -0.52 3 r4t;'-O;:----;g(fir:-(f-(}-~"i>1niF;,(,;l 

/oIATERIAL 

C PROP~RTY fABLE ILINEA~ INTERPOLATTON) TFMP C TEMP C 

O~O-· t ~-OOOO- --- --. - ;?3(1(j":-O-----T~0-OO() • 

t.lATEiHAL 2 

C P,WPFRTf fAt:Lt:: fLlt\EAfi rrvrFPPllLAT1(JNl T~~P C rE~p C 

-- O. t1 1.0000 -2300.0 - 1 ~O()(} 0 

t.fATEIHAL 3 

C PROPE"T, TAHLE lLINEAR I~TF~POLATTONl 

TEMP C TE~P C 

-- 0.0-- 1.0-00 a -----iBOu-:.lJ ----I ~"IlO 0 0·--

/oIATERIAL 

OFNS PROPERTY TABLE (LINEAR INfERPOLATIO~l 

TEMP uFNS TEMP OFNS 

-. -- -0 ~ 0 1. a a a 0----- ---t!3"no:-cr-r.:!TOO 11 --.-----

t.fATERIAL 2 

OFNS PROPEi<Tr fABLE (Ut,EAR INTfRP[)LATflHJI 

TEMP GENS TEMP OENS 

----0.0 -- 1.0000 .--- -""""""230():-O--·1~·noon 

"'A TER IAL 3 

DENS PROPERTY TAbLE (LINEA~ INTERPOLATIO~l 

TF~P OENS TEM~ DENS 

.. - 0.0 ---1.0iJOO----·---Z!0-(j~()" ----T.-ooon 

~AMI~UM ~ATE~IAL N~~HER= 3 

127 
***** ~ASTE~ DEGREES OF FRFFOUM ***** 

NODE DEG~EES OF FREEDO~ LIST 

NIIMAER OF SPECIFIED !VASTER D.D.F.= 0 ""i'fTACliilJMijE'"ROrMAST~o. 0 .F. = 0 

TNTEGFR STO'RAGE REQUIREMENTS FOR MATERIALS. ETr. TNPIIT CP: 1(-..Q40 TII<IF= 10.10826 ~FMORY--.lE__ 92 /#'Et"OR Y II= 0 TOTAL=--- _____ 9? ~_E~Oln_~VA ILABLE= __ 3t1QO(j!l 

*** LOAD STEP OFTIONS SPECIFICATIONS 

NTTTER= -10 NPRINT= . 10 NPOST= 10 

AI L PRINT CONTROLS RESET TO 10 

-ALL POST DATA FILE CG~TROLS RESET TO to ---------.-

Ar.EL= O:OOOOOE+OO O.OOOOOE+OO 0.00000E+00 

~E G A ~-.!l~0 0 0 0 ()E;.+~.O~O OJ.LQ .. QJ; + 0 0 0.0000 () F + n () 

nn~EGA= O.OOOOOE+OO O.OOOOOE+OO O.OOOOOF+OO 

_ r.r.l,Jlr:;:;:-..Jl .• _QO!l.O_QE+JlJLiLs..Q.Q.Q.iLQE+OJL.jl~Q.j)_Q_Q.nt_±Jt_0 ________ . ___ . ____________ _ 

Cr.OMEGA= O.OUOOOE+OO O.OOOOUE+OO o.OnOOOF+oO 

.....DJ1Iv1E:.G A = Q. 0 \I tj_Ollfo_t.tI_O._O.QJLQ P..Qt.!..<l!L..ii_,Jutn () ()F + 00 . 

KTEIAP= o o 

~HEADY STA fF. cOr'jvl:fiGEr,CE Ci-II rER [ON=, 1 ~O()OU TRANSIENT UPTI~IZATIC~ Ci-IITEK[A= ~_Onno 

TFMPERA TUNE L [\11 T= O.OOOOOE+O(J K F Y T 0-- TE K loA rNA Tf,--Rlj 1\ '1 F' -N a- C O","VFRITFNI.E :-- -.. -0 .~. 

I nADS STEPPFD Ta FI~AL ~ALUES FOR ALL [TFRAfTUNS (kR~= I) 

RFACTHii-J FCRCE KF.:Y=I ---------

, IINTI'U';'" fE",r'E"ATuRE= i).OOl) (IREF= (I.OOII) 

RnUNOARY CONO rrrON p"rr;rK[Y=- 0 

128 
_LQAL~IfJ:_NU},J~ .... E,",-,k,--",="---.... __________________ . .,._ .. _______ ... _ .... _. ___ ._ ... _ .. '' ...... __ ._ 

*** LOAD OPTIONS SUN~ARY 

.J1..!'1E = O:!LQ. 000 E'-'+--"O'-"O'---'-'N'-"Ic.:T..!T...,E"'R:.;:=:......_--e·:....1...."O'--_________ ._ .... _ .. _________ . __ .. __ ._ .... _ ... ,_. ___ _ 

A~EL = O.OUOOOE+OO· O.OOOOOE+OO D.OOOOoF+on 

ONEGA = O.OOOOOE+OO O.OOOOOE+OQ O.OOOOOF+OO 

nnMfGA= O.OOOOOE+OO O.OOOUO~+OO U.oooooF+no -r. t. ure ':: o~-O () IlUlJE+lTno·~lJ·O-ITOlJ ITo-oo:OlltYn fI r:+-tfo .--.-.... -.--., ...... - .. -.. -.. _-. - .. -... . .. --_._.--.--. -'-. Ct.OMEG: O.OOOOOE+OO C.OOOOOE+OO O.OOOOOF+OO n~GONE: O.OOOOOE+OO O.OOOOOE+OO O.OOOOOF+OO ~~ ~~~~H~B.~·d-~~~·~t~__'r_U_~ d ~ E P 0 0 IT FR AT I QN.=. ___ o ... _. _________ ._ .. __ ._. __ . _______ .. 

IlNfFORM TEMPERATURE= 0.000 (TRFF: O:O(JO) 

PRrNr--·-E-[-EME:NT-TO-~-C-tS-ANlr~EACn-(fNF-trR_CF.1r-CKR-;:-: 1 l ----,"----.- ... ---- '-~- --_._---- -, ... --.-. 

InADS STEPPED TO FI~AL VALUES FOR ALL TTFRATIONS (KAC= 11 

-N~ ~g h~l ~~h~g~ ~·f1 ~ ~Tf~I.Ls':-oo--AoQJlD-1L_---·-- ----.-.-. ----.-.. - "'- .. -"'-' .. -.-.-... --.' .. --.----... - TFMPeRATURE lIt<!IT: D.OOOOO·E+OO 

KFY TO TER~I~ArE RUN IF NO CCNVFRGENCE: O. ~Il!)~.?.. 0 I S nl~._. __ 1 -----_._----_ ... _----.. _._-- ._ .. 

NPRINJ= 10 NPosr= to REA~rION PRTNT FRFa= 

nISP. POST DATA FkE~= 10 REACT. POST nATA FRE~= 

.. n rSPLACEi,IENT PR [!~ r-F!1--E'.JLt'NtTtS .---.. -. 

FREQ NSTRT wsrOF ~I~r. 

10 1 32COO 1 

- "FLEf.1FNT Pff'INT -'-A'No~--P-rrs r- DATAF"RFQ'Ut~jC IF-S" ~----- .--

TYPE STIFF SfRESS FORCE STRFSS OATA FORCF 

NC. PQII\T PfllfH I)AH lFVFL I)ATA 

.S'i 1 0.1.0 .. 1 0 'I 1 0 

if;,,,;,,, SPECIFlt:u TE'''PE"ATlHIES ***** " _.-------------_._-_ ... -

NonF TEIIP 

A'i 

"Af, 

R7 

All 

Rq ·--··-qO 

ql 

400.0UO 

400~oao 4no.oo,) 

(lOu.QOO 

400.000 

'-'-'-'1J 0 a~'u 0 0 

400.000 

;,***" SPECIFIED HEAT FLOW RATES ***** 

NO: __ .. _._N..QDE __ O_. OF FR. VAlUF ----'--_. __ .- ---_ ....... _. 

I t HEAT O.q70RAOE-0~ 

~ 7 HEAT O.q70RAOE·n? 'I 2 HEAT 0.lQ417'iF.-Ol ------a -' ··-·---r--r.tA'--lr:TQ LrrT5E;; 01 

'i 4 HEAT O.lQ4175E-Ol h 5 ~EAr O.194t75E-Ol " __ ... ________ 7 .___ 6 HEAT O.194t7'iE-Ol _ ___ e_· ____ .. _ . __ _ 

1 U 

10 

***** LOAD SU~~ARY - 7 TE~PERATuRFS 7 HFAT FLOwS - ~'----"-"--'-'--'---

o CONVECT rm!S ***** 

TlH E G FRS T (] RAG E R I: Q Ll REM E N T S F () fI U) A 11 nAT.l J f.j PI, T C P = ? t .? 110 TT 1.1 E = 1 U • 11)..Q.1 /.IFMORY I=--216····tJn'GRY TI=·········---'T TOTAI= ?1f.. MEMORY AVAILARlE= ~uOOOO 

R A t-J G F 0 FELl: i~ E iii T IV A )( r r~ L M CON n u r: II v J T Y T N (;I fJ HAt r. n () R f) ft. '" T E S ~AXr~U~= 0.tdS014E-Ol AT ~Lf~~~r 7~ 

- /<irrHrvlJr~= O.1'Jilf,Q..,E-04 'AT ELEfJEI~T '" 

129 
T t n F G F" 5 T (j i< A I~ to R tall f.I E '" l: h ., S F (j K E Lb·, F r J T F { II{ 1'1 ill A r 1 () I~ C f-' : ? " • c, 2 (, T I j. F;; l!.; 1 .... - /.IFr-'OkY I: ?lit r"tIVC~Y II:: (I TOiAI:: ?IIJ 1,'Er"Ui,y AI'AILARLf::: ~~40()()(} 

*-* FlEME~l STIFF~ESS FOk~LLAT[O~ TINFS 

TYPF NUMAER SIIF TOTAL CP AVE CP 

,------- I' - - 7 i' 

TIME AT Erm OF ELEtJHT STIFFNESS F()R~1IJLATTt)I~ CP : 26.5£10 

-' --- -_ .... " 

: /.IAxI/.ILJ~ lI~·COKE V<AliE FRor,T ALLO"FD FnR RH"lJfS1F[) Mr-~1ORY SIZE:: St;O 

Tl\TEGFk STGflAGE REQLII1EMENTS FOR v..AVE FRONT MATI-iJX SflLljTIOI~ CP:: ?CJ.730TJME= _10.115:;,. 

,-MFMORY-T: --. --214- i'lE1'OR'fTr:----l?f---TOTAt': '1'1C:; r>1EMORY AVAILABLE: 340000 

2!OtJ!Jt.1 __ I~~_~9BLW_"V"; FRONT (EQUA T ION~_LtJ1if_n: ___ ,_ q 

*** MATRIX SOLUTICh TIMES 

REAn T~ ELEMENT STIFFNESSES CP= 1:4~0 NODAL COORD. lkANSFOR~ATIGN CP:: 0.120 

--M A n I)(T k I A NG U[ A f-i 12 A TTOrr-----t P:------- 0 ~ R 80 

TTME AT END OF ~ATRIX TRIANGULARIZATION r.P: ;:>'1.750 

-rNTE GFP--S TDR A (~r kcliCnn:METilB'- to o kRTc 1\ 5IJHST ITllT J 01\ CP= ~o ~ 11l (j 1 J'" F:: 1 C. 11,,: -

IIFMuRY J= 214 i"tMORY II= lAb TllTAI = 4(1) f.1E:MURY tVAILAHLE:= :\/H)(;() 0 

**_ ELEM. HT. FLlft CALC. T]i"FS -r YPE' NUMBER 5 T 1 F- --1 a T ALCP-----,l;VE'CP-----

55 2.390 '0.033 

--* ;0'" -"MID A eli 1'. FLm;- LA'[ L.TT~ .-"--,---

TYPF NUMRE~ STIF TCTAl CP AVF CP 

*** LOAD :IT!:P ITER rOMPLFTFf)~ , T1 t~ E = (I • 00 (I 0 () 0 E 1- 0 0 K uSE:: 0 C U "'1 • I T E ;; • = 

i TNTEGFR STORAGE REQLIREMEhTS FOR HEAT FLOW CAl r.ULATIOhS CP: ~3.8qO TI~F= IO.lls?' 

--llFMORY 1= . ?14 tJEtJORY 1'1='--- 1;,,"- TOTAl:: "i7A MEHORY AVAILAf,LE= 31100(10 

**_ PRO~LEtJ STAlISTICS 

NO. OF ACTIvl-: LJEGREI:'S OF FREFIlN~:: 1;11 ;---R:; ~-. S:;-- W A VEE RON rslZ E'-=---.. -T~--"------------ --

*** ANSYS RINARY FILE STATISTICS 

-RIIFFFR- SIlrGSEu:'2ITlfo--

I RlI?... ~J(jt<O S J 

POST OATA ~RITTt-:N 01\ FILE12 

RFSTART DATA nRITTE~ C~ FILE 3 

TFMPFRATURES ~RITT~h LI\ FILE 4 --.. --- --" ._--_._---------._"""'-;:;.......--------

*** LOAD STEP 1 ITER 2 ~ •.. lHEADY STATE C(JNVERGENCE VALUE: COMPLETEn~ TTME: 

?7 .4 ... 0 I AT NODF 

O:OUOOOOE+OO KUSE:: 0 CUM. ITER.:: 

?2 CkITERION = 1000UO 

2 

"-·~·..,*:-.*:-:l* .... [-,0"rA..,D,.,--S~lrE~p--.,1:---IT"T~EC"Ra---... 3-...,.C'TD"'Mn:PtT[O:F~TF"IT7_'lTr,fF=--o-:rrn-oOOOr'tO(j---KliSF:-o--ruv.-:-"TTER-:-::----y----

RTEADY STATE CUNVERGE~CE VALUE = O.C:;~RII AT NODF ?2 CRITERION = 1.0000 

*** SOLUTION CONVE~GED - LOAD STEP 1 CONVFW~ED AFTFR ITERATION 3 CIIM. ITER.: 3 -----NF Xi--TTE RATlmrTIOENi IF lEO As-TTFRAlTON- 1 01 SATISFIES PRINTOUT' OR POST DA TARE{WEST:---

130 
~Axr~UM TE~PERATURE= 37bO.S 

MINIMUM TEMPEHATURE= 400.00 

A T NON 

AT rWrJF 

t 

91 

--_._-_ .. __ ._ ..... 

*"'*** ELEr~Ei~T HEAT FLOW RATE5 ***** TUIE = O.OIlOOOOF+(JO 

FI EMENT \. 

'-HE A r-FLOw·-Ii~TuNOm:--2~(Od7<fE:';;-(f?--·-··--

HFAT FlO~ INTU NODE q 0.q70H7HE-O~ 

... FAT FLOIN HdU NUDE 8 0.<;7087oF.0;1 

HF ~ T FLl)lt~ __ .J r4Lll iIJ().DE: ___ L -0 •. 97 (jHI\i'J::~Q~_._ ___. __ 

FI t:f.1ENT 2 

HFAT FLOw 

HFAT FLOW 

INfG NOUE 3 -O.970H7~E-O~ 

INTO NUDE 10 0.q7087hF-O~ 

. HF A T FI G':i 

HFAT FLm~ 

li~ TON ODE-----q--O-:-c17 'IB ("iF;;'-');>"'-

INTO NODE 2 -O.970d7nF-O? 

1=1 F.r·tH,T 3 

HFAT FLml Ii\lTfr NaDt"-'-'''~O~-qT<r87C;F';'-0? 

HFAT FLOw rrno NODi:: 11 O.g70(H"iF.-O~ 

HFAT FLOW INfO NODE 10 0.970875E-o;> 

HFAT FLOW INTO ~~Q.Q; __ ~_~1!~?'jF.:. ... .il? 

FI FMFrn I~ 

HFAT FLOW INTO NUDE 5 -O.q7DI\75F-0~ 

... FAT Fl m~ r'~TU NUDE It! O.9701l75F-O? HFAr-FLOri TNTO NODE---n'-"-U-;-GT0875E';;;O?" 

HFAT FLOW INTO NODE 4 -O.G70875F.·O;> 

EI Et-IENT 6 

HFAT FLOIN INTO NODE 

HFAT FLOW INTO 1'1100£ 

-l'FA'T -H-ow--n,TG" "'NGDE 

HFAT FlO~ iNTU NODE 

FI EMFNT 7 

7 .. O.cH08AOF-t)? 

