Mechanical design of a Curtis turbine for the oxidizer turbopump of the M-1 engine

APPENDIX A S:IT_JC_.L DESIGN CRI'.FERIA Compiled by F o D° Ronkovich and L. K, Severud I, GENEP_&L All components shall be capable of withstanding limit and proof loads without undergoing excessive permanent deformation and without deflections which will adversely affect the performance cha r acteristics of the turbopump assembly.

All components shall, be capable of 21 re-uses and a total operating duration of i0;500 sec. Adequate structural integrity and high reliability is to be a primary design objective° The ability of all components to meet the strength and deformation require- merits shall be substantiated analytically. Experimental substantiation of structural integrity shall be performed in cases for which the reliability of the stress analysis is questionable.

II° DEFINITION OF PRJ_SSURE TERMS MoE, 0oP°.---The maximum expected operating pressure Proof ............ ]..2 M°E°0,_i° ..... .design to yield Burst ........ 1.6 M°E°OoP°- .... design to ultiraate I!!. DEFINITION OF !_IERjYIAmT__ND L I_ LOADS Limit Load ...... The critical load or combination of loads and environment_ the occurrence of which is expected at least once during the life of the component Design to yield .... 1.O x Iimit load Design to ultimate ..... 1.5 x limit load IV, SHOCK AND / OR DYNAMIC L OAD CRi!%RIA Equivalent static loads derived f r o m dynamic "analyses are to be used and these equivalent static loads are subjected to %he same factors as given in Section IIl of this Appendix° V, TEMRER A 'IURE DESIGN CRI_ERIA Components will be designed for t[he critical operating conditions created by the combination of p r essure and ie m perature gradients existing during turbo- pump assembly transient and steady-state operation° Metal. temperature shall be based upon heat transfer analysis° Experimental data obtained from test programs for similar hardware shall be factored into the analysis whenever such data are ava ilab le° Page A=2 VI. DEFINITION OF STRESS CATEGORIES _11 _ A. PRIMARY STRESS A stress developed by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments.

The basic characteristics of a primary stress is that it is not self-limiting.

If a primary stress exceeds the yield strength of the material through the entire thickness, the prevention of failure is entirely dependent upon the strain- hardening properties of the material.

B. SECONDARY STRESS A stress developed by the self-constraint of a structure. It must satisfy an imposed strain rather than being in equilibrium with an external load.

The basic characteristic of a secondary stress is that it is self-limiting because minor distortions can satisfy the discontinuity conditions or thermal expansions which causes the stress to occur.

C. PEAK STRESS The highest localized stress in the region under consideration. The basic characteristic of a peak stress is that it does not cause any significant distortion and is objectionable mostly as a possible source of fatigue failure.

Examples of peak stress are: i. The thermal stress in the wall of a vessel caused by a rapid change in temperature or a large temperature gradient.

2. The stress at a local structural discontinuity.

VII. CYCLIC LOADING AND FAT!GUEANALYS!S All components shall be capable of withstanding cyclic loading associated with 21 turbopump assembly re-uses and 10500 sec operation.

Consideration of a stress which fluctuates about a non-zero value shall be accomplished by use of a modified Goodman diagram. The mean value of the stress used shall be an adjusted value, defined as follows(2) Let S' = Basic value of mean stress (calculated directly from mean loading cycle) including stress concentration effects° S = Adjusted value of mean stress mean (I) Criteria of Section I!I of the ASME Boiler and Pressure Vessel Code for Nuclear Vessels, ASME Publication, 1964 (2)ibid.

Page A-3 ,: /' i_£_ !' _'rl " !_ i i I A \rKX_.C _ i< < _ i : i_ b_ i <£p_ _:S'J':i:_ _ J_4 ) Sal t = Amplitude (half range) of stress fluctuation including stress concentration effects S = Yield strength Y If Sal t + S v _ Sy, S = S' mean , mean mean If Sal t + S' > S and < Sy, S = S mean y Salt mean y Salt If Salt _ Sy, Smean = 0 For determination of the allowable number of cycles, an equivalent alternating stress component, Seq, is used in entering the fatigue S-N curve. This value is defined as Salt S = eq S mean i S u where S = The smaller of the tensile strength or the 50 hr. stress u _ rupture strength.