14 O.G7087bE-O? 

1 3 f) • G 7 0 Ii 7 I'!l:-1TT---·--------

6 -O.'l70871.1F.:-O? 

-I-lF AT FL OW rNiO"'I\:ODE q ~-qT(fIT7ht-l\? 

HFAT FlO~ INfO NQUE Ih 0.970H7hF-O~ 

~FA~ FLOW I~TO NODE 15 O.q70~76F.:-O~ 

HFAT FLOw INTO .'~U9_t ____ 8_::Q~G~1\7bE-O~ 

FI Er-.lFriT 8 

MFA T FLO W HI T (J N lJ 0 E 1 0 - 0 • G 7 0 8 7 t; F·;);> HFAT FLO~ INTO N~DE 17 O.G7nM7nF-O? 

HF" A T Fl UW IN TC" NOD E '--n; . -n'~"HO 8 ThE""; 0;> 

HFAT FLUW INTU ~OOE 9 -O.G7C87bE-O? 

FI EMFNT q MFAT FLOW INTO ~ODE 

131 

LUAO STEP: ITER .: 10 CUf'.1. ITER, 
... FAI i- L lJ ~". 11\ I L ,';uUt 1 ~ t: • ..., 7 t: t'. I~'" -11 r 

HFAI ~ L (IV, JhTl I\ULd:, 1 I O."7Ch7'if-C;; HFAT Fl (jr: Ii, TO ,\juDI: 10 -L.C;70"7,,F-O? 

FI E.~'F /, T 10 

HFAT FLOW Ii, T L M1DE 12 - () • C; 7 \) I' , h r - (, ? 

HFAT FUjV\ 11<1(' r,QUE ", 1 '1 (i.'170(;7S1--0? HFAT FUJI'i- INTO ~.llll E fe ',u • '17 (j i': , bE';' n? -. HFAT FLOI'i INTO tiULJE 11 - 0 • 'l7 0 h 7 5 E - (I? 

FI EMENT It 

'HF Ai F[ OLi.'" I NTCNOl1[--f3--.-(;-.-c:;nIl7't;F;;;(J;;--

HFAT FLOI'i INTO N00E 20 O.'l70h7bF-O? 

HFAT FLOW I~TO NODE 1'l o.Y7081bF-O? ..Ji£AI,£.hO~_ IlIIJ:Q_N QJ:>;_,J.?._-_(j_~1.Q Ii 7 5 F - O?_ 

FI EMENT 1 iJ 

HEAT FLOW INTO NODE ]~ -O.'l7087bE-O? 

HFAT FLOW INTO NODE 21 O.970A7bE-O? -HI" Arnll rr'Il~TO"fiWD E"'-2-(j-U:-crHI 81TI:'-ri2---

HFAT FLOw INTG NODE 13 -O.C;7087bF-O? 

FI EMENT I'! 

-H FAT F L 0I'r I N TO Il on E--lll-o:-il t OB-7bE::O?-

HEAT FLOW INTO NODE 23 O.<;70B76E-O? HEAT FLOw INTG NUDE 22 O.'170B76E·O? _~UL F LWL.I IHD "f\; 0 OC ..... ...J.5.-.::..Q, .. ,21JlJl..H!.E..~QL_. 

FI EMf NT I ~ 

HFAT FLO~ INTU NODE 17 -0.Q70H7hF-O? HFAT FLOV. INTC NUL)E 211 (;.<;70B7bF-(,? HE AT FLOv."'IfvTD NOut-'-'?r--u~ 'l7lT&7br.;()?'-

HFAT FLOW INTC ~ODE 16 -O.q7~87bE-O? 

FLE/o',FNT IS 

H F'AT FI. OW nn c l~ 0 Df' - 1 B -';;lr-:-GTllrfTGF;"-CP-'--

HFAT FLO~ INTU NU0E 25 0.Q70A7hf-O? I-IFAT FLO,," If,,TL (H)UE <'II Ii.G7ue7hE-O? 

HFAT FLOV'; 1,,,T( I,OUE ""L7......:.L." .. LOIl}bt-('? 

FI EMEfl:T 10 

HFAT FLm,1 JI~TG 

HFAT FLOw Il\fO ~ODE 19 -0.G70R7bF~O? 

NeOE 26 V.G7087bF-O? '--HFAT FLOW rr,TC 

HFAT f-LOW HnG ~;OD E-- "-25' '-0' • ., 7 0 87 hF -Ii?' 

NGuE 1~ -D • .,7C67"F-O? 

FI EMFNT 17 

HFA T Fl OW' INTO NUDE -'-2n"-O'~'C;7 tTB76E';'07 

HFAT FLOW IN10 NODE <'7 G.~7087bE-O? 

HFAT FLOW lNTU NODE 26 ~.<;70876F-O? 

HFAT FLOW INTO NODE 19 -0.'170876E-07 --.- ---- --_. - -_._ ... _--,--- -. __ .---_ .. _---- --". ---.'~----' 

FI EMFNT 18 

HFAT FLOW lr,T() [.GuE 21 -0.G7ui':7bF-fJ? 

I-IFAT FLOW linG "OUE 2I:J 0 • .,70h7bF·O? "f.!FJlT FLOW"TNT() JliODE-'?7-0'~'i708ToE-O?-' 

HFAT FLOw INTO NODE 20 -0.'l70876F-O? 

EI EMENT 1 q f.! F ATFI. 0 Vi Trn'(;"-r'iG D E"'--. '2 3"'-~G 70-8 TbF';'C, ?" 

HFAT FLOW INTO NODE 30 0.970B7hF-O? 

I-IFAT FLOw INTO NODE 2'l O.S70h7bE-O? ~!_.F_~9_~ _,I N.!Q ,,~~.9.Q~ __ ,2 "2 -.Q.~~LQ.il..?"'~:-_O ~ .... , ", 

FI EMENT <'0 

!-IFAT FLOw INTO NODE 24 -0.<;70876E-O? HF~T FLOW INTO NODE 31 0.97087&E-02 

FI FAT Fe 0 iV'lNit'i-O -N Ol'"rDrrEr:------3'r-o;;.--,Cri-=1 ."t17ll1iTf;t;;~ 

HFAT FLOw INTO NODE 23 -O.G7CR76F-O? 

FLEMFNT 

HFAT FLOw 

HFAT FLOW - ffF"1iT FLO W 

HFAT FLOW 

22 

INTO ~ODE 

INTC rWDE 

I l'JIOI'V ODE 

INTO NODE 

26 -O.970R7bf,-O? 33 O.'l70B76E-IJ? 

32 1I.'l7'!T'BT~nT--

25 -O.970B7bE-0? 

132 
1'WA-f-ftfrW-f1'tTO t1(t[H:~1I o. CjTtT8Ttrf'-..-tr;>--

HFAT FLOW INTO NOD~ 33 0.970B7hE-02 

HFAT FLOW INTO NODE 26 -0.Q70876E-02 

EI EMENT 2£1 

HFAT FLOW INTO NODE 28 

~~H-fEg~-i ~~a8-~-H H~AT FLOW INTO NODE 27 

FLEMENT 26 . HFAT FLOW INTO NODE 31 -0.Q70876E-02 

~~ H-ftgH~+g-~8gt--1~ g :: ~!t-~ f ~~.~~~ HFAT FLOW INTO NODE 30 -0.970R7bF.-O? 

FI EMFNT ?7 

HFA T F L b i~-I Nj1)-N OO-E-~--;;-O .9 7{i"a'r6 E - 0 ;rHFAT FLOW INro NOOE 39 O.Q70876F.-02 HEAT FLOW INTU NODE 38 O.97087bE-O? 

.....IiE.A..LFJ .. .!l.1.'LlNlD_NO'o.E _.3.L~il....9.L_QaI.bF_·:dJ?_. 

FI EtAENT 28 

H FAT FLO W IN TO NO D E 33 - 0 .. 97 087 h F: -.1? 

_.H.F A T FLOW .. I 1\ TU _NODE . ___ 4..0 .. _(t •. 9} Oil7 hF -:.()2 .... _ HFAT FLOw INTO NODE 39 O.97087hF-0? 

HFAT FLOW INrc NODE 32 -U.970876F-02 

_FI EME;IIJT .. 2G ........ __ .. __ ... _______ ... _._ .. __ HFAT FLOw INTU I\OO~ 34 -u.q70~7hE-O~ 

HFAT FLUW INTU NODE 41 0.970876E-02 

HFAT FLUW [NrO NUOE £10 U.97087bF-n~ 

_.HFA T. fL,iJ \IIJ,,!T C .. ~ OGf;_.D _::jJ .• ~LQ.IH.!?f .::l!L. 

FI Et'lFNT 30 

HF"AT FLOW INTO 

H FAT FLOW 11~ f 0 

-HFAT FLm~ Tf~TG 

HFAT FLUW INTG 

FI EMENT 3<' 

HFAT FLOW INTU NUDE 38 -O.97087hE-O? 

HFAT FLOW INro NOD~ £15 O.97087b~-O? -'HFATFLOW ·INTOl'WDE·-·-1l4--0·:'1TOa7bE~(f? - HFAT FLOW INTO NUDE 37 -0.97087bE-0? 

FI EMENT .B 

··HF"A T FL owrr~TO-r.rGT)E" ·-·--39--;;;-U:'1TuOT6f;;;·O·? 

HFAT FLO~ INTO NUD~ qb O.970A7bE-0? 

HFAT FLOW INfO NODE 45 O.970d76F-O? ~~!"'£.!::..0~/IJ.lQ..~ E 38 - () • '1 7 () 8 7 I) f - ~L .. 

FI EMENT 3£1 

HFAT FLOW INTO NUDE 

HFAT FLOW INTU NOUE 

-,:wITT ·Flow-nrro NODr 

HFAT FLOW INTO NUUE 

FI EMENT 36 

HFAT FLOW INTO ~ODE 

HFAT FLOW INTO NUUt 

-HFAT FLOW ·TNTO ·NODE 

HFAT FLUW INTU NUOE 

Fl Hll:NT 37 

£12 -O.970A7hE-O? 

£19 O.9701:176F-0? ·aa "1T~q(OR(hE-O? 

III -0.G70A76F-O? 

·-HFAT FLOW INTO rlOOE -'-"4if''';O~970iT7nF-O? 

133 
-1"" A 1 F- UJI'V 11~ II ",ullt HFAT FLOw HdL I.OUl 

HFAT FLO~ l~l~ N0U~ 

1"1 EMf J~T .58 

~ 1 (·.-{J7l!~ I hI- -o? ~ 0 () • C, 7 u i< 7 r> f- - 'I i' ,,~ - Ii • C; 7 (J "- 7 bE: - \I ? 

HFAT FLDw INTO NODE 45 -0.G7no7hF-il? 

HFAT FLOw INTO NuDt 5<' u.G7687oF-O? 

HFAT FLOvr INTO NODE---5C-(I.G7()il76~~O? 

HFAT FLO~ INTO ~OOE 44 -O.q7067bE-O? 

FI EMFNl 39 

H FAT F[ 0 W-- rr. T 0- l'JODF----Zr~--;;Tf::-G Hi b 7t:'-E-';;O"-

HFAT FLO~ INTO hODE 53 G.q7~87of-O? 

HFAT FLO~ INTO NUUE 52 u.q70R7bF-O? _~_A I._F LO~ IJj T O..-!'! 00 E _~2~,~J!Ji ~6E' -O~_ 

F.LEMF.NT 40 

HFAT FLOW INTO NUDE 47 -O.C,7DB7bE-02 RFAT FLOw INTu NUDE 54 0.970876F-0? 

RFAI FraW-TNTO-NOOF--S3-IT-.-q-Tli/i Tb-r--O;>--

HFAT FLOw IIJTO r,ODE 116 -0.G70876F-02 

FLEMENT 41 

-HF A T F L QW-rNTO-r-·iODE---ifa---;;;-(f:-q7 Ca7bf'..;-tf2--

RFAT FLOw INTO NODE 55 O.q70B7bE-O;> HFAT FLOW INTG NODE 54 O.970876E-O? 

JiF A _L F _L_ Q.I'L_l.NIJLN 0 D-E __ "LL.::.51. .51QJi I b E _~.O .2-_ 

FI EMFNT 4;> 

HFAT FLO~ INTO ~ODE 119 -~.q7Uh7bE-O? 

RFAT FLOw INTU NLDt ~h ~.G7U87hF-n? 

-HF Ar F L ().~-- IN T O-~,ODE---- 5'5--0. 97fJfl7 f', E-;;; ,);:>-

RFAT FLOW IhTG NODE 48 -O.Q7067bE-02 

FI EMFNT 43 ... FAT FI-O~-lNT(fNiYDt----51---(j:-~Tolin;r-li-?-

HFAT FLO~ I~ro NODE 58 n.G7067hF-O? 

HFAT FLOW If',TO t.ODE 57 O.C;7oI'7nF-CI? 

_H.£ ~T. F L.1l.\:'!. Ji'L r C __ 1\ ()_ D.L _____ 5_0_-JL.!..-'?_~9 6 ? _b_t~(i_C_ 

FI EMFNT 

HFAT FLO~ 

HFAT FLOw 

HFAT FLO'.'." 

HFAl FLOI'> 

44 

If. TO I\ODE H,TO NODE 

HTO fiODE 

l,Hll NUDE 

1"1 EMFt,l 115 

HF AT- FL nw--nn-o-NODE----S3---rr:-<;70EifiF-Or 

HFAT FLO~ INTO NODE bO h.q70fl7bF-0;:> HFAT FlO~ INTO NODE 59 h.Cj7087bF-n? _ HF~~ '=-!:-lJ.~_ I flJJQ_.![)DE __ ~? __ -:_~~_~.1g_~ 7 hE~n? __ 

FIE~-1FI\T 4b 

HFAT FLOw II,TO l\ullE 54 -u.Cj70A7bF-u? HFAT FLD~ INTG ~OOf 61 O.G7UH7bF-02 

--HF AT Ftow--TTITU--NO D~----C U- To l171l1l7 6 ~"';1l rHFAT FLO~ INTO NODE 53 -O.~70B7bE-02 

FLE/1ENT 47 

uloiFAT F["OW--J1';nrI'i'Gm:-~s-~u-:-q-r 0 0 1 h E -CT';:-

HFAT FLOw INTO NODE 62 h.970B7bF-O? 

HFAT FLO~ INTO NODE 61 0.970876E-0;> HFAT FLO~ INTO NODE 54 -O.Q70876E-02 

Fl EMENT 

RFAT FLOW 

RFAT FlOW 

-..FAT FLOVj RFAT FLOW 

41l 

INTO 

INTO 

INrO 

INTO 

FI EMFNT 49 

NODE 

NODE 

NODE 

NODE 

56 -0.970876E~02 

63. 0.970B7bE-02 

62 (I.Cj/08/6E-02 55 -0.970876E-0" 

-HFAT~LOW--TNiO\llrt"Ort"D<1:E""""-----'5""'8O--:-""On-. e-rn...--r.-OV8"71 of -1)"- HFAT FLOW INTO NODE 65 O.q70B7bE-O? RFAT FLOW INTO NODE 64 0.970B7bE-0? 

HFAT FLOW INTO NOUE __ ~5~7 __ -~0~.~9_7_0~B_7_b~E_-_O_2_ 

FI EMENT SO 

HFAT FLOW INTO NODE 59 

HFAT FLOW INTO NODE b6 -1WA1-rtow-TNiU-~OOr-6 5 

RFAT FLOW INTO NODE 58 

FLEMENT 51 ""11F1;TrL01f-yNTIrl'm-nD-eE----"6,..,Or-::-,...,o.,.. • .,C;.,...,,.,C,..'8~1 o""'FI:"'-~tl ? 

134 
FI.Ei"FNT '52 

HFAT FLOW INTO NODE 61 ~0.970876E-O? 

~H+ftg-~-+~f-g-~8-g-~~~H-~-i ~t:2~-- HEAT FLOW INTO NODE \ 60 -0.97087bE-02 

El.EMENT S3 

-H-FAT -F CD l~n.J "-r .... 0--r:-N""O""O E~--;--6'i!-::::--"0 -. n9 "7 c1i""a:r-7"6' E-:;-(~ 

HFAT FLOW INTO NODE 69 0.97087of-O? 

HFAT FLOW INTU NODE 68 0.970876E-0? 

HFATE1Q!!.-±NTO NODE 61 -0.970876E-0;:' __ _ 

FI Et-AEIIJT '5S 

t:iFA T FL-O~lNTcr'NOT)~5---;;o:qTo87bE-;;02--

HFAT FLOw [NTU NUDE 72 0.970876E-02 

HFAT FLOW INTO NODE 71 0.970876E-0? 