When a complete S-N curve is not available, the relationship between S and N can be taken as follows (_ (4): S = E in i00 + S 4 _-_ " i00 - H A e where E : Elastic modulus (psi) RA : Percent reduction of area in tensile test S = Endurance limit (psi)>for i08 cycles e S : Strain amplitude times Young's Modulus Low cycles fatigue (4) 65) (6) which occurs in less than 104 cycles, caused by large plastic strain fluctuations characteristic of thermal stress fatigue may be evaluated using the above formula where S : Yield strength, 0.2{ offset e S : Strain amplitude x Young's Modulus (3) ibid.

(4) Langer, B. F , Design of Pressure Vessels for Low-Cycle Fatigue, ASME Trans.

Journ. of Basic Engineering, September 1962, pp. 389 - 402 (5) Coffin, L. F. Thermal Stress Fatigue, Product Engineering, June 1957, pp. 175 - 179.

(6) Manson, S. S., Thermal Stressq_ in D__esii_ 2art 2 - Basic Concepts of Fat__ in Ductile Materials, Machine Design, August 7, 1958, PP. i00 - 107.

Page A_4 The calculation of S is based upon the assumption of elastic behavior. However, if instability such as plastic hinges or other forms of non-!inearity exist, plastic analysis is required to obtain the strain amplitude.

When a component is subjected to a variety of stress cycles during its lifetime, failure will be taken as when the cumulative usage factor, which is the sum ____I + _2 + _3 + ..... is equal to 1.0 N I N 2 N 3 N I = Cycles to cause failure at stress level S I HI = Cycles of stress level SI applied VIII. MARGIN OF SAFETY allowable stress The margin of safety is equal to calculated stress - i For Primary_ Secondary, and Peak Stress Levels in Brittle Materials (elongation !5_), and Primary Stress Levels in Ductile Materials (elongation> 5_)_ A. MARGIN OF SAFETY BASED UPON YIELD Allowable stress is minimum yield tensile, compressive or shear stress_ yield bending modulus or yield torsional modulus at operating temperature. For welded joints, use 85_ of tensile and shear values unless test data justifies otherwise. Calculated stress is derived from proof pressure or design yield load.

' B. MARGIN OF SAFETY BASED UPON ULTIMATE Allowable stress is minimum ultimate tensile, compressive or shear stress, ultimate bending modulus_ ultimate torsional modulus, endurance limit, i0 hour stress rupture or critical column stress at operating temperature. For welded joints, use 85_ of tensile and shear values unless test data justifies otherwise. Calculated stress is stress derived from burst pressure or design ultimate load.

C. COMPOUND OR COMBINED STRESS When compound stresses and, in some cases, combined stresses are present, stress ratios may be used in the calculation of margin of safety. For particulars, see MIL-HDBK-5o For Combined Primary, Secondary and Peak Stress Levels in Ductile Material: Page A-5 The margin of safety is based upon fatigue life, and is the minimum of E.S.

M.S. - 1.4 S i eq N M.S. - i 2n where E.S. = Endurance Strength for N cycles S = Alternating Stress Amplitude eq N = Cycles of Life for S eq n = Cycles of S Applied eq Fluctuating Primary Stress Levels in Ductile Materials: The margin of safety is determined as follows (7)(8)(9)(I0)" D or i0 hr stress rupture OA : OD = yield strength, 0.2% offset, strength C _ _ __-_ / _U_ OC = Endurance strength for N cycles OB : Tensile ultimate

"__ Z /// _"

see VII. mean' _Z I _ F°r definiti°n °f Salt and S' _<_ I "_'- _ The margin of safety is:

/" "_';_ I _ _ Ox

0 j -- _ M.S. - 1.25S' 1

_( A B Mean i7) Section III, ASME Boiler and Pressure Vessel Code, op. cit.

(8) Grover, H. Jo, Gordon_ S. A°_ and Jackson, L o R., The Fatigue of Metals and Structures_ Battelle Memorial Institute_ Prepared for Bureau of Navy Weapons, Dept. of the Navy, 1954, Revised June 19607 NAVEEPS 00-25-534, pp 127 130 (9) Benham, P. P. and Hoyle, Ro_ Thermal Stress, London, Sir Isaac Pitman & Sons Ltd., 1964, pp 285 - 293 (i0 Horger_ O. J._ Editor, Metals Engineering Design_ ASME Handbook, Second Edition, New York_ McGraw-Hill, 1965, Sec. 7.2, pp 192 - 198.

Page A = 6 • / When stress-range curves determined experimentally are available_ they may be used in place of the failure lines constructed per the above technique.

IX. DEFORMATION CRITERIA All deformation which could adversely affect the performance characteristics of a component (i.e. turbine wheel tip growth_ etc.) are to be evaluated with particular attention given to creep. Acceptability of deformation magnitudes will be based upon the effects such as performance and clear_ances.

X° ADDITIONAL CRITERIh FOR ROTATING DISCS Burst speeds are to be 1o4 times greater than the nominal operating speed° The average tangential stress of the disc_ which is a measure of burst speed margin, shall not exceed 50_ 0 of the average ultimate material tensile strength.

XI. ADDITIONAL CRITERIA FOR PUMP IMPELLER VANES AND TURBINE BUCKETS Operating vane and bucket natural frequencies in the range of possible stimuli (i.e. nozzle passing excitation_ low order rotational speed excitation_ etc.) shall be determined and evaluated for potential fatigue failure. The margin of safety against fatigue failure shall be determined as set forth in Section IX considering the following_ Alternating Stress = .3 (Fluid Bending) (Dynamic Magnification) Amplitude where, Fluid bending stress includes a stress concentration factor appropriate to the article geometry_ and Dynamic magnification includes effects of closeness to resonance_ receptiveness of vibration mode to excitation, and damping if appropriate.

Mean Stress = Maximum combined stress level caused by centrifugal loading at shaft design speed_ fluid bending loading_ and the steady-state thermal stresses.

Page A_7

APPENDIX B TEMPERATURE ANALYSIS MOD Ii OXIDIZER TURBINE INLET MANIFOLD - BACKPLATE ASSEMBLY TABLE OF CONTENTS Pag____e I. Summary and Introduction B_3 II. Discussion B-3 A. Description of Configuration B-3 B. Selection of Insulation B-3 C. Thermal Conductivity B-4 D. Heat Transfer Analysis for Insulated Configuration B-5 Eo Heat Transfer Analysis for the Configuration without B-7 Insulation F. Heat Transfer Analysis of Main Joint and Nozzle Shroud B-7 LIST OF FIGURES No. Title B-I Backplate - Inlet Manifold B-8 B-2 Mod II Turbine Manifold, Conductivity of Insulation B-9 B-3 Oxidizer Turbine Mod II Inlet Manifold Wall Temperature vs. Time B-IO B-4 Mod II OTPA Turbine Manifold - Backplate Assembly, Outer Joint B-II B-5 Oxidizer Turbine Mod I! Inlet Manifold Outer Corner Temperature B_I2 VSo Time B-6 Mod II OTPA Turbine Inlet Manifold, Inner Joint Backplate - B-13 Manifold B-7 Oxidizer Turbine Mod II Inlet Manifold Inner Corner Temperature B_I4 VSo Time B-8 Oxidizer Turbine Mod II Inlet Manifold Outer Corner Temperature B-15 vs. Time B-9 Mod II OTPA Turbine Inlet Manifold, Redesign of Outer Manifold- B-16 to-Backplate Joint B-IO Oxidizer Turbine Mod !I Inlet Manifold Outer Corner Redesign Bml7 Temperature vs. Time Page B-I TABLE OF CONTENTS (CONTo No. Title B-II Mod II Oxidizer Turbine Inlet Manifold Heat Transfer B-18 B-12 Mod IT Oxidizer Turbine Inlet Manifold Heat Transfer without B_I9 Insulation B-13 Mod II Oxidizer Turbine Inl.et Manifold Heat Transfer without B-20 In sulat ion B-14 Mod Ii Oxidizer Turbine Inlet Manifold Heat Transfer without B-21 Insulation_ 2 sec B-15 Mod II Oxidizer Turbine Inlet Manifold Heat Transfer without B=22 Insulation_ 4 sec B-16 Mod II Oxidizer Turbine Inlet Manifold Heat Transfer without B-23 Insulation, 6 sec B-17 Mod ii Oxidizer Turbine Inlet Manifold Heat Transfer without B-24 Insulation_ 8 sec B-18 Mod iI Oxidizer Turbine Inlet Manifold Heat Transfer without B-25 !nsulation_ ID sec B-19 Mod II Oxidizer Turbine Inlet Manifold Heat Transfer without B-26 Insulation_ 14 sec B-20 Mod II Oxidizer Turbine Inlet Manifold Heat Transfer without B-27 insulationj 50 sec B-21 Oxidizer Turbine Mod II Main J oint Thermal Analysis B-28 B-22 Oxidizer Turbine __od II Maim 0oii{_tThermal Analysis_ Parts 1-5 B-29 and 26-28 B-23 Oxidizer Turbine Mod IilMain J oint Thermal Analysis_ Parts B-30 5-13 and 22-24 B-24 Oxidizer q / arbine Mod i- Main Joint Thermal Analysis_ Parts 14-21 B-31 B-25 Oxidizer Turbine Mod i! Maim Joint Thermal Analysis_ Parts 29-33 B_32 B-26 Oxidizer Turbine Mod Ii Mai_c J oint Thermal Analysis_ Parts 34-,39 B=33 Page B=2