.l:!EH ... FlP W 1 NLo...lLOJ.lf_f'-!L~_'UJ)J3.I~~O ? ___ _ 

F!. EMEN T 56 

HFAT FLOW I~TO NODE 66 -O.g70S7h~-O? 

HFAT FLOw INTU NODE 73 0 97~87hf-O? 

--HFATFLOW --tNTU-i~UDC--T2-o.q7{fili6E-O';' 

HFAT FLO~ INrc NODE 65 -O.97087bE-0? 

-ELEMENT _ HFAT FLll~" 

HFAT FLOw 

HFAT FLOw ...J:I.f~T fLUw 

EI EMEiIIT 

HFA T FUH~ 

HFAT FlllW 

-HFAT FLO •• HF A r FLll.~ 

i ~fO--j\rlTDt--6T -0. 'flO /:l 7t;}'-';'-i)? 

INTO NUDE 74 0.970876E-0? 

l,HO ,\lODE '3 0.97Url7hE-il? 

jl\j.LU_ N0.QJ; ___ t>.o_:9_,,-QLo. __ dLh f - 0 L_ 

58 

11-1 fa 

INTO 

If~TlJ 

wro 

NUOE bd-O.97u87bf-O~ 

NODE 75 O.970ri7bE-O~ r~tJDE ------]'1 -i.i~-4 7IJ" 7 of";O? 

NODE 07 -u.Y70/:l76E-II? 

Fl.Et~EI,T 59 

-HFAT FLOW INTO-NODE - -69 ';;-0:970B76E-Ofl 

HFAT FLOW INTO NODE 16 0.970~16E-O? 

HFAT FLO~ I~TO NODE 75 O.970d76F-O? 

HFAT FLllw Ij'<JJGi"9lJ.t: ____ ~I3_~_c1X_Q/:lr6_t_;-O,::> 

FIEt-1F.NT 60 

HFAT FLO~ [~TU NODE 70 -O.Q70A7bE-O? 

HFAT t=LU\'i II'.IU NODE 77 0.'l70li7nt-O? 

-HFH FLO\~INTO "NODE--n,--u:'17087bE;';O;> 

HFAT FLO~ INTU NODE 69 -0.97087bE-O? 

FI EMENT 62 

HFAT FLO~ INTO NODE 

HFAT FLOW INTO NODE 

rln r -FLOVvlNTcr-NDUF 

HFAT FLO~ INTO NUDE 

FLEMFNT .,3 -HF AT FLow-rIHO· NOUF--".---.-u-~-'1T(JaTf:iF-O" 

HFAT FLOW INTO NODE 81 O.97087bF'-O? 

HFAT FlOW [NTtl NODE ao O.Q7087bE-O? HF A T F L o~v II'JTtJ NQD.t: ____ 73_~9_._91..0 In bE - O? 

FI EMEIIoT 04 

HFAT FLOW UHO ~;ODE 75 -0. 97 (j,~ lbf'-O;> !-FAT FLOW II,TO NuDt /j2 o • "I 7 i) Ii 1 hE - 0 ;: -HFAT FLm', IrHG NODE -- 81 -0. QT0876F-il? 

HFAT FL(III INTU NODE 74 -0.970!l7bE-I)? 

FI. E t'iF N T 65 

HFAT FLO", r,.TG NODE 76 -0".'17Gil7,:"F-t1? 

135 
I-' ~ Ai,. L I'" j .. Il I'. I" h: c:.~ u.l;/Ut',it"):'·I".: HFAT FLU~ INfC ~uuf 

HFAT FLlH~ 11\1(, hUUl hi.' lJ.G7(lclnF-i!? 

I":; -li.'i7(JI"7br-tl? 

FI f:MFNl 

HFAT FLOI'> 

HFAT FLOW 

-'-MFA T FL 01'0 

HFAT FLOv; 

bh 

f~TO NUO~ '7 -O.G70P7bE-0? 

INTO hGOE 84 O.970~76F-n? II'; T 0--1\ UDI:---13 :r---'lr-. g 7 Il tl7 6t";'O?-' 

INTO NOUE /6 -O.G70B7bE-0? 

FI EMf'NT 67 

-'HFA T F LO,~ 1 rnO-iW1YE'---,-g----;;;u-:,G-, O!lrbT;' (j;iHFAT FLOw INTO NODE db O.G7lJB76E-Il? 

HFAT FLO~ INTO NUDE 85 O.G7087bE-O? 

_lifJI l __ F~~~Jl~LL~(,JDL_7 8_~fL. GI 08Z!,E-t1?~ __ 

FI EMENT 68 

HFAT FLOW INTO NODE 80 -O.G7087bE-02 

HFAT FLOW INTO NODE 87 0.G70876E-0? --HF Arnl)l'rTiJTD1~ODr--8-b--u~GTljb71;r:..-(f~--

HFAT FLOW IhTC NOUE 79 -0.9708761'-07 

FI EMENT 6G 

-'HF A T--Hl)W-T~JT"o--NOO-E---iiT~o_:-q708-1()E-ir2--

HFAT FLOW INTO NODE 88 0.97087bE-07 , HFAT FLO~ INTO NODE 87 O.G7087bE-O? __ ,JiE.A.LEL,OJ'L,I_NIO_f\;O!J.f;-----.a.~_.~li8~.E..?02.. __ 

FLEMFI'IT 70 

HFAT FLO~ INTO NODE 82 -C.G7087hF-0? 

HFAT FLOw INTC NUDE 89 O.97087oE-O? 

---HF ArFLOW IIHC NODr--lra---O: G7 08 7bE-';;O?-

HFAT FLOW INfO NODE 81 -O.G70b7bE-O? 

HEI-IFNT 

HFAT FLOI'> 

HFAT FLOW 

'-HF AT FL Ovi 

HFAT FLO", 

7? 

INTO NODE 8~ -O.G70H7bF-O? 

INrO NODE 91 O.G7087bF-0? 

IN TC T, ODE-:;-(T--O. G7 (j 8T6F - O?-

INTO NODE 83 -D.G70P7~F-O? 

AN-S-'Ht--E-l\"fll-I'IE-i-R-t'f\:"G-~t:-'t"~~'ffo'-~FV-ISTnN-4-. fr-E 2---- ---SOt t.-f? -f-URR INF.9 ---A-uG --t;1 q-,;-~ SWANSON ANALYSIS SYSTE~S. INC. HOUSTON. PENNSYLVA~IA 15342 PHONE (~12)7~6-330~ TWX 51 

THRUST CHAM~ER MODEL 10.1357 SI 8/84 C 

***** NODAL HEAT FLO~ RATES ***** TIME: O.OOOOOE+UQ LOAD STEP: I ITERATION: 10 CU~. ITER.~ NOTE - "REA-rFcaW RA TES aRE PaS ITIVE INTO THE NITDE -----------------------

NCOE, _____ ~H~E~A~T ________________________________ ___ 

85 -O.q7087bE-02 8b -O.lq~17~E-Ol 

87 -0.IQ4115E-Ol --.;r8g----rr~Tq-4T7'5E"";.Olr--------------- --------- -,-----------------------,-- aQ -O.IQ~175E-Ol 

QO -0.lQ~175E-OI Q t -0. qJ Q..§ 7 bE -0 2 _________ _'_ ____ _ 

TOTAL -0.11b505 

*"'* LOAD STEP 1 ITER 10 r.OMPI.FTFO: TTME: -STn-D v 's n TtCnriiVE-"-GT1\CE'-VAlLHO':- 1f:-c;3iHI---AT NODE O:OOOOOOE+CO KUSE= 0 CUM. ITEk.= 

? 2 C RITE R rON = 1"0 0 U 0 0-- ' - , IJ 

END OF INPUT ENCOUNTERED ON FILE27 

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IlSER3: [Ar~SYS. TONY1 SNSTIIL .OUT; III 

USER3: IANSYS.llJlnj SNS14L.OUI: 111 

lISER3: IAtJSVS. TUN Y} SNST tiL .OUT: 111 

77777771 

77777777 

77 

77 

77 

77 

77 

77 

77 

77 

77 

77 

77 

77 

USER3: [ANSYS. TONY) SNST4L.OUT II 

II S E R 3: [A N S Y S .1 0 I. YJ S N S r ilL. U U I 1 1 

US t. k 3: [A N 5 Y S • r u I. YJ S I~ S T 4 L • U u 1 1 1 
********** ANSYS INPUT DATA LISTTNG (FILEIHJ ********** 

12 

v 

1tl 

V 

24 

V 

30 

V 

138 

~6 

V 

4? 

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41l 

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54 

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r,' ;.. I r f! j I~ l i ..... t ••• : :: t :- 1: j~. , t I r I J" I." ~, t .. 

r: 0:: 1. 1I,j (j \; ,I (l 

FRCWFtoiTY TAIiLE.DI'NS JVAT: NUt-'. PuTNTS: ~ 

,TEMPE"klllU.:IE DATA Tn;i"Fr<IITtJhf DATA 

O.OO()OOOOE+OO 1.000000 ~~Oi'./l(}O 1.000/l()0 

-~A 1£ f.I lAC ;> "-'-- --C'OEFF Ii': H';'\TS "Of:"-nFNs V S":. H::..1P FIJIIA TI 01\ 

CO: 1.000000 

PROPFRTY TABLE. DENS JVAT= ? NUM: POTNTS: ? -~ TEi"PFR" TURE----·D·ATA-·-·---·TFiv,PFt-1A TIlRF fJA T A 

O.OOOOOOOE+OO 1.000000 ?300.00n 1.000000 

MATE R lA L 3 CO EF I I r: I F hi T 5 () F f) F Iv S VS ~ .lE..!.1 F' F ,J IJ A Tl 01\ -'-CO' =······T::nOOooa--------··---··-···-··· 

MATERIAL 1 COEFF ICIFNTS OF tllUXY vs~ TEMP FlJlJATIOt\ cn = 0.?30.00,lH) ______ _ 

PROP~RTY TABLE.NUXV ~AT= NU~~ POTNTS= ? 

TEMPEkATURE DATA TF~~PFRATIJRE DATA __ . __ O.J!O.ooOOOE...+Q.L 0 ... 2 3.0o...Q.oJL..---..-.2...'1'<W~Jln .. o.___ fI .2'100000 

MATFRTAL? COEf-FICIENTS OF NtiXY VS~ li"MP FIJlJATIUt\ 

CO = ().~300000 

"PROPERTY TAo1lE";-NUXY- ~A""=--?-"i\iIJM~ POlNTS: ? ,TEMPEf.lATU~E DATA TFMPF~ATIlRF. naTA 

O.nOOODOOE+OO 0.2300000 ?'!oo.onn 0.2'100000 

---VA TFi< n'['---3- --------C01TFTcTEIH:3 'UF",lIiJX V v S: T E I~P F (JI1A TI at- co = 0.3550000 

_F:R_OPFRTY TARLE.I\UXY "'AT: ~ I\U~~ PUHJTS: ? . , TEt.1PFRA TUi<E--"'- DATA- ._-.... TFI.jPF RA TilliE flA TA 

O.OOOOOO~E+OO 0.3~50u()0 ?300.000 0.3~~nooo 

NnDE 

fI: (1 OF 

K C S = 0 >. • v • J .=_ 1 .• ~ q~ 'i 

7 KCS: 0 x.Y.l= 1.?9q5 

O:QOOOOE-O? n~n()nOnE+OG 

(I'.OIlOOOF+OO n'.1I111l00F+{)O 

F T L L "i P lil in S 8 E hi tf' t\ 1\ ODE 1 A h l) illlli> F 7 -. STA'RT \'lITH NODE'- '2" A'ND-rNr:RFMFNT HY" , ..... . 

~FNERATF .4 TOTAL SETS OF NODES ~TTH rN~RFMFNT 7 

SET IS FROM 1 TO 7 IN STFPS OF, 1 

--'GEOf,lFTRY Tr--C·RFMENTS--lI"RC-U. T3T!3F';;O? O. OOOOOE + 0 0 0 ~ 000 OOE+ 00 

~FNFRATF 4 TOTAL SETS OF NODES ~TTH TNr:RF~FNl 7 

SFT IS F~O~ 22 Ie 28' t~ STEPS OF, 1 --. GEOr.<ETRY··TNCRCi,fE1\TS-'A"RE -o·.-snu-ool";;o r o. OOOOOF+OO O. OOOOOE+OO 

NnDE "iO KeS: 0 X,Y.l= 1.3039 n~qOOOOF-r.2 O~OOOOOE+OG 

... NODE' .. ' 'ib ···KC·S=---O---X·;Y;-z=-r:31l3Q-----·0: i1 0 0 0 01" + 0 0 0: 0000 OE + 0 0 

FTLL r:; POINTS BETftEEN NUDE 50 AND NODE 

~~B..~_~ITH NODE 51 AND INCREI-'ENT RY 1 

NnDE "i7 KCS: 0 X.Y.Z= 1.30B6 n:QOOOOE-02 o:onOOOE+OO 

--.:N.:..:n~D~E::.-_.:.:":=.3_~K~C:..::S:.::=:......:O~-'X::..!..... :"Y.:.., !:..Z.::.=_..!.I..::..-=3~0~B:..:;&'--_--.:0~:=_:O:...:..O 0 n OE + 0 0 o'~ Q.~.I!.Q.g..!..O.9 __ _ 

FTLL ~ POINTS BET~EEN NODE 57 AND NODF 6'1 

sTART wITH NODE sa AND INCREMENT AY 

~nDF'"-fi'~IH:s::-n-- X. T • Z= 1.31Inr-·cr~'T01)nnf-;.rr2·-n.!~ 000 OOE + 0 O' 

NODE 70 KCS= 0 x.Y.Z: 1.3140 , , o.oooonF+OO O.OOOOOE+OO 

F TLL -"P(JT1<> I 5 BEIWFE~ODE hI; AN"uNon.-rF-.......,.7-nO· · ___ h •• - •• - ••• '_.-••• - •• -. 

sTART ~ITH NODE bS AND INCREMENT BV 

. ,. ~N-'-'-n=-D.:..:F.---'-7...:1--'-K.:...:C:.:S=-=-O=----.:.:.X~ • ..:..Y..:..-=Z-:-...:1...:.:.o3:..:;>:...:0:.:.';J---0c....=-Q.c.o_O_O_O ____ E •. ~~.o.~~og.E.-t:Q.O_ ... _ 

NnDE 77 KCS= 0 X.V.Z= 1.3202 o:ooonOE+OO O:OOOOOE+OO 

FTLl 5 POINTS BET~EEN NODE 71 AND NonE 77 -:;T'A~'TTl1NODE It! AriD lNCREMFNr-13-,----·r-----·----·-

139 
- .------.-·--.A,~;:.,.::l·- t-1'ir,lNt:bk-i";G··-JtI't,'1t-V:HS---!T'MlTtt~· -wFVISIllN 4.(t 1'2 -SUt:AR--fUHHHIES -t\'UG"'l, tqt:\l---

S i', A N S tJ N A i" A L Y SIS S Y ;j T E /,1 S, HJ C • H II U S l' t) r~. P f' N illS' L \I A i\ I A 1 ~ 3 U 2 PH (j N t (U le J 7 4 h - 3 304 Hu, 5, 

TITLE 10.13tl9 51 tl/84 I 

~  ANSYS ANALYSIS DEFINITION (PREP1) 

'-m;--W-TI TCr=-TFiRmt----c-F1'M B E R MG 0 E [ 

PRINTOUT RE_;iUMEO BY (G""'O"'-'-P __________ _ 

ANALYSIS TYPE: 0 

** *** 

T!=MPER~.ILJRE OFFqET FfiC~ A B SOL UTE Z E flO =~_Q.QJ1 ___ .. _____ .... __ 

REFERENCE TEMPERATURE: 70.000 (TUNIF= 70~OOO) 

4~5~h{!~h-L~-S-ESUlli-T~-"'4-'=-2_-....2---".O--0;<""----,O..--

ISOPAR. STRESS SOLIO. 2-0 

-cli'RRENT-N'OOrr--oiJ'F'-Sn IS Ox 

T~O-OIMENSIONAL STRUCTURE 

Uy 

---- _ ... _--_._._---

--tiAIER'I1C-- 1 cOEFrIC l~u-F--"E""'i4" -'V7i::;. n:i~p -EllUAf ION----------' ------------.-... ---

CO: 6000000. Cl: -2000.000 C2 = o.oonoOOOE.OO 

C 3 = 0 • J 0 0 0 0 0 0 t: l' 0 0 C 4: 0 ~ 000 II I) n 0 E.O 0 

-pjmpERTYTA-aCE~T)(--rrAT=-r--N'UM::-POnJTS'=--2-----

TEMPERATURE DATA TEMPFRATURE DATA 

O.OOOOOOOE+OO 6600000. ?30o.000 2000nOo. 

~A TER TAL '2-'COEFFrcrEiHs OF-EX 

CO: 0.182500uE.O~ Cl = l~uO.noo 

C3: O.~OOOOUOf1'Oa C4: O.UOOOOOOE.OO 

-V3. Tt:,'IP FiJIJA T 1UI\ 

C?: 0 • 0 0 U I) (j 0 0 I: l' 0 0 

--PROPERTY TAffLE',-EX ---~A-'''=--'''''T'NUl.,r:--POTNTS=-7. 

TEMPERATUkE DATA TEMPERATUkE DATA 

0.00000001:1'00 0.1825000E.Otl ?300.00n 0.2tul000E.Ob 

~ATERrAL 3 ---CUEFFTcr~NTS OF EX 

Co: O.llOUOOOEtOa 

PRUPERTY 'ABLE,E~ ~AT= 3 ~UM. POINTS: 2 TEMPE~ATURE .. - -OATA------ --TEMPERATURE DATA 

O.OOOOuuOE.OO 0.170uOOOI:+Od ?300.000 u.1700nOOI:.Od 

t-lATEI<IAL 1 COt:FFICIEfHS OF ALP;(J5. TEI"P felllArIOI\ 

CO: 0 • 3 G 1 () 0 0 0 E - () ~ ---C 1 -: . 0 ~ 1 ? () 0 Ii 0 0 F - 0 El C 2: ,) • 0\) " Il 0 0 0 E l' 0 0 

C3: O.OOOOOOu~+OO eu: O.OOOOOOOE+OO 

PRUPERTY TAHLE,ALPX ~AT: 1 ~UM. POINTS: 2 ---- ----TEi-1PERATURE --UATA----------TEMPER-AnJRE DATA 

O.OOOIlOOOE1'UO 0.3970000E-05 2300.000 0.6730000E-05 

t-IATERIAL 2 COEFFICIENTS OF ALPx vs. TE~iP EfJUAlIUf', -_ .. co :O.7529aOOE'';'0s- -oCT :". O'~-43fnOOOE';'OB- '--C2: ,1.11)0701;OE-10 

C3 : -O.7202000E-l~ C4 = 0.178&000E-17 

PROPERTY Tk~LE.ALPX ~AT= 2 NUM. POINrs= ~4 

------TEMPERA TURE -'-UATA' ---------rEMPERA TURE - "-OA TA 

O.OOOOOOOE.OO 0.7S2900uE-0~ 100.0000 0.dOhnq17E-0~ 

?OO.DOOO 0.d753642E-05 300.0000 0.q~70?I~E-05 

400.0000 O.10u7819E-04 ~oo.OOOO 0.1144~~AE-n4 ------f)° (J .-000 0----U-.J7zr=-Q1'l3F.;.uu,ITo-;:-IT 000 . -----0 .1 ,IJ A q q ~ E - 0 II 

1'100.0000 O.14S~43~E-U~ qOO.POOO 0.1~~~'~4E-04 

1000.OCO 0.16566001'-04 1100~ono 0.1/~~an~E-04 

1?00.000 0.1854779E-OU 1300~noo 0.lq~?340E-04 

---- --raoo. 000----,) ~70S'OT? lE'"04--T~00-':-00 0 ·--o.j) llJq~ARE-O 4 I~OO.QOO Q.~2S2634E-04 1100~000 0.2~hl~R?E-04 

IROO.OOO 0.2U791d5E-04 1900.000 O.~hORh~'E-OU 

'000.000 0.27S3500E-01J "oo.ono 0.~q17K~AE-04 

· .. ·??oo.OOO ---- 0.3TOh163E'-oa --;;r,oo.ooo 0.33;J33n7E-04 

MATERIAL 3 COEFFlr.JENTS OF ALP)( VS. T"'Mt> FiJlJATlOi\i 

Co: o.qaOUOOOE-OS ---.------- ----

PROPFRTY TA6LE.ALPX ~AT~ 3 NUM~ POTNTS= ? 

,TEMPERATURE DATA TFi4PFRATIJI-IF DATA 

0.0 OullOO OE .00 _9. 'id o(j9n OF. :-£lS?30o. 011 0 0.91'1000 nilE -() 5 

140 
NODE 7R KCS= 0 X,Y,Z= 1.327? o:q~~~;~-;;? O~~-;;-I;~;E~)O- .. -.-- .. 

NODE A4 KCS= 0 X,Y,Z= 1.3~7? O:OOOOOE+OO O:OOOOOE+OO 

FTLL ~ POINTS BET~EEN NODE 7R AND NOnF A4 :-.S.l...A.RL...~lTH NQP_E-__ B AND INCREt-iENT BY t _____ . ___ ._._. _____ ._. ____ _ 

NODE 8S KCS= 0 X.Y,Z: 1.3350 0:QOOOOE-02 O:OOOOOE+OO 

FTLL ~ POINTS BET~EEN NODE 85 AND NOOF 91 

START WITH NODE 86 AND INCRFMENT BV t 

-VAIE"lHA"CNUP"BE"R SET TO 1 

FI.EMENT 1 2 ~ 8 

, r.F~~~~ l~-l-i3T Al,~-HS S9~pgL-EfENrs--\H..I!LN.ll.D-Cl.f'!CBf-M.ffllI ._OF __ .1 . 

. Nllt-IAER OF ELEMENTS: 6 

-~F.~~~Ag_l__.~gIAL.b.s-~~s-s~h§-L5~-E.NI~L\\lJn: NilDf. lNCRfMFNT OF 

NIIt"RER OF ELEtJENTS: Id 

r.FNF.~ATE 6 TOTAL SETS OF ELE~F.NTS WITH NonE TNCRFMFNT OF 

SET IS lq TG . lq IN STEPS OF 1 --".ili M A ER- m='-E'i..-E t,fEN n-: --'--2'ij' -;-.-.----------.---------' .- ---- .. ---- .. 

r.FNERATF ~ TOTAL SETS OF EL~~ENTS WITH NonE fNCRFMFNT OF 

SFT IS 19 TO dIJ 1~ STEPS OF 1 . t\'1W,RF.11 OF ELEMENTS: "30 . 

~ATERIAL NUMHE~ SET TG 3 

FIE MJ' NT. _.3.L. __ ..... .. I! IJ ____.51 _____ 2fr........ __ IJ :~ 

r.FNEiiATE. 6 TOTAL SETS OF ELE;"l-:NTS WITH NIJOF TNC'IFMFNT OF 

SET IS ~7 TO 37 [N STEPS OF 1 

NIIMAER [JF ELEfiEN fS= 42 

r.FNERATF h TOTAL SETS Of ELE1'II"NTS ,vlTH NOliF IfH':RFi~FNT OF S~T TS ~7 TG 4~ I~ STEPS OF 1 

NI!MHF~ Of ELEMEi~rs;:: .... JL __ 

NTTTER: -10 NPQINT= 10 NPQST= 10 

ROUNnANV CONDITIUNS STEPPED DUE TO NEGATTVF NITTEI1 

AI [-PIHNT CONTROlS"'RESET'To--"---U)-'-'---

AI L POST DATA FILE CCI\TROLS RESEr TO 10 

-KTEMP= -'--'1---10 

TFMPERATU~ES FROM LOAD STEP: LTF~ATTnN= 

. STF.PSOUNDAin"·CCThIJTTTUWl\EY= -..----.--.. -----

~FACTION FO~CE KEY: 1 

10 

__ 'J. 

7 

-g-PECIF'TFcr 1nSp-uy-f'Rcr~--rroOF' ',0" NnnF' ---qrTN -'STFPS OF .-. T 

VALUES= O.OOOOOE+OO O.OOOOOE+OO AODITTtJNAL DnFS= 

SFT_~_!!:1U~. COUPLED II\ODE SET SIZE Til 

r.FNE~ATE ADDITIONAL I\CDES ON SET: 

STARTING NODE: 1 ENDING ~COF: 

1 nTRFcrrON: IIV A~ N(1IlF INC.: 

--r.nllPLFO "SET=---r-u-nrE"CTT01\: lIY'---TnTAL"NonFS: 

NnDES: 1 8 IS 22 ?q ~h a~ 

71 78 85 

rJ ,U/III.UM-CQUPLFTI 'SET'l\uMBER:' ·-1------·-...... ---·· - -.. 

DATA CHFCKF.D - NO FATAL ERRORS FOUND. , rHF.CK OUTPUT Fa~ PCS5IBLE ~ARNING ~ESSAGFS _ _ . -.- --_.------._----- .. - .. ---.----~- _ .. -_ .. _--..... ---_._-

ANALYSTS DATA ~RrTTE~ ON FILE27 ( i'4h LT.NFS) 

AI L CURRfNT PREP7 IJATA \,;lITTFI\ TG FILflh -. FOR POSSr'1LE RF.3UIVE FRC:-cTHTS 'FnTi,T 

141 

n 

'i0 

7 

57 64 
----;t{I~-H-ri£H~~T ~OPP~TEW-~n-t;-~-.-l~~-N~F~F-~~~~~O~\J-~~~~hIS------------------.-----. ---- SYSTEMS. INC. NOk THE CORPORATION SIlPPLvtNr. THE CO~PUrER 

FACILITIES FOR THIS ANALYSIg ASSUMF ANY RFSPONSIBILITY FOR 

------.;.b~~~fTA.;.~~k5-~-~-f{~U~~~H_~~__s-~~~~E~~~! 6~;Hn~iJ~~\~~~~~ T~I$--- ._m ______ . _____ . _________ _ 

OWN RESULTS. 

------.... SWAI"fSOf.fANALYS ISS Y5TUTS":TN1':-:--r;<;-t:NhF'AVrlRTNt.10Mll< E THE -- ---- --------------------------- ANSYS PROGRA~ AS COMPLETE. ACCURATE. ANn FASY TO USE AS 

POSSIBLE. SUGGESTIONS AND r.OM~FNTS ARE WFICOMEO. ANY 

---~~~-~8~-h·~~fg~~6E~~_n5:~~-tr7~-~-y T~~nB~~¥~~T~~~n~T~~~t15N:----------------------

__ . ____ . ____ ~ __ *~.1!.__.AN~t...LS_I.S_QPT to~.*~~ __ ~ ________ ._ 

VALUF 

-~~fJ~l_~I~-So1!-~·h-sO-F·FR1-EOO·;-:~1.a~,.,.K~~-;-;Y---7V----------.----------.------.------.- -----------------... ------ FLEMFNT CONSTA~T TABLE h 

REACT TON FORCE KEY ••••• t 

MASTFR OOF RE'O KEY ••• '. _ 1 -MA TER IAT.. --T .fBeE -ENTRiES-. --- :-:__..----;;-a----

REFERFNCE TEMPERATURE . . .. 70.00 __ !INJF_9.RM LFM~_EBALUI!E_._. __ .• __ .. --,, ___ 7!1.e.OO 

TYPF STTF DEseR IPTIll~~ KFY OPTTC~_S 

? ~ 4 '5 b 

- , I SOPAR. - STfi"ESS-SUI:rrr; ~-O fl II 200 

Nllt-IHFR OF ELE .... EN r T vFE 5= 

NO~ ***** TABLE OF ELE~~~r REAl CONSTANT~ ***_. 

NIIMAFR OF ~EAL CUillSrANT SETS= 11 

142 

7 A q 

000 o 

HJlll P 
~~u* ELEi'iENT-DtfTi~rn-ON~f***"'._------­

FI EMHT NODES MAT TYPF CLA5S ELEf'.lErJT REAL CQI\STAi'JTS 

-----·~1---2----~ir 1 -.. -----.-... 'f --, 0 

~ 3 10 9 2 1 I 0 ~ 4 11 10 3 1 1 0 

4 5 12 11 4 1 1 0 .,----- 6--1 r--f2---S- -------C-- 1 0 

r. 7 14 13 6 1 1 0 

7 q 16 15 8 1 1 0 

R to 17 16 9 1 1 0 ---- -- Q---n--Uj---l 7--1o-C 1 0 

10 12 19 18 II 1 1 0 

11 13 20 19 12 1 1 0 

: ~ ----: ~--H-----~ 2---} §~----- -- +-----! 2 

14 17 24 23 16 1 1 0 

IS 18 25 24 17 1 1 0 _ 110.__ __19 ___ <'6 __ ..25_-.J_8*-_________ 1 1 0 

17 20 ?7 26 19 1 ,-- () 

III ?1 21'1 2-( 20 I 1 0 

19 23 30 29 22 ;:> I 0 

20 2a 31 30 23 2 1 () ?1 ,--- 25- 32------3C--2/j------------ ------, <' 1 0 

?? 26 33 32 25 2 1 0 ?~ 27 34 33 26 2 1 0 

?IJ 28 35 34 2,--;.7,--_____ 2 1 0 ----?"--30- 3(--- 31;---29 -------------- 2 ,- (, 

?6 31 38 37 30 2 1 U ~7 32 3G 38 31 2 1 0 

2R 33 ao 3G 32 ? 1 0 --- 20:;-- -]<r-41-- 40----3:r----------- ? 1 0 

30 35 42 /jl 34 ? 1 0 

31 37 44 /j3 36 ? 1 r. 3~ 36 45 44 37 ;:> 1 0 ,~ 3 q 4 I> Ii '5 38- ? 1 () 

31J 40 47 46 39 ? I 0 3~ 41 4M 47 40 2 1 0 

3" 4t> 49 48 41 2 1 0 

37 1.111"-- 51---- s(J----1B------------- :5 1 0 

3A 45 52 51 /j4 3 1 0 

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40 47 S/j 53 /j6 3 1 0 

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4~ IJY ~6 55 48 3 1 0 

4" 51 58 57 5 0 :~ 1 0 

41.1 52 59 'i8 51 :5 1 0 ---u-.:;--sr--6-U---'S-q---S2 ---- ------- -" 3--- 1 0 

4f, 54 61 60 53 3 1 0 

47 55 62 61 5/j 3 1 0 

41'1 56 63 62 55 3 1 0 ---~-·£lG·--·~8--bS--t,.1l--SI .--.------- 3---- l' 0 

50 59 no 65 58 3 1 0 

51 bO 67" 66 59 3 1 (J 

5? hI h8 67 00 3 , 0 --~,_--o;r_n..____68 61----------- T------ ,--- - 0 

54 63 70 69 62 3 1 0 

55 65 72 71 64 3 1 0 

51. 6b 73 72 65,. 3 1 0 51 ... / It! 13 66 -r---T---O------------ ------------

58 b8 75 7q 67 3 1 0 

Sq b9 76 75 68 3 1 0 

60 70 17 76 69 3 1 0 ---6,---- 7~(q 78 71 ----T--- , ---. - 0 u_ 

6? 73 80 7'1 72 3 1 0 n~ 74 Al 80 73 3 1 0 

h4 75 82 81 7q 3 1 0 --fi-.;-~R_r____ll_e_75 -- roo- ,. r 0 ---

hh 77 8q 83 76 3 , 0 

67 79 86 1:15 78 3 1 0 

h8 80 1'17 86 7q 3 , 0 

'--., q----g 1---------gr-a-r-8 Q r ---- ,- -- o· 

70 8<' 89 88 81 3 1 0 

71 83 90 89 82 3 1 0 __ .;.-7?'-. _84 91 90 83 ,3 1 .0 ______________________________ ._, __________ ,_, _

--r-Nff-r,~-tJf(-~E~m..~f'/tEIJT:'l Ft.~ EL~MF~~-·-----------_____r:p_::_'--n-:t'40---TtME-=-t'O ~-t/j'5'5i 

MFI"URY I= 1001 t-'E/\IURY 11= tA~ TOTAl.: l1A~ MEMORY AVAILABLE= 340000 

MAXIMUt-l NOiJE NUMBER FOR AVAlLAflLE AUXTLTARY ivlFMOR't ST7F= 169a<l8 

NUMAFR OF ELEMENTS : 72 MAXIMUM NOnF. NIJMf1FR USFD = 91 

143 
lJOOI:--- ·--x--·····--··· --.'y------ ... ---.- T H X 'i 

(Ot< R) COR Ti-tETA) (OR RTl 

, 1.294S n.q~O"OF-O? D.OOOOOF+OO ~--- --·-'-1 ~ ;>q45 -0 :7S-()oOr-Il~--- 0;' OOO·nOF+OO-- H 