APPENDIX B Prepared by C. E o Klessig I. SUMMARY AND INTRODUCTION The inlet manifold-backplate assembly of the Model I! oxidizer turbine is exposed to temperature extremes of -320°F to lO00°F. The pump discharge housing and the bearing housing are chilled-down before operation causing the backplate of the assembly to be chilled to -320°F. During operation of the turbine, hot gases of lO00°F pass through the inlet manifold causing it to heat up rapidly.

However, the backplate stays at approximately -320°F resulting in a large thermal gradient in the two short structural members joining the backplate and inlet manifold. To reduce this thermal gradient, the manifold was lined internally with an insulation.

Without insulation, the .lO0-in. thick spherical manifold wall will reach the gas temperature of IO00°F in 20 sec. Insulated with i / 2-in. Thermoflex insulation, this same wall will reach -230°F in 20 sec and 650°F in 400 sec_ the length of a normal run. The insulation not only reduces the final tempera- ture reached by the wall but greatly reduces the rate of temperature rise allow- ing adjacent areas time to warm up, thereby reducing thermal stresses.

II. DISCUSSION A. DESCRIPTION OF CONFIGURATION The backplate is a box-section with one wall common to both back- plate and inlet manifold (see Figure B-I). The inlet manifold is internally insulated with a ceramic fiber blanket. A .063-in. thick sheet metal liner is placed over the insulation for retention and to protect the insulation from high velocity gas. In turn, the liner is held in place with a i / 4-ino pin and washer combination.

The areas of thermal interest are the two corners where the hot walls of the inlet manifold join the cold backplateo B. SELECTION OF INSULATION Many insulations were investigated. Min-K and Thermoflex were the best two choices. The others were eliminated because of: i. High conductivity 2. Installation difficulties 3. Ablative - particles pass through turbine 4o Too heavy Page B-3 5o Not reuseable - good for one run only 6o Not good for required temperature extremes Min-K has a lower cond.uctivity than _ermoflex _der normal atmospheric conditions; however_ in a hydrogen-water vapor atmosphere at 200 psia it loses its advantage as the conductivity of the gas controls. Because Min-K must be pre- formed before installation and tends to powder under vibration 7 _ermoflex was selected for insulating the manifold° Thermoflex is composed of a ceramic fiber mating and is opaque to radiation.