~ 1.2q45 0.600001:-0? n~OOOOOF+OO 

4 1.;>945 O.45000E~0? O.OOOOOF+OO 

5 1.2945 0.30000F-0? 0 OOOOOF+OO 

fl- -·'·---1 .2CJ45--·----0-~·)C;oO·OT;;.-0~-·--- n:OIlOOOF+OO· 

7 1.2445 n.OOOOOE+nO O.ODOOOF+OD 

R 1.2958 0.90000F-0? o.OOOOOF+OO 

q 1.2958 0.7S000F-0? O.OOOOOF+OO 

10 f. 2 g5!!------O-. hOO-O Oi;-:.O?----O.O 00 (i OF + 0 Ii 

11 1.2qS8 0.45000F-0? O~OOOOOF+OO 

I? 1.2958 0.30000F-0? o.OOnOOF+OO 

1 ~ .. ___ 1._2.95IL. __ . __ ._O .• J.5.!tOJIF- .-O.? ___ 0 .. 0.000 OF + 0 0 __ _ 14 \.?YSe IJ.OOOOoE+On fJ.nn(JOOF+Of, 

l'i 1.2'172 f).ilOOuOE-O? fI.OfIOOIlF+OO 

1 hi. 2 9 7 2 0 • 7 'i 0 0 0 F - 0 ? (J • n () 0 0 0 F + I) 0 

1 7 1 • 2 q 7 2 (I • I- 0 (, n II F - J) ? 0 l' Ii [) () 1111 F + 0 n 

lA 1 ~?q72-----(5-.-4cHfOOF';0?--· o.nnOMIF+O(,· 

lq 1.2972 O.!OOOOE-O? fI.OOOOOF+OO 

20 1.2972 0.15000E-0? O.OOOOoF+On 

21 1.2972 O.OOOOOF+OO OOOOOOF+OO 

n-·· L 2 q 8 S--·----\I-,;-·g 0·0-0 () F-- 02----0 ~ 0 II 0 (,OF + 0 0 

23 1.?9~5 O.7S000F-n? n.DOOOOF+OO 

?4 1.~9b5 lI.hOOOOF-O? n.nnolln~+on 

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?7 1.298'S O.15000F-O? O.OOOOtlF+OO 

2A 1.2985 O.OOODOF+nO o.OOOOOF+OO 

?q 1.?990 0.90000E-O? D.nOnOOF+OO ~ 0 I • 2 q 9 0 --.- .- .-.- 0 • 7 '3 0 ,)() F - () ? () • () n ,1 ,HI F + 0 0 

~ 1 1 • ~ 9 9 0 (, • h () 0 I) 0 F - 0 ? Ii • 0 0 () (} Ill' -+ 0 n ~? 1.2qq() O.4"iOOOF-IJ? O.ilOOnOf'+oo 3~ 1.2490 0.30000F-n? 0.II0000F+\l0 "'14 r~ ?9CJO --.---.. -.- -0. I 5n-0 Or-.;.O? 0.0000 OF + 0 0 

~"i t.2990 O.OOOOOE+OO o.nnooo~+oo 

"'Ih 1.2995 O.ClOOOOE-O? ().OOOOOF~OIl 

~7 1.?995 O.7S000~-n2 u.OOOnOF+OO ~Il- .. - 1.2995-··------·-o.hOn·\iOf-O?n.010nOnF+00 ~ G 1 .;> 9 ~ 'i (I • tJ 'i 0 n () F - O? C. • 01 0 0 II n F .. 0 fJ 

aO 1.2995 O.3nOOOF-O? n.OilounF+OO 

41 1.2995 0.15000F-0? o.ononOF+OO 

--·-·--4 ?---.-.- r~ 29 9'5-------O-:-01l 01l·0Ei- 0 (1.-- -0-. onrro 0 F + 0 n--

a3 1.3000 0.QOOOOE-02 o.onuOOF+OO 

a4 1.3000 0.75000F-fJ? o.oonooF+OO 

45 1.3000 0.nOOOO£-02 O.OOOOOF+on -·---41i-----,.:3lf,)() (). ZJ'5OlJOr;;l1;r--o-:nnOOliF+OO-·· 

47 1.3000 O.30000F-0? o.nonnOF+OO 

Ull 1.3000 O.l"iOOOF-O? n .. ooouOF+OO 

4G 1.3000 1i.00DOOE+no 0.000001'+00 

"i 0---·--_· r::31l3 Cj --U--~ q () 0 0 () F .. rr;r--tl .. 0000 OF '+ 0 0--

"it 1.3039 0.75000E';'02 o.onoonF+on 

S2 1.3039 0.60000E-02 OpOOOOOF+OO "i] 1.303Q 0.4~000E-O? O.OOOOOF+OO 

----.;4 1.303Q 1I.30CullE-tl? u •• l1ill1l!TF+orr55 1.303Q 0.15000E-0? O.OOOOOF+OO 

"ih 1.303~ O.OOOOOE+OO O.OOOOOF+OO 

"i7 1.3086 O.GOOOOF-O? o.nOOOOF+OO --. -, ,,----·1-~·:31Hl6 n-;;:-T51JlJllF.;;;O ;:r---II.: 0 0 0 OOF+ 0 If----

"iQ 1.3086 O.nOOOOE-O? O.OOOOOF+OO 

&0 1.308~ 0.45000E-02 OpOODonF+OO 61 1.3086 0.30000F-02 O~OOOOOF+OO ---6,,----1:-311'80 -u-:-pmOllF;;-o·;r-- rr-,::IT01TOOF+on-·-

h3 1.3086 0.00000[+00 o.nooooF+OO 

kU 1.3140 0.90000E-0? O.OOOOOF+OO &~ 1.3140 0.7~OOOF-0? O.OOOOOF+OO 

---fj" 1 • 3TZI 0 0 • 6-mrmrr-rr;r-rr;. no 0 0 0 F + 0 II . h7 1.3140 0.45000E-0? O~OOOOOF+OO 

bA 1.3140 0.30000E-02 O.OOOOOF+OO 

h'1 1.3140 0.15000E-02 O.OOOOOF+OO --70--- r-:::31Lio u :0-0-00 on 00----n . 00000 no 0--· 

--7'1 1~2'(r2 ,,;-4HllltlnF-1tI' tl~fti"tit(tF"o-rt·--

72 1.3202 0.7S000E-02 O.OOOUOF+on 

73 1.320~ O.hOOOOF-O? o.nOOOOF+OO 

7a 1.3202 o.u~nOOE-O? O.OOoooF+on 

7"i 1.3202 0.30000E-02 O.OOOOOF+OO 

7h 1.3202 0.15000F-02 o.noOOOF+OO 

77 1 320 2 ~O 0 0 0 F + 0 () 0 ~ (HHI () 0 F + 0 0 ---'11 t ::ri1:? 0 . ':I 0-000"-;:;-0;:' \) ~ 0 oif6("-F~-1l Ii 

7'1 1.3272 0.75000E-0? O.OOOOOF+OO 

80 1.3272 0.60000E-02 O~OOOOOF+OO 

---~ 1----t·:i~H- g : ~'a-g~g-f::-~~--g:g ~~-a g ~t g:~ 8~ 1.3272 0.15000F-02 O.OOOOOF+OO 

84 1.3272 O.OOOOOE+OO o.onOonF+OO 

Il"i t .3350 o.qOOOOE~O? ~nOOOF+OO 

---a·.,----C33-5"0 0 • 7 S 0 0 0 F - o;r II .. lill (Hrll~+O 0 

87 1.3350 0.60000E-02 OpOOOOOF+OO AA 1.3350 0.U5000E-0? O .. OOOOOF+on 

-·--~ri·· ---- -}:Hgg-·----g:rg·g-gg~~g~--g:ggggg~:2g 91 t .33S0 O.OOOOOF+()O o.oonOOF+O') 

_XM rN.;; __ .1 .• ?.95 _____ ._l.(~~1L=~l3~. __ \'t1t.N::_!).onOOF+no ~i'l1\ X= .0 • 90 OOE.,",02 ... _]./<JIN= It. 00 O.OE+O.O __ ZMA.~ 

TNTEGER STORAGE REQUIREMENTS FOR NOOE INP(IT CP= la.970 TI~E= 10.14771 ~FMORY 1= 0 to'Ef'iCRY II=·, , C:;Uk TOTAl = <ian MEMORY AVAILAAlE= 340()OQ __ . __ ~A X Tf.'.Uf.'_J\jODE -"tU!1...e .. EiLE.OJLAlIAl.L...4f.1.LE.JIJX II IIIR Y __ MFt~OR f S T IF: 1133.3 3 __ .___ ______ ._ . ________ . ___ . ____ _ 

144 
-----".-'-- A:';!'lT~ - t-l\blr,t:.tf-Hd, ,11'lAL1:>1~ :>v';ii-~· f.F-vl:illJl; /J.I: E? SCI ,~? TtJto".li!~_::' tlu\, 1.1YII} 

S f\ Al-J S [) h A I', ~. L Y :; I!:: S Y 5 It, t·, S. 11. r. • rHlU:, r n i, • ,.> F fJ i; S Y L v A 1\ 1 A } 5 ~,,? ;.> H Iii J r (" 1 ? 1 7 " h - .5 3 \l 4 r ;\ X '" 1 (J - b 'i l) - II 

IRUST CHA~~ik MUOEL 

***** ***** 

,TERIAL 

'X PROPE~TY TABLE (LINFAR INTfRPOLATJONl 

M·P f:X TE 1"1 P E)( TF M P 

"}lr:lI-lJ-:-6E on 0 E"+ 0 7 2 J"Oo.ll-~(}IHf(fF +" () 7 ----------

fERJ AL ? 

X PROPERTY TABLE (lINEAR I~TFRPQlATTONl 

MP EX TEMP EX TFMP 

-; -0=0- -0':-1825 oe: +'lfEf--2"3() (1:- o---o.?Tol OF+oi\--'-'-

,TER 1 AL 3 

'X PROPERTY TABLE (LINEAR I~TFRPOLATTnNl 

MP Ex TEMP F. TFMP 

-n-:: 0-- o-.lrOOon 6S--23-olf:0----O-:-1Too-on:OA 

TERlAL 

LPX PRUPERTY TABLE (lII~f:AR lilTE;:<PULATTOfn MP ALP X TE~P ALPX TFMP 

-0.0 O. 3CJ70 OE'';O 5----2300-::0 --(i-::-f, 7 ::IOoF -oc; 

TERIAl 2 

LPJ( PROFfidY TAfiLE CUI\F.AR II.HRPOUTTIW) 

Mf' ALPX TE"'''' ALPX Trt.1r' 

--0,.0" 

SOO~o 

000.0 

500~0 

'011 0-.0 

O. 7S ?CJOf:- O-S--- 10 0."Ir- (J.ROb I OF -Oli 

O.11449E-OU bOO.O 0.1?qbOF-04 O.lo56bE-04 1100.0 0.17ShaF-oa 

0.?149bE-Oa 1600.0 O.~?5?hF-Oa 

O. n 535E" -0 .:r-- 21 OU-:O--' '0;;>9] 7 9F - () a-

,TERIAL 3 

,LPX PROPERTY TABLE (LINEAR INTERPOLATJON) 

;>00.0 

700.0 

I~OO.() 

1700.0 

~;>OO.1i 

:MP AlPX TEMP ALPX TFMP 

--0-:11' o."';;OUO"£;;' U .... 5--2,-,3""01"7(') ."o,---,(n-) -.' l:rI'flmO F ;;;'Oc;'------"'---------

,TERI AL 

--------------------------- IE~S P~OPERTY TABLE (LI~EAR I~TERPDLATTON) 

:MP DE~S JEMP DENS TFMP ~ . 

IO.14M£: 

TEMP 

EX fEMP 

TEMP 

ALP X TE'MF ALPX 

ALPx TEfoiIP ALP X 

• O .. 8753hE'-05 300.0 0.9<;71);:>E-05 (),.13119()~-(j4 bOO~O O.14S;>Qf-Oq 0 .. 1~548E-oa 1300.0 0.195?3E-Oa Op?~blbE-OQ tson.o 0.?479?F-OQ 0.310b2E'-OQ ;>~()(J.() fI.33~33E-04 

ALP X TEMP ALPll 

DENS TEMF DENS 

, () ~ () 1 • \l U 0 Il 2300 • 0 1 • (rrrO"tTIl7r II----- ----,-------------------------

,TERIAL ? 

" lENS PROPERTY TABLE (LINEAR INTERPOLATIONl 

:MP DENS TEMP DENS TEMP DENS TEMF DENS ".~' , f 

-lI: -U---l--:-rnrn-o"'-o--

ITERlAL 3 

,..--------'------------------------------"--------

lENS PROPERTY TAHLE (LINEAR J~TERPOLATJON) 

:MP DENS TEMP DENS TFMP DENS TEMP DENS 

~~(Jr~~()r-~1-.~~O~----~2~ .. 3~(~)O~.~\I--~l~:~n~o~()~()·--------

145 

51 h/84 Cf': 

TEtvP Ex 

TEf'J,P EX 

nrvp EX 

ALPX 

TEI"P ALPX 

aoo.O 0.10478E900.0 u.15552tlQno.o 0.20:'01f:-

1900.0 O.2hOllor:-

TE"'P ALPX 

TEtvP DENS 

TEt-'P DENS 

TEMP DENS 
~ATEFlIAL 

NUX' PROPERT' TABLE (LINEAR INrE~POLATrON) 

TEr,lP NUlIY TEMP NUll' TFMP NUXY TEMP r~ux' TEt-IP 

----lr:O.---n::?-30ll0 . 2 3'0 0 • 0 0 • ;J 3 0 0 \) ._--_ .. _--_._ ..... _- ...... ---.---

MATERIAL 2 

--------------------'-------------_ .. _._._-_._ ..... _----_._-------_._--_. 

NUXY PROPERTY TABLE (LINEAR INTERPOLATION) TF.MP NUXY TEMP NUXY TFMP NUX' TEMP NuXY TEr,··p 

---·O·~O-o·~3"(nr.,,-o ----"2·)"(fO. 0 O. '·Hriyo----------····--·· ... 

MATERIAL 3 

NUXY PROPERrY TABLE (LINEAR INTERPOLATION) TEMP NUXY TEMP NUXY TF~P NUXY TEMP NUX' TEMP 

----cCo·-oJS"SO"'i'f----2-300. 0----0:-3.,-';-00----·--.... · --.------. 

~AXIMUM MATERIAL NU~8ER= 3 

-_.-_._._-----------------_ .. _--_ .. _-.. --- -. __ ... -._-- ..... --... _ .... _.--.. _---_. -.. _-------_._ ... _ ... --._ .. _ .. __ ._- ••••• COLPLEO DEG. OF FR. nFFINTTTnNS • __ •• 

SF T. __ J1. (). t'_. __ . __ ... _~L!~BE_fL. ___ ._S'_'O.~jELf.!LJJ(JlJ!' s 

U'( 13 1 

71 

II 

7A 

l5 

85 

• - Ii-;i .-i .-'-f! A' 51 tR "[JE'GR'fES'-OF "F FlF'!=' iHl M -. It. It * * 

NOOF DEGREES CF FREEDOI-I I rST 

NIIMRER OF dPfCIFIEO /1IASTER 0.0.1".= 0 

TnT .l L N I j j~ H f' R 0 F j.l A :: T t " D. J • r • = 0 

29 36 50 57 

'TrHEGFR STORAGE REGL'IREMErITS'-TOR f\l,ATERTALS. FTr:. TNPtlT CP= 17./HO TTI-,F= 10.t4Q3'l 

II F 1-1 [) R ¥ 1= ? 1 7 2 /11 E t' 0 In II = 0 TO TA I = .:> 1 l? r~ t: H 0 k Y II V II I L A H '- E :: 3 £! 000 0 

••• LOAf) :HEP 

NTTTEk= -111 NPRIhT= 10 I~POST= 1 \) 

__ .AJ.L .f~ Ii'ir _ COt,:..TJ!P_L"S RE..i~J_-'-tJ ___ J..9 ____ ._. ___ . 

AI L POST OA r A FILE ceNTROlS RESET TO 10 

. ~r:_E ~= o. 00 U 0 P ..1":+ O..'L_.Q • .o _<1.9. QO E.!.Q.Q . .Q...,iLo ° D.'! E+ on 

n~EGA= O.OOOOOE+uO o.ooonOE.OO O.OOOO"F.on 

_12.~_~.GA= __ (J. 0_0 ~_Q.f: .uQ . ....2~oo Of. + 0 0 o. 00 o_Q_of~~~_ ._ .. __ ._.,_ . 

~~LOC= o.onOOOE.OO O.OOOOOE+OO O.OOOOOF.no 

r:~~f.G.A=_.Q., ~Qo...QQ~Q.~O a 0 OOE+OIl...O: ooooC!f +O~) . 

nnME~A= O.OOOOOE+OO o.ouonOE+OO O.OOOOOF.on 

KTEt-AP= I 10 _. --_. __ .. -. ---~---------.-. 

TFMPERATURES FROM LOAD STEP= ITFRATTON: I 

PI ASTIr:TTV r.Cfl;VfRGEr~CE CRITt'ldor.=, D.lll00 "~RFFP- nPTTi-AT7ATIITN' CRTTEWrA= .... ----·O:pTI)O 

lARGE DFFL. CGNvf:RGEI\CE CRITERIA= 0.001000 

nTSPLA~EMENT Lr~IT= O.OOOOO~.OO 

KFY TO TERMINATE RUN IF NU CUNVERGENCE= o • . _- ~ -------- - -----_ .. _---_ .. ---------_ .. _------ -_. 

RFACTION FO~CE KEY: I . II N I F (J R M T E I'A P f R II r u R E = 7'U • 00 () ( Tf-l F F = 7o.onol .-•• -- •••• ¥_ ••••• -

146 

10 
·-_·-·· __ ·_--t\flt&V-:;-···"€I'ffi-I Nt-E-ft"tf'ltf-A"ni\"c'f"S"f S--t;\"5Tl"-r--j(F v I g T HN- II. (1'--E 2' ·--.--set a ~ -T fJ~H TNF S' -'-A UG ·-t,··-t'f!tt--

SWANSON ANALYSIS ~YSrEMS. INr_ ~OU9TON. PFNNSYLVA~IA 15342 PHONE (412]74b-3304 T~X ~ 

THRUST CHAMBER MODEL 

. __ LflALS1..E.P N utA H£..!CRc.....=='-----'~ __________ _ 

*** LOAD OPTIONS SU~~ARY 

TT M E = 0: 0 000 0 ~~+~O~O:-....~N!.-'I...!T...!T-"E,""R,-=:,-_-=-:...!l,-,O,--__ 

ArEL : O.OOOOOE+OO 0.60000E900 o.nOOOOF+DO O~EGA = 0 • .10000E ... 00 O.OOOOOE ... OO O.OOOOOF+OO 

nOMEGA: 0.OOOOOE900 O.OOOOOE+OO O.OOOOOF+no 

10~lU99 51 8/8£1 

-r.-Gl.:-aC"" --::'If .-000 OOE +-ITOlJ.IJlTlTllllF+lllJO: lr(f<T(lOF+-t1lr .--------.----.. --------.-----.--.. - .- -.-.-. -_ .. -.--. -- ... ---.--

rGOMEG: O.OOOOOE ... OO 0.00000£ ... 00 O.OOOOOE+OO 

nrGOME= O.OOOOOE ... OO O.OOOOOE ... OO .O.OOODOF"'OO 

TFMPERATURE (KTEMP) LOAD STEP= I ITERATION: ._1_0. ,'-1 fF"FN ESS'RF1JS E--'Kn'-TKUS~o --..... -.----.--

IINIFORM TEMPERATURE: 70.000 Crr<FF.: 70:0110) 

-PRJ N T'FLEM8-!irC RCrr-Al'ifDKEAi:TTOW-~ ORCF 5--1 KR 1": \') 

IOA05 STEPPED TO FINAL VALUES FOR ALL TTFRATION5 (KAr: 1] 

-~~·:~~J5-p-¥~~-~~~t7-cR-I~@·Hl~~-N= v. ?o:g*OJL._-.. -.. -.... --.---.-.---... -.~ -... -.-----.. -.-.---.---.--.-- -.--.--.--. --.'-- '--

J ARGE OFFL. CUNVERG. CRITERION: 0.001000 

nTSPLArEMiNT LI~IT; 0.00000E900 

KFY TO TERMlNATE RUN IF NO CU~VFRr,F.NCF.: O. 'ODI:="'-"--O -·rSyM=---t---··-----· ..... -.-----..... -

NPRINI= 10 NPCST: 10 R~A~rtGN PRTNT FREQ: 10 

_or.SP._ P(jST_JlATA_f...RJ:G.=. ___ 1Q ___ RJ-:_~c:r. YOSTOA}A FRE;J: 10 

OISPLACEMENT PHINT F~EQUENCJES 

FREq NSTRT NSTGF Nl~C 

1 II __ 1_ .. 3?Q Y 0 ________ 1._ ._ 

FLEMFNT PRTNT ANU peST DATA FREQUF~CIFS 

TYPE STIFF ~TRESS FORCf STRFS~ nATA FORCF 

NODF 

7 

14 

;l1 

?A 

30; 

a;> 

4G 

5", 

'1;";\ 

70 

77 

84 

------q I 

~U. PRINT PRINT DATA LFVFL0ATA 

<J2 10 ····Te "'10 , Iv 

It It It It It SPE"CrFIEO' DTSPLAr.Er~F.NTS *,,**It 

UX uy 

-... ---- ---0-.-0 "ern (j'0T""~ () 0" 

o.oOOGOOF+oo 

O.OOOOOOHOO 

O.OOOOOOE+oO - -.-- ..... --.. --------0:t100ITCOF +110--

O.ooooonE900 

O.OOOOOOE ... OO 

O.OOOOOOF+OO .. --. ------·-----0. (J(;"O'O l1(fE+O 0 

O.OOOOOOE900 

o • 0 0 0 v 0 0 E ... \)() O.OOOOOOF+OO 

------0 ~-ITr. OO'IT OF. +"00-

TF~PERATURE DATA READ FRO~ LOAD STFP: ITFR: 10 CUtl. ITER: 4 ON FILE 4 

-_. __ ._--_._---_ .. _-------------

*ltltltlt LOAD SU~~ARY - 13 DISPLA~niFNTS () FORCFS o Pf<ESSLJRFS Itlt*ltlt 

TNTEGFR STORAr,E REGLIREMENTS FOR LUAU nATA INPllT CP= ;>3.1130 TIt',E: 10.1SnQ IJFfVlOflY 1= 131b tJE~ORY II: 0 TOTAl: t,lh MEMORY AVAILABLE= 31100{)O - ,-- -._---- ---"--_ .. ---

,,1t.,..,.1t CEI\Ti10rO, {vIAS~. Arm ~IASS i'11ll'<'FNT5 e,F TNFRTIA 

rALCULATIONS ASSUtJE ELE~ENT ~ASS AT ELFMFNT ~F~T~nTn 

T (j TAL ~ A :; S : . .0. ~ 0 11 1 E_~O 2 . 

147 
~·ASS Atd·, HJ~'Ef\1 VALLLS /ltd, EVALLl<lH; Hit< ~~nil~ [)1'1:kf-F MJIlt:L. 

THF RFLlJW T':'HLI' IS tiASf:lJ ON A 1'I1:'1(:;hTFllll"-AS.''i Pi-/(,PliiHTIINC,1 10 kArdUS) ~XT1iHlI.,Flldr. SFr·fIIJI,. 

SQUAriE k(]Ol OF 

FIWSl FIRST ~OMFNT/~ASS SFr.OWi S F r(lf~lJ IV· 0 1'1 E NT /I·IA S5 !>10MEN.·-······_· - (tFHTROIOT· 

RADIAL 

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l,rOE -31> 0.11,,4(0 .. 

~LC~ 37 O.11h4Gh h0v~ 3lJ -O.IO~15R 

-':>7-- NCnE'S=-- 3;:-- '3'1-- 311-~!--; " 

YC= 1.;>9Q 0.5250~-02 TEMP: ~ln.D 

~t.SIG2.SIG3: -112.90 -35'?9. 

lTtO FONCES ON ~OGE 3~ -0.103157 

PoH: Fil~r:ES Oi; r,CDE3"--r.. II h40,,' 

lTIC FO~CEb ON ~ODE 38 0.lln405 ~11[ FOAr:~S UN hCDE 31 -U.IO~157 

~lAT: ;> VOl = SX.SY .SH .57: 

-'5a77. 

-1;>.0;>111 

-11.01 it" 

" .111 Il/< 1,.0;>'" 

MAT:';> VOL: 

Slt.SY.SXV.S7= -~~477. 

-12.0;>111 

-11.01 IlFl 

1 , .01 Ak 

1?021<1 

MAI"- 2- VOl = SX.SY.SXY.S7= -~5477. 

-P.G2111 

-11.01A"-

11 .U I Ill< 

1;>.0;>11' 

-qt-.7"~ 

S.1.= 

O.'j'/aIF-Ub 

-96.75/\ 

5.1.= 

O.<;741E-Ob 

-'16.758 b.I.= 

O.971l1E"'-Ob -q".-,,,O S "t;',T .. = 

O.97u4E-Ob 

-11;:. gO 

S.1.= 

O.G7~aE-Ob 

-II? gO 

S.1.= 

O.G74a~-tJb 

-112.9V 

S.1.= 

0.'l71J4E-Ob 

-112.90 

S.1.= 

--;> .. --NonEs::-'"311--z!,----lfo----'3,-'~---'MAT=-2-- VOl = O.G7IJ4F-Ob 

-q7/;'~." 

~nC;'l." 

-97a~'.ii 

%,'1.5 

- .:J",r"\ qa; .. a~7~g. 

.aIl1l9'i. 

a879CJ. 

-aAii'l5. 

aH79Cj. 

-4111l9'i. 

41179Q. 

-~'ia77. 

~'i:~h5 • 

-'~C;U77 • ~'i~b5. 

YC= 1.299 0.2250E-02 TEMP= 51n.0 SX.SY.SXY.S7= -112.'10 -~C;477. 

• 1.SIG;>.5IG3= -112.90 -35~?CJ. -~~477. S.l.= 35365. 

lTIC FOR[ES ON ~ODE 34 -0.1031511 -12.02Rl .1 Ie FOReE-s-l!N,'.;oOE u-r-u;-rn;1nr~----n .-011111--'-' --- ------ --- --,----------- ~TIC FORCES UN ~CDE ao 0.11&40" 1'.01RA 

.'1[ FON[E5 ON ~OOE 3~ -0.103157 t?U;>1I1 

-,-n---- 'Nunc: 5 ,,---:rs---u"'C'r---." ..... I.--- ,~;---' M n ::---7"- VOl:: 

~C= 1.~Qq O.7500E-03 TEMP= 'ilh.O, SX.5Y.SXY.S7= ~t.SlG2.SIG3= -112.90 -3512Q. -'C;477. ~TIC FOkO~S ON ~ODE 35 -0.1031~q ~1?02Al 

IfTT[rDROFS--UNr,OOE 4", O.T'-.,aa,-----"'11.011lIl' 

HII: FONCES ON ~ODE al 0.lll1a07 11.011111 

lTlf FORCES ON hCDE 34 -0.10315A 1~.02Al 

0.97a4F.-Ob 

-112.90 

S.1.= 

O.G7IJ8F.-Ob 

-123.96 

S.T.= 

151 

O.b27a~t-O~ -<;71b.O 

ST6i':= 9645.b 

O.~bq76t-03 ·Q716.0 SI6E: 9;'45." 

2-r SCLID'-

O.5'i~qlt-OU ·97tb.U 

SIGE:: 9b45.b 

O.44?B~t-03 -IJB7b3. 

SIGE= ~H732. 

2-1) 5GLIC---

O.6833IE-a3 -UIl7,,3. SIGE: 41\732. 

n .. ,) h r.l q 7 t. - () .~ - u'" j !"):: • 

SIGi-: 4ti7~2. 

2-1J SOLID 4, -O.2hb'lSE-03 -ue763. 

SIGE= 48732. 

.2-0 SOLID 4< 

-O.h8331E-03 -a87b3. 

STGE= 4A732. 

2-D SOLlll 

2-0 SCLID 

O.S2424E-03 -3532Q. S I GE= 35;>'10. 

';-0 SOLID O.94q~IE-0~ -3532Q. SIGc: i~?90. 

2-0 SCLID ,,-~< 

U.435CJSt-03 -35329. 

SIGE= 3~i'90. 

2-D SOLI D 

,---- ", ---------,-------- --------'2-0 SOLID-- a. 

-0.94951E-03 -35329 • SIGE= 35290. 

,---,---------------,--- ---------- 2-0 SOL 1:lr 4. 

-0.S?4?4E-03 -35329. 

SIGE= 352'10. 

-,-------,-- "----2-D SOLID 4, 

0.a8380E-03 -22399. 

SIGE= <'2355. 
-~ ht r H~ Ftlt<Cf::i' t:!-r···"CUE· --_·U·3-""'1j-.T',>t!tlUA ... _ ... - .... /13 127 ~H<lltr. t'UkCF:> UN I,LiCE 30 -0.11 ... 1107 7."?M~'1 

FI;: -"2 NunES= 38 115 1111 'n . MAT=" VO'-= O.9741!F.-06 ~~.'r= 1.100 O.o7~OE-02 TEMP: 4"2.". SX.SY.SXY.S7= -123.9" 

2-0 SeLlO 

RIG'.51G2.SIG3= -123.9b -22..,9'1. -2?~~R. S.I.= 

-~a~w-~g~~U g~~g£~·--t~+'Ht-~g·~ .. ---~:~~~Jj ... --------.............. - .. -.-.... ----.-... ---.-.. RT4Tlr. FORCES ON NODE 1111 0.1?1I80" ".831?7 R~4TIC FOWCES ON NODE 37 -0.1161107 7.82R..,'1 

"r:--:4 ... -·I~ODES=--~C;.. lj" 4S 311 • "f;(AT""·;i-"VOL:·O.97118E-06 ~r..YC= 1.300 0.52S0E-02 TEMP= 1I11?" SX.5Y.SXY.57: -1?3.9& 

lHGf.SIG2.;l!G3= -123.96 -?2..,'l'l. ·-2?~~A. 5.1.= 

RTAJIr. FO~CE5 ON heCE 39 -0.1, ... 1/0" -7.B2H..,Q -'IITA lIr. FORCE S ON -I\iOOE-----.u;----O';T>liao-i> .. ---.. -;;i>: <1 31 ? 7 ~~Alir. FOACFS ON wODE 45 O.I?IIAO~ ".831?7 

STIIIC FO~CES ON NODE 311 -0.116110" 7.821<"''l 

I"r=-.. -.II ...... NIJDfS=--qlf~1 lib -----rq----;----MiiT::---,:;· ... VOI.= 

rr..~r.= 1.300 0.37S0E-02 TEMP= IIII?" SX.SY.SXY.S7: 5r~I.SIG2.SIG3= -123.96 -??~q'l. _?2~~A. 

~ f:H~~gSg~~ g~ ~gg~---U~~·~riL:-Ha~g~ --.. -·_ .. :r: ~~~ Jj ~T.'IC FORCES ON NODE 116 0.1?480" 6.M31?7 

RT&lIC FOHCES ON ~OGE ]9 -O.1161/0~ 7.B2A"''1 

Fl'·:£' - .~~ ,.l.IDt:S= q·I .. -"Iili--ifl---4ii' .... ~·~ ... MAT= 2 . vOl: 

~r..t(: 1.300 O.225uE-02 IEMP: 482."', SX.SY.~XY.S7= 

~rG'.SIG2.SIG3= -123.g& -??''1'1. -??~~A~ 

~T~rIC ~ORCFS UN ~UCE ~I -0.llh407 -7.i!?R'Q S U riC FOiiCF. S" ON ~'\ODI'-"·"-I/B-O ~"I?aIlO" ~ .. ~. ";'~. i\j 1?7 

RTATrr. FU~CFS ON ~ODt 117 O.I?IIHU6 ".B31?7 RT'tT~ FORCES ON NOOE I/O -0.1'''110... 7.H2~'G 

:, .. =.... ," .. NOI,I'S=- .. /i;;--.. ·,j"q·-,j·i\-li.,... .. · .... ;- '''",A Ti: .. 2 VIII = t'r..,'!'r.: l.lOU O.7500E-03 TE~'P= 11112.", SX.SY.SXY.S7: 

R1G',SIG?SIG3= -12!.96 -??..,GY.. -2?~~M. 