C. _ERMAL CO_UCTIVI_ The effective conductivity of the H 2 and H20 vapor m ixture plus Thermoflex ranges from 1.15 to 1o55 B_ / hr Ft2°F / in. depending upon te m perature as shoal in Figure B-2. The other conductivities shown in Figure B-2 are: Min-K in air 7 _ermoflex in air 7 gaseous hydrogen_ and H 2 and H20 vapor.

Re conductivity of the H 2 + H20 vapor and Thermoflex was based upon an equation for porous material.(1) i -b i ----_P K s K : K a s 1 +.b_ K - Apparent Conductivity a Ks - Conductivity of H 2 plus H20 Vapor K - Conductivity of Insulation P a - 3 Ks (2K + K ) _' ' s p b - Ratio of Insulation Volume to Total Volume The above equations do not i n clude i° _e conductive heat transfer provided by ' the insulation shield retaining pins° 2. Allowance for co::wective heat transfer caused by gas flow between the insulation and the manifold wa_!Is or within the insulation.

A thermal conductivity was calculated with pins and insulation in parallel heat flow.

(i) Jakob 7 M._ Heat Tra:o.sfer_Volume 17 194-97 pp 83 - 85 Page B-4 The result of 2.34 BTU / hr ft2°F / in, was rounded off to 2.7 BTU / hr ft2 °F / in. to include convective effects in and about the insulation. All subsequent calculations are based upon an effective thermal conductivity of 2.7 BTU / hr ft 2 °F / in.

Do HEAT TRANSFER ANALYSES FOR INSULATED CONFIGURATION To determine the effectiveness of the insulation_ heat transfer analyses of the walls and the corners of the inlet manifold were made using the following three assumptions: i. The insulation liner reaches IO00°F immediately.

2. The insulation has no heat capacity.

3. No heat loss from the uninsulated side of the walls.

A temperature versus time analysis for the three walls of the inlet manifold was made. The walls were considered infinitely long and wide so it may be analyzed as a one dimensional heat transfer. The walls were initially at -320°F. Charts were used to determine the temperature.(2) The results of how the temperature varies with time are shown on Figure B-3. These temperatures were used for the end points of the manifold wa!is_ such as H 6 and M 7 of Figure B-4_ for the numerical analysis.

The corners were analyzed by numerical techniques. This procedure is to divide the area to be analyzed into sections as shown on Figure B-4. For each section_ a heat balance is written.

Example - Figure B-4_ Section M I.

dTl KA KA M1Cp _-= (--L)2,1 (T2-T1) + (--P4,1 (T4-T1) + (_)g,l (Tg-TI) - (--_)i,o (T1-To) Rearrange • C AT1 = [(_)2_1 (T2-T1)+ (-_)4 , 2(T4-Tl) + (hA)g , l(Tg-T1) - (-'_)I , o(TI'To)]_IAt P T = Temperature °F AT = Temperature Increment At : Time Increment _2) Heisler, M. P._ Transactions of the ASME_ April 1947, pp 227-236 Page B- 5 K = Thermal Conductivity in BTU / SEC in2 °F / in.

h = Hea% Transfer Coefficient BTU / SEC in2°F • 2 A = Area mn L = Distance Between, inch M = Weight Section, ib C = Specific Heat BTU / Ib P This same equation can be derived for each section. With the above equation_ the change in temperature of a section can be determined for a small increment in time. Using initial temperature as a starting point, a temperature versus time history can be obtained by adding the temperature and time increment to the initial temperature and time, then repeat adding the temperature and time increments to the previous temperatures and times until the desired time is reached.

The inlet manifold outer joint was analyzed starting with an initial temperature of -320°Fo The corner was sectioned as shown on Figure B-4. The results, see Figure B-5, show that the insulation greatly reduces the temperature in the corner and walls• Without insulation, the thin spherical wall (point MT) reaches IO00°F in 20 sec. The insulation not only reduces the temperature but allows the adjacent areas to warm up reducing the thermal gradient.

The inlet manifold inner joint was sectioned as shown on Figure B-6.

The corner was analyzed using an initial temperature of -320°F. It can be seen on Figure B-7 that the temperatures in this corner stay mader 200°F up to 400 sec, the length of a normal run.

The outer joint was reanalyzed with different initial temperatures.