RTIIT~ FOkCES ON ~OD~ qi -n.l1"~tl7 -1.8?H'4 ~UHr. FORCES Qf'" I\CDE ..... --q9---0:1?4ilOI. "';'''~il31;>7 

~TITIr. FOiiCE6 ON NOOE 1/8 O.I?IIBO" ~.831?7 

RT&Tlr. FORCES UN ~OCE 41 -0.11hIl0" 7.821<'G 

'I :: ~7 j-JOOF. S; 44 -51- '50 .~. '''", ~IA T= .~ VOl = tr.u:: 1.~02 0.~~Stl"'-02 fH'jJ: Llh?11 sX.;>Y .:-iXY .. 'i7: ~ IGI ,SfG2. :H,,3= -132.1b -27 "<;.1 -;>4"~.'i 

~HI!T1f. Fu';C':S t)I'J NCet LlII -0.1 ;>LlIIO~ -" .2Il? III ~TiHlr. rOflCF S ON i\OCE SI ··· .. 0.11?9" -f.,.5QQLI~ 

S!AI,TC FUiH:~ s Oj~ [\OLt: 50 o .1~?9'11 h ... 594~C; 

~HUIC FuRr.E-S ui' f'"CCI: 113 -,i. I ?uAOI- ".20;>111 

"I :: 'Hi NOOES: qS" 5'2 '51 LlII 

<16">:0 

MAT: 3 vOL:: 

tr..I'r.: 1 • 'n;:> O.b750E-1l2 TEMP: SX.SY.SXY.S7;: 

S[(;t.SIG2.;;11;3= -132.lb -?7A~.1 -"'1;>~.'i 

STUlf. fOtiCF.S ON «OCt: 'l~ -O.I;>IIAU<; -il.2()21<.I 

RTIITIr. FORC!'S O,~ [,<CCE -5;> "-0'.1 ~?q~1 -".5941/<; !HAITC FOflCES ON I\CGE '51 0.n?9J' ".''19a<; ~r"'TJ~ FURCES ON "CIlE <11/ -0.1,.>1/80" 11.20?1<.1 

'I :; '--"'~q NO(;ES= 

cr..VI:: 1.30;> ~ItTt.S1G?SlG3= 

SlArrr. Fu"'C~S 0 .. -~TATIC FORCES ON 

.. Lli>' "'''-5 3-,-.,;r-"--Ol"l· , .... i"A T= 3 VOL= O.!~SUt-O~ TE~P: <I,,2_P SX.SY.SXY.S7;: -132.16 -;>7A~.1 -;>'1;>~.'i 

I\OCE lib -0.1?4"01- -A.2021Y 

NCCE ·-"-'5T"-rr. n?""'1 ·~"h. S9QII~ 

!HITJr. FORCFS OIY 

!IT II fIr. FORCE ~ Or, 

NUDE S; 0.13?'l~1 ~.'i9~1I'i 

hUGE liS -~.1?4AO" H.?U?I11 

:r=" "·1I1l· --NODFS= ..... -. '~7 ·......,,-zr·-.;T-,!1; .. _ .. , 

cr..'C: 1.3n~ O.37~~E-02 TE~P= ~"?P 

~I~I.SIG2.SIG3: -1~?16 -~7A~.1 

. MAr: '3 i/Ol.= 

sX.SY.SXY.S7;: -?'l2'1.'> ~T~rlr. FO~r.FS O~ NGCE ~7 -0.1?41<0" 

Rr411(: FORCES ON ~CCE S4 ····O.13?'l11 

RT,rrr. FORCE9 UN ~UOE S3 0.1~?'l11 

R~4TJC FO~CES O~ NCDt <lh -O.I~<.I"Oh 

aa"<;5 · .. ~<;4 .. il7 

O.ii~Ot-n2 TE~P= a,,?:p 

-"."O~la 

-".5'1'lIl'i 

".'iQYII~ 

1\.2021" 

::'1:" al -NonES:: MAT= 3 VOL: 

cr..Y\,:= 1 •. ~0; 

SrGI.~IG;;.SJG.~= 

!HATJr. FJj~CE~ ON 

STAnc FORCES ON S·TArrr. FlJ"C~S 0" 

SlATIC FCRCES LJI~ 

-132.lb -?7R~.1 

S~.:iY.SXY.S7= 

-2'1?'.5 

~I = 

'(1.= 1.. ~n? 

(;l .. SIU'r\. :'l('~= 

'A r Ir H"ir.~~ u" ATiC: H)~r:I':; u:' A T I r. ;: (;I<C" ~ 0,', AlIr. HJR.:FS (JI'; 

~cDE <.IR -~.I;aHn ... NCOl:: - --5S-0~ n?'l11 ~CC~ Sa 0.13?'131 ~OU~ 117 -Q.l?IIBO~ 

4G '56 

{. .. I'::. ()!.'!':' - ~J " 

•. ~!,;~.I~ . ·"JCt ~C; 

r,cL'~ c,~ 

!\GLt 55 hCCE iJf! 

'is 

it!l'l~= !.Ih? .. 11 

-~7""'.1 - Ij .. 1 ,..>4r. 01"1 

(,.~~?4~1 

(J. 1 ~? q :11 -O.I?4ROn 

-A.eO;'" 

-".'5'l'lIl~ 

... 59'111~ 

".20;>111 

MAT= 3 

~" .. h1.S1.Y .. hl:: -;><;;>'.~ 

-F, ... i?O? 1" 

-n.:;"Cj"", 

... ,"4a'i R .. ?(,?, u 

O."7UIlE-lib 

-123.96 

9. I. = 

o.<;7alll'-ob 

-123.4& 

5.1.= 

O.~7ar.E-(jb 

-1;;3.'<b 

S.1.= 

II.l",qo;F-<l5 

-13;;.11> S. I. = 

O.759'iE-Il'i 

-1~2.lb 

S. 1.: 

il.7S9'iE-05 

-13".16 S~I.= 

O.7S95E-()5 

-132.16 

S.I.= 

0.75G5F-OS 

-Ui:.lh 

9.1.::: 

,,~ NODFS= 

yr:= t .10n 

Gl.lHG2.SIG3= 

ATrC FOl<CES UN 

-A1'IC FuRCFS 0" 

A rrr. FOI;CE S [j,. ATIr. FURCFS Ur, 

I~AT= 3 1101 = sx.SY.SXY.S7= 

-M27.00 

0.91a4;:-05 

-137.17 

-~.b7AIl? 

-1.3bIQI' 

1.~611l11 

·'.n71\1/? 

'MAT= 3' VOl: 

SX.SY.SXY.S7= 

-fl27.00 -~.&71l42 

·-1.3614R 

1. 3b 1 41' 

'.n71l4? 

MAT= 3 VOI.= S~.SY.SXY.S7= 

-fl27.00 

-"'.,,7I<a? 

-1."'bltlR 

, • ,,, 1 1111 

'.b71l4" 

MAT: 3 VOl: 

SX.l<y.SXY.S7= , -1127.00 - .... b71114? -I.,&IIIA 

1 .. ·~b 1 Lin 

'.b7RIl? 

S .1.= 

0.9111410-05 

-137.17 

S.1.= 

O.GlqUE-u5 

-137.17 !:i. 1.= 

(,.GI/j~I'-05 

-137.17 

5.1.= 

152 

-22~'i1l. 

??434. 

-??'i~8. 

224::14. 

-;4?":.5 

;>7qj .4 

~;>4F':-' .. ~ 

?l U 1 • tJ 

-R?7. 00 

"'H'l.8? 

-An.OO 

"'11".B2 

-A?7.00 

"AY.II? 

-1\;>7.00 

"1l'1.82 

.... ---.-.-............... -- -· .... - .. -·---.. ----·------~ ..;O-S·ULnr 

O.43320E-03 -223'19. 

SI6E: 22355. 

.... ·· .. ---~----E.;;O- SOITlr 

-0.43320E-03 -2239'1. 

SIGE: 22355. 

-O.91?39E-03 -~2399. 

SIGE= 22355. 

2-( SGLIrr 

2-0 scurr O.13731~-03 -?7H~.1 

SIGE= 27211.8 

D.n"UJ3t-oa -2785.1 

SlGE= 272"./j 

... 2''-0 SeeN 

-ti.hIl437E-OI/ -2785.1 

SIGE: 2724.8 

2-0 SOLTO 

-O.1373IE-03 -2785.1 

516,,= 2724.8 

2-0 SCLTiT 

.... P ... 7!l'::l4e-t-OLJ -n~5.1 

:»(,1:= r',?4 e r! 

2-D SGLlD -0.ICj7411t.-Oa -b83.40 SIbE: 630.41 

-O.3bbIl7E-OII 2-u SCLID tl -b83.4(J SIGE= 630.111 

"0.171~8E-oa 2-lJ SOLlu 

-b83.40 " SIGE= &30.41 

- --0 •• __ • ____ ." 

0.171551:.-04 -b83.ljO SlGt.= 630.41 

coD SGLID 
MA I:: ~ V(JL: 

5X.5Y.SXY.S7: -A?LOO -'.b7A'I? 

- j • :~61 4R 1.361UR 

_.b7~a;> 

--~H-I-I~--FOi<Cf S cll'rl';c(''t-- --e-..---~·H------ r~ :t'l~'1*, 

STArrc FONC~d ON ~UDE 60 -O.I~~3?6 -".~7RIO 

Iff": "'3 NODF.S: be 69 I>Il (,1 • MAT: 3 VOL: ~r.'c: 1.111 O.~250E-02 TEMP= U'lU.~ sx.SV.SXV.S7= SIGI.SIG2.SIG3: 1782.7 1"011.7 -1''I.?6 __ sH T I r _FOkCE S _ ON _ /liCDE ___ --.I>?--!'!!l..J_353?6 ________ ll_. 27 H I tl 

SIAIIr FORCES ON NODE I>~ O.I?RO~I 7.'1'1.'16 STATIC FORCES ON haDE 68 O.I?RO~I -7.44,96 

s,HrIC FORCeS ON NODE bl -O.13~3?6 -'I.27AI0 

(J.9141lF.-05 

-137.17 5.1.= 

0.9144E-OS 

-137.17 5. I.: 

u.I071E-04 

-134.2" 5.1.: 

0.1071E-04 

-134.26 5.1.: 

--FI--;---<;-'I-NOOES:i--,,3--70~--"-?--;------MAr=-- 3---- VOL: -0.1 !i71 E:;04 -

~r.rc: 1.~11 O.7500E-03 TEMP: 444._ SI.SI.SIY.S7: -13u.~b 

!HG1.SIG2.SIG3= 1782.7 16/11.7, -I_4.?1> ti.I.; SIAlIC FOMCES ON NCCE 6, -O.13~3?6 a.27Alo ---!lTATIC: -FORCES ON - /\00£ ---71l--0-~-j?80r:;1 ------7 .44'9/' 

~TATIr FOkCES ON NeDE bq 0.1~A051 -7."'I~'1/' 

STATIC FORCES ON NODE 02 -0.1353~6 -4.27AI0 

--FY:- ---~'i---NOOES:---65--r2--tT--64---;----MAT=----3- - VOL: -0~-12~~E-04 

~~.YC; 1.~17 O.~250E-O~ TEMP: u3~.~ SX.SY.8XV.S7= -119.23 ~_IGI.SrG2.;lIG~: 4b04.5 "47R.~ -I1Q.;>3 5.1-.: 

-----~f! H~ ~g~~~ ~ 8~ ~~g[---9~-~g:-t ~~~n ~~ :~;~t STATrr FORCES ON ~CDE 71 O.10h570 -;>O.U~Ql 

S'~TIC FORCES ON NOOE b4 -O.I?A051 -lb.271" 

"I :- ---C;f,-I~On-F-S=----;'~ - n--f?--6'i----, --M-A T::i- 3 -VOl.: 

xC.ve: 1.~17 O.n75~E-d2 TE~P: a3~.a 5X.SY,;lXV.S7= 

SIGI.STG?SU;3: 4b(Ju.S 1147".1 -ll'l.?3 SIATIC FORCES U~ NLCE 66 -O.I?~O'i1 1/'.~7a*, 

STATIC FONCE5 UN NCO~ 7~ -0.106170 ;>0.4;'11 STATIC FO~CtS UN NGuE 72 u.1U6'i70 -~".U~'1l 

STATIC FORCES UN ~GCE bS -u.I?RU51 -lh.~701*, 

Fi': ~7 ,;()D~S: 67 ]"f,--- -/3-- -- -f; ... - , MA T: ~ VOl: 

XC.YC= 1.311 O.5?50E-u2 TEr~p: 43~." SX.SY.SXY.S7: SH;1.S162.SIG3= "boa.s ""71i.' -IIQ.;>3 S_liA TIC FORCr:S t.ii. "CCt b7 -0.1 ?110,1 1 h .2711,., STATlf: FORCFS-Of< ;,Cut::-------74 ---0.101><;70 --- ;>O.II?'11 ST.TIC Fn~CE~ Q~ NODE 7] 0.101>570 -;>0.4?QI SIATIC FG~CE~ UN ~CCE b6 -~.1?A051 -1".274" 

-iH: <;F i;Or;E5: 

xr.yc= 1.317 S.ll;I .SlG?~!r,3= 

sf.rlr f0NC~o U~ 

STATIC "UR~ES aN stATIC FOHCES uN STArl~ FO~CtS eN 

6875 -74 67 , C.!7SDE-02 TFMP= a~~.N 

.hOu.S qU7A.~ 

~GCt od -O.I?KO~l 

~CCE 7~ 0.101>,70 ~CCt {4 0.10,,570 ~ULE hi -O.I?ROSI 

"I = ~C; 1-;CrFS: - 6q--T6 Tt5 _ .. ';f\ I 

IC.'C= 1.!17 O.225uE-u2 TEI<,P:: U3'.~ 

""71l.' -Il. I ?RO'i1 SI~I.aTG2.~IG~= 4604.5 STATTC FO~r:ES J~ ~CCE hq STATIC FORCES ON ;;GCE- 76 -0.101.'i70 SEArrr Fn~CE5 O~ ~CCE 7~ 0.106'>70 sFATrC FOkCtS UN NOOE 6R -0.1?1I051 

FI':: 1i0 NCnES= xr.vc: 1.317 ~[GI.SIG;>.:'IG3= 

STATIC F"I-ICf~ Oi\ -~TArrC FORCES ON ~TA TIC fui<CES UN ~UTIC FllRCEti ON 

70 77 ---7~ -,,9- 0.7~OUE-03 TEMP: .3~.1 

Ub04.5 41171l.~ 

NCGE 70 -O.I?Ali'il NCCE 77 -O~10/'~7n 

NOQ~ 76 0.IOb'i70 ~OGE b9 -O.I;>AO~1 

MAT: 3 Vlii.= :i x. • ~ y , ;j x 'y • S 1:: - 11 ... ,,-s 1 .... ?7"" ;>0.4;>'11 

-?0.4;;ql 

-I .... 274h 

i~AT= ~ VOI.= SX.SY.SXY.S7= - 11 q.", 1 .... 2711" ~0.4?'1l 

-;>0.4?Ql 

-1".27,," 

"'AT= 3 VOI.= SX.SY.'>XY.s/: -II Q.?_~ 

Ih.21l1h 

;>0.4?'11 

-?1l.4;>Ql 

-11>.2.,46 

--F! = ----hI - -iliODES=----- Tr-,"-----Tl!-----71-----;-----MAT: -3 VOL: 

xr. YC: I .3?4 O.~e50t-')2 TEr4P: iI?I.0 SX.SY .SXV .57: ~IGI.STG2.jIGJ: 1772.1 767R.O -H7.aOq <;_TATIr. FORCE'; fiN NOCE U -O.IOh570 '?'l~?~ 

~rATIr. FOMCES ON NGCE 7Q O.6~77~0"-01 ~H.2IRI 

STATTC FD~CES UN ~CCE 7R 0.1>~77;>n"-nl -'ft.?lkl S'ATIC ~OHCES UN NOCE 71 -0.10,,570 -~?9';>R 

,I = .,;> r;OOF:S=- xr..YC= 1.324 SIG1.SIG?.:ilG3:; 

,~TAT[C FO~CFS (J~ 

!HATIC FllliCFS uN ~TATlr FONC~3 UN ATAITC FU~Ct3 Of< 

FI ;: 

7! 80 G.b7S0E-u" 17 7 2.7 ~;GC E n ,\CCE ijO 

:\CCE 7'1 ~GGE 7? 

74 P.I 

7G n- . TEMP= 4?1.1I 

7,,7H.0 

-O.101><;7n 

O.h~77?I\F-OI 

O.h~77;>"F-,jl 

-0.10b<;70 

FG 

i~AT= 3 VOl: SX .. SY.SlCY .. S7; -"7.010'1 ~;>.'n;>11 

'M.?IlIl 

-~fl.21HI 

-~~.9~?A 

-0.12?'1F-';04 

-119.23 ti. l.;: 

Il.I;;'9F-UII 

-11'1.;>1 5.1.= 

O.12?9~-Oa 

-IIY~?3 _ S.1. -

O.1229F-04 

-II q .23 

S.1.= 

,).12,o9f'-oa 

-11~.2J 

S. I. = 

o .139il("04 

-d7.40'l 5.1.: 

O.13qO~-O~ 

-~7."O" 

S.I.: 

153 

-R?7.00 

hIl4.1l? 

-A?7.00 

61i9.82 

1 ""1.7 

191".9 

0.3&&a6£-04 -6M3.AD 

SU;E.: b30.<l1 

G.19743E.-O" -6H3.40 SIGE.= 630.41 

0.lbM84E-OS 1782.7 STGE: 1850.5 

-O.?/l7B7E-O~ 1782.7 SIGE= 1850.5 

2-D SCLID 

2-D SCLID 

2-D SOLID 

2-D SOLID--

2-D SCLID 

2-U SOLID 

--------------------- -------------- ----C!--o-!fOCIlJ". 

1114 1.7 IqI6.'l -O.lbB88E-O~ 17R2.7 SIGE= 1850.5 

-- ------------ ------------------------------e-;;osOLTO, 1147R.! O_~0343E-Ob UbU4.j 47~3.7 SIGE= 4b61.9 

4471\.1 

47?~.7 

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47?~.7 

7671'..11 

7RhO.1 

7678.0 

711hO.1 

0.70M42~-Ob 11604.5 SIGE= 46bl.9 

1l.453?eE-06 4h04.5 SIGE= Ub61.'! 

-O.q~28~E-Ob 0604.5 ~IGE= Nhhl.~ 

-U.70H07E.-Uh 4hO •• 5 SIGE= 4h61.9 

-0.?U',,2C[-06 4604.5 SIGE= 116bl.9 

-0.12bUbt-Ob 7772.7 SIGE= 71l13.2 

-0.35137E-OI> 7772.7 5ISE= 7813.,0 

2-0 ~CL1a, 

2-0 SCLIIJ' 

-2-D SCLro-. 

2-0 SGLIU--' 

2-0 SCLIO-' 
dtl :;:~~j~~~~-i;i -~~:~i~~ 

7~ -~.lnn~7n -~?9~?A 

Fr= hU t,UlJEcS= XC.VC: 1.~24 

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STATIC FUNCES Uh 

-s-HT Ie HJ~Cf:~ UN 

SIATIC FORC~S o~ 

STATIC fO~CES ON 

B 1 7 a , IE,·P: ~?I.1l 

7i>7A.0 

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SIGI.SIG2.SIG3: 7772.7 767n.0 

STATIC FORCES ON ~COE 7i> -O.10b~70 

--~TATIC FO><CfS UI, r,cilE-------- 83--- -U~-b~77?OF--OI 

STATIC Fu~CE5 o~ KCDE b? O.h'i77?OF-OI 

sTATIC FONCES UN ~CUE 7~ -O.lnh57a 

-'fi hh NUDES: xc.Vf:: 1.3;>U 

SIGI.SIG;:>.SlG3: 

STATIC FO~CES ON 

STATIC F-UKCfS UN 

SIATlf FU~C~S UN 

STITlr FONCES UN 

-I'r= 1>7 i,ODES: xr.. YC: 1.331 

SlI;I.SIG?SIG3: 

STATIC FORCES UN 

-STATIC FORCF5 ON 

STATIC FOkCFt D~ 

STATIC FO"CE~ .Oi, 

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5TAT1C FOKCES UN 

--STATIC FORCES G~ 

RTATIC Fu~CF5 ~~ 

RTATIC FORCES Ow 

77S. 83 7& 

0.7500E-03 TEMP: 4?1.0 

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NGOE_ 77 __ -",O.JOh~70_ 

~COE HI u.o'i77?OF-OI ~UCE b! U.h~77?nF-Ol 

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STGI.SIG?SIG3: 112"17. 11?'Ih. ~lATlC FORCES ON r,CIlE 81 -O.b<;77?OF-Ol - STATIc FORCES OK ~CDE BE -~.9h?7R~F-l? 

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STATIC FORCES U~ 

STAIIC FO~CES UN 

- e 2""- - - a-q-- -- 88"- ill----;-

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MAT= 3 VIII: 

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hi .. rJiJ77 

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f.-AT: 3 VIH_: 

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hI. ,",,77 -ttl.41177 

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""F1--:---- 71 I<fOOES:--aT-----.:ro----R<r--R?- ---;-----MAT=-- 3 -- VOI.= 

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" ~lGI.SIG2.SIG3: 11277. 11?~h~ -1~_707 

4 n:Hf.-~~~gf~-tgg~ ~fi :~~~~~g~:~--~~:~~ii---q STATI~ FOHrES UN ~CCE 8ll O.3A708~F-Oq -hl.ga77 ~TATIC FONCES O~ ~COE 81 -O.b~77~OF-O' -~a.dR~7 

-n:-- 7;>-- NUUFS::--- 81r--'n-- qU-- IIr-;----"MAT: -3 VOl.: 

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SIGE: II?':IO. 

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SIGE: 112'10. 

c-r, SOLID -0.74?91E-07 11277. SIG~= 11290. 

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-1.2110-; .. -"~IH)~_ 

-LIo~fIO ~ .. 

-2MM.'? 

-2~";>;>~ 

-r:?NM.?;;l .. -;>I'';';>? 

-';1>1'.;>;>. 

-;?I1I<;>? 

-'>lil\"';>. 

-107<17. 

-107<17. 

-107a7. 

-10747. 

-107117 • -107'H. 

-107<;] • -18~7 .• h 

-18?7.1-

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45A,.O~ 

4':i1\.0~ 

4~fI.n3 

31fJ3.h 

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310~.f, 

310~.f, 

310~.f, 

,,100.7 

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b100.7 

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n100.7 

6100.7 

h 111( •• 7 94~~9.1. 

9419.0 

'l4~<;.O 

'1'13'1.(1 

9 .. 'IQ.0 4,,~q.O 

4:J O;~.:) 

15101'.. 

UIOA. 

131 OR. 151""_ 

131 Oli. 131011. 

131 r.A. 

***** RLUTINE CU~PLErED ***** Cp = 5.h~~0 

-'/FOF ENt:OUNTEREO-Ulll'!1:nS----------- -- -,-----'-----

156 

SIGI 

-1.8710 

-1.8725 

-1.il7S3 

-1.8749 

-1.87'53 

-1.8725 

-1.8770 

-4H.847 

-4/1'.842 

-4b.84~ 

-<lM.1l45 

-4LI14':i 

-48.842 

-Uh.Il117 

-14.9ul 

-79.8'l1l 

-i4.d'1'1 

-7'1.900 

-7".~9q 

-7 Q .898 

-74.901 

3142.1 

314d.1 

3142.1 

3142.1 

31~2.1 

31 "2.1 

! 142.1 -llJh.23 

-I nb .2_~ 

-106.23 

-IOb.2.5 

-1l1b .. 23 -IOh.21 

-I ('n .':>:; -I 1'1. / ~ 

-llo.71l 

-119.78 

-119.7H 

-119.78 

-11 Q.11l -II Q. IH 

-I"'R.~" 

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-1?il.54 

-128.53 

-12M.54 

-1'>8.54 

-1?Il.S4 

-13b.8lJ 

-13b.llu 

-13b./lU 

-136.8<1 

-136.80 

-13b.tiO 

-13b.80 

458.03 

4!iB.03 

4Sb.U~ 

4'ib.lJ3 

tlSA o IJ3 

USH.03 

458.03 

3103.b 

3103.b 

3103." 

jIO~.b 

3103.0 

3103.b 

:JI03.6 

6100.7 

6100.7 

bl00.7 

0100.1 

bIOO.7 

61Uu.7 

~ I 0".7 S43'1.u 

9434.0 

91139.0 

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943<;.u 

'143'1 .. 0 

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131 ull. 131 tIll. 13108. 

13108. 

1310t!. 

13101>. 

131 OB. 

81; 

13108. 

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-54841-

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·'8Ii/-on. 

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-4?003. 

·u;>OO L 

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-~fi"2'>. 

-10747. 

-10747. 

-10747. 

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-1627.6 

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314.00 

31~.OO 

314.00 

314.00 

314.00 

314.('1l 

314.00 

?qI>7.4 

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?'i67.4 

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?467.4 

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S'ltl".n 

5'18"." 

598n.6 

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'i'i~h.b 

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",\hb.3 

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93ob.3 

43b6.3 

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13101. 

13 j 0 I. 13101. 

13101. 

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5 

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10.1721 51 8/811 Cl'= 

SIG3 

;;54906. 

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-38064. 

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-?2f538 .. -nS3t. 

-;>2530. 

-22~3b. 

-;>68'1'1. 

-?bb9~. 

-;>b8"~. 

-':be9G. 

-;>hIlS9. 

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-,lh<l~~. 

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",~el"':. 

-4?14S .. -4(l1 tJ'.: .. 

- ... ~ll.~ .. -4;145 .. -?~97b • 

-?~"7b. 

-';0'1/6. 

-20970. 

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-I08G~. 

-tOIj~,! .. 

-1(,895. 

-IO~"5. 

... 10BGS .. -IOb.,S. 

-108<;5. 

-1'1b9.~ 

-1'lb9.~ 

-1969.5 

-19,,'1.5 

-1409.5 

';;'1'16<;.5 

-1%9.5 

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-1.5<;.0)1 

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~ 139. (,B 

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-131.~8 

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-131.~8 

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-109.57 

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-104.57 

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-109.57 

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-H.4i?9 

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-I>H.42Y 

-;>.5574 

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-2.5574 

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7 -54906:· 

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, 54J:l72~ 

54872. 

541;12. 

54b72. 

54872. 

54 III ;;. S487 2. 38021,. . -----311026. 

38026. 

38U26. 

3802b. 

3802b. 

38u2&. 

2"a51. 

,22451. 

224'::11. 

224S1. 

224'::11. 

2i'4'S I • 22'151. 

~00U3. 

3000.1. 

300(13. 

3GOllj. 

30003. 

3000_l. 

~OOVj. 

4IYbl'. 

4lYbn. 

419bo. 

419ho. 

4196~. 

'11 'iocl. 41966. 

2tl7bO. 

2A71lu. 

267110. 

2M/bU. 

28/ou. 

210 7 ;"1. ~~71lG. 

1,,6"4. 

IObY4. 

IOb94. 

IObq4. 

10bY". 

IOb94. 

1o,,<;4. 

17bO.0 

170b.0 

17ob.0 

17b6.0 

17b6.0 

17bh.() 

1766.U 

53'1.70 

539.70 

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53".70 

5~9.70 

5 ,~". 7 0 539.70 

3Ib<;.2 

3Ib<;.~ 

31b9.2 

310<;.2 

31b9.2 

JlbY.2 

3169.2 

b15A.0 

b15u.0 

b15a.0 

6154.0 

hl~4.0 

"154.0 

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13101. 

13107. 

13107. 

13107. 

13101. 

13107. 

7 54872.-

4.770 

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54'104. 

54904. 

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549u4. 

54904. 

311U36. 

38u36. 

5tiUjb •. " 380.10. 

3e03b. 

3aU36. 

38U~b. 

22454. 

22459. 

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d24~g. 

22q'jG. 

2d4'::1". 

.:24~g. 

3UU41. 

30U41. 

3<JU41. 

j (, G 1'1. j \) \J iJ 1 ... 

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3 (J iJ '" 1 .. l.I.'::tJ ~ ":0. 

<.It?u~''1. 

4d(j3~. 

42v~q .. 

I..ldUj~ .. 

42U~~ .. 

4t- '.J 3~. 

28t~7. 

28e~7. 

288:7. 

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2b057. 

2nn57. 

2b~~I. 

107H. 

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I U 7 b 7 • I u / ~ 7 • 10767. 

lu767. 

lulU. 

leo 5" • 7 '183?7-" 

1832.7 

1632.7 

1832.7 

1832.7 

led2.7 

591.ll 

5,7.11 

SG7. II 5~7.11 

5''] .11 5g7.11 

597.11 

323~.1 

3235.1 

323~ .1 

323'::.1 

3235.1 

3?3'::.1 

!23::.1 

621U.2 

6210.2 

621lJ.2 

621U.2 

b? 1 (,.ibi'lU.i' 

o21u.2 

451) 7.: 

95u7.5 

95U7.5 

95u7.5 

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13110. 

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7 54904~ 
l_ 

i 

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VAX/VMS AIOIH G ll-t<AY-1 qa~ 10:10 LPAO: I\-MAY-I'I!lq IIJ:n lISHn: IAIISYS. In",Y} G.LIJG: J 

VAX/VMS ATOIIIY Q 8-"'AY-1984 1 0: 1 0 LI'AO: II-MAY-IQ/l4 IO:? 7 IiSFR.~: (ANSYS.1Ut.VI !i.LOG: I 

AAA rr TTl IJUlI " I~ Y 

A A T () II N N Y 

A A I II () NN N Y Y 

A A T () (J N N N Y 

AAAAA T 0 () N NN ~ 

A A T () II N N Y 

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IlQQQflll 

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all (l(l 

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LL 0000110 GGGGGG[;[; I: II 

L L· COCOOO GGGGGGGG I: 11 LL 00 00 GG :: 1111 

IL 0(1 UU GG II 1111 

LL 110 00 r,G II 

LL- 00 ·00- GG J I 

U UO 00 Gr, :: :: 11 

LL Oil 00 GG I::: II 

LL U(J 00 GG GGGGGG :1:: II 

LL- ----- 00 --.----- 00- GG GGGGGG II:: II 

LL 00 00 GG bG :: II 

LL 00 00 GG 6G II II 

LLLLLLLLLL OUOOOO GGGGGG :: 111111 

LLLLLLLLLL oocoeo f,GGGGG II 111111 

AAA r TTTT oun N N Y 

A A T n (J l- N V 

A A 1 (I (I HN N Y I 

A A T 0 II IJ N N Y 

VAX/VMS 

VAX/VMS 

VAX/VMS 

AAAAA T II (J N NN Y 

A A T u (l N N Y 

A A 1 (JOO IJ N Y 

ATONY Il 8-MAY~1984 10: I 0 LPAO: h-MAY-198~ 10: 27 LSfR3: (A.NSYS. TOrH] (;I.LOG: 1 

AIUNY Il 8-MAY-\Qa4 10: 1 n L'·A" : 1I-1·1A Y -1984 10: 21 l'Sf-IU: lAl'JSY~.IUNYllJ.LUb: 1 

ATONY Q 8-MAY-!91\4 10:111 LPAIl: II_MAY_148a 10: 27 USFR~: IANSYS.1DNY] (I.LOGII 

/--D~l:Ofl f ,-- - -' .---... --- - PA6E 1 

--------- ------------.-

ACCO~NTING INFOR~AIION: 

B(JFFFREO I/O COUNT: OTRFCTIIO COUNr:-

PAGE FAULTS: JLAPSFD CPU TIME: 

157 

to:IO:;>~:14 

P~AK WIlJ./UrJ[; ~FI SIZE: PFAK PAnE FTrF RIlE: MIIIINTEO IIUIIIMFS: 

500 

4q.~B 

o FLAP!'lI'(l rrl"F: () oO:04:37.;J? 

APPENDIX 3. 

PHOTOGRAPHS OF TEST SPECIMENS 

159 

This appendix is supplied to document the coating test specimens used for 

this program through photographs. 

Figure 90 shows two flat and one ring specimen before testing. 

Figure 91 shows two flat bend specimens after testing. The top specimen was 

bent with the coating in tensions; the bottom specimen with the coating in 

compression. The flake shown with the bottom specimen is a piece of coating 

which spalled from the middle span. 

Figures 92 and 93 show close-ups of two bend specimens after test. Figure 92 

shows a coating failure due to compressive stress and Figure 93 shows a coating which has been subjected to substantial tensile bending stress. 

Figure 94 shows four bend specimens after test, and Figure 95 shows a closeup of one of these. 

161 
Figure 90. Specimens Prior to Testing 

Figure 91. Two Bend Specimens Ater Test . Upper: Coating in Tension 

Lower: Coating in Compression 

162 
Figure 92. Compressive Coating Failure 

Figure 93. Tensile Coating Cracks 

163 
Figure 94. Specimens After Test 

Figure 95. Bend Specimen After Test 

164 
End of Document 

r·