The backplate and backplate closure were left at -320°F, the spherical shell was changed to 70°F and section M 4 has a linear gradient from 70°F to -320°F. It was believed that this was a more realistic initial condition° The thermal conductivity used was from Figure B-2 and the curve for H 2 + H20 and Thermoflex which is a deviation from the otherwise consistently used K = 2.7 Btu / hr ft 2 °F / in.

This resulted in much lower tempera%ures as seen on Figure B-8. The change is mainly the result of the reduced thermal conductivity° A configuration change (see Figure B-9) was made to the spherical shell where it joins the corner to ascertain if it would reduce the gradient in this area. The analysis was made under the same conditions as the first analysis for comparison purposes. Comparing Figure B-IO with Figure B-5, it can be seen that the temperature at points Mh and M 5 have dropped some 200 degrees F. with the new configuration.

Page B-6 The above studies show that the thermal gradient must be taken into consideration when determining the stress in these parts. Should the stress get too high because of thermal stress, configuration changes such as the suggested redesign of the outer corner will help reduce the thermal gradient.

E. HEAT T_&NSFER ANALYSIS FOR THE CONFIGURATION WITHOUT INSULATION Initial Temperature Backplate and Backplate Closure -320°F Spherical Shell 70°F with a one-inch section having a linear gradient from 70°F to -320°F.

The analysis without insulation was made to ascertain if short run tests could be made without insulation. The results shown in Figures B-If through B-20 indicate that even for runs of i0 sec_ very steep thermal gradients are obtained.

F. HEAT TRANSFER ANA L YSIS OF MAIN JOINT AND NOZZLE SHROUD An analysis was made of this area to ascertain if any steep tempera- ture gradients exist and how the expansion patterns would affect the stresses in the main joint weld. The analysis set-up is shown on Figure B-21 with the results shown on Figures B-22 through B-26. No steep temperature gradients were found in the main joint and the temperatures obtained did not cause any high stresses.

Page B-7

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TEH P ERATURE \ / GRA D I ENT FIGURE B-I BACKPLATE - INLET MANIFOLD Page B - 8 Backplate Flange Manifold Spherical , Shell_ .lO0

, 5

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Gas Temperatu r e i000 OF Figure B-12 Model II O x idizer Turbins Inlet }. ' [anifold-Heat Transfer Without Insulation Page B-19 Figure B-1 3 i%_odel 12[O . xidizerTurbine Inlet E_anifold- I1eatTransfer Without Insulation Page B-20 ,00 h=497 BTU / HR FT 2 °F b O @ Gas Temperat u re O 0 I000 O F FIGURE B - 14 MOD II OXIDIZER TURBINE INLET MANIFOLD HEAT TRANSFER WITH - OUT INSULATION TIME 2 SECONDS Page B-21 GAS TEMPERATURE lO00 ° F _ / I h=479 BTU / HR FT 2 ° F - I0 ( i FIGURE B-15 MOD II OXIDIZER TURBINE INLET MANIFOLD HEAT .TRANSFER WITHOUT INSULATION TIME 4 SECONDS Page B - 22 H-497 BTU / HR FT 2 ° F lO00 ° F -2 GAS T_4PERATURE - 1 0 FIGURE B-16 MOD II OXIDIZER TURBINE INLET MANIFOLD HEAT TRANSFER WITHOUT INSULATION TIME 6 SECONDS Page B-23 Gas Temperature -_, lO00 ° F h=497 BTU / HR FT 2 ° F

/

FIGURE B-17 MOD II OXIDIZER TURBINE I_LE_ MANIFOLD HEAT TRANSFER WITHOUT INSULATION TIME 8 SECONDS Page B-2_ GAS TEMPERATURE - I O C lO00°F I h - 497 BTU / HR FT2 O F FIGURE B-18 MOB II OXIDIZER TURBINE INLETMANIFOLD HEAT TRANSFER WITHOUT INSULATION TIME lO SECONDS Page B-25 I '_ _ 0 GAS TEMPERATURE lO00 ° F h-497 BTU / I-H_ FT2 ° F 4 00 FIGURE B - 19 - MOD II OXIDIZER TURBINE INLET _NIFOLD HEAT TRANSFER WIT H OUT INSULATION TIME 14 SECONDS Page B - 26 -3 0 0 00 h=497 BTU / HR FT2 ° F GAS TEMPERATURE i000 ° F FIGURE B-20 MOD II OXIDIZER TURBINE INLET MANIFOLD HEAT TRANSFER WITHOUT INSULATION TIME 50 SECONDS Page B-27 .. . 3" _- .. .-. - - .z _ , . _ ....... _- , ___ - _t - , . . - --_ _ _"_ _, _ _r-_ -- ; - -_ --4 _ ,! I __- - _.___" -- _' . ----_ - -_--__. _ _'_ ! .8 i 9 ,

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APPENDIX C STRESS ANALYSIS OF M-I M0D II OXIDIZER TURBINE _I_LET MANIFOLD AND BACKPLATE ASSEMBLY I!I_IIA _lilI_ 511¸'¸!! " _'_ii_!I IiJ ¸ TABLE OF CONTENTS Page I. Introduction C-I II. Summary of Results C=2 III. Design Criteria C-5 A. Definition of Pressure and Loading Criteria C_5 Bo Definition of Inertia and Thrust Loading Terms C_5 and Criteria Co Margin of Safety C_5 D. Calculated Stress C_5 E. Margin of Safety for Primary Stress C-6 F. Determining The Margin of Safety for The Combined C-7 Primary, Secondary, and Peak Stresses Go Low Cycle Fatigue Life C-8 IV. Material Properties C-IO V. Method of Analysis and Loading Condition C-II A. Backplate Assembly, P / N 286506 C-If Bo Manifold, P / N 286501 C_12 VI. Stress Analysis C-13 Ao Backp!ate Assembly C_13 B. Manifold and Inlet Stack C_13 Co Stress Summary and Margin of Safety C_14 VII. Conclusion C_16 LIST OF TABLES No. Title C-I Critical Stress in Backplate and Manifold C_4 C_II Mod II OTPA Backplate and Manifold Stress Summary C_14 Page C=l ]:!ASA C:i;_511 ( _?J; nd_}?_:_d:i.:-C_'-: .- Table of Contents (cont.]

LIST OF FIGURES No. Title Page C-I Elastic Range and Plastic Strain C-8 C-2 Beam Section C-IO C-3 Mechanical Properties of Inconel 718 C-17 C-4 Mod II Oxidizer Turbine Backplate and Manifold Loading C-18 Condition (Limit) C-5 Mechanical and Thermal Loads-Backplate Assembly C-19 C-6(a) Line Loads at Manifold Inlet C-20 C-6(b) Line Load Distribution at Manifold Inlet (Limit Condition) C-21 C-7 Tangential Stress Dist2ibution-Backplate Assembly C-22 C-8 Meridional Stress Distribution-Backplate Assembly C-23 C-9 Manifold Inlet at Section A (Limit Condition) C-24 C-IO Manifold Inlet at Section A (Ultimate Condition) C-25 C-If Manifold Inlet at Section A_ Limit Load + Thermal Load C-26 (Temperature at 400 sec) C-12 Manifold Inlet at Section B_ Limit Load + Thermal Load C-27 (Temperature at 400 sec) C-13 Manifold Inlet at Section B (Ultimate Condition) C-28 C-14 Manifold at Section C (Limit Condition) Cm29 C-15 Manifold at Section C (Ultimate Condition) C_30 C-16 Manifold at Section C_ Limit Load + Thermal Load C-31 (Temperature at 400 sec) Page C_2

APPENDIX C Prepared by T. Chinn_ L. W. Bartholf_ and L. K. Severud I. INTRODUCTION This is a discussion of a structural analysis of the Model II Oxidizer Turbine Inlet Manifold and Backplate Assembly_ P / N 286501.

The structures were analyzed for the most critical operating conditions° The loading conditions for the backplate assembly are pump cavity pressure_ turbine manifold pressure_ bearing load_ and temperature gradients between the hot gases in the manifold and the liquid fuelo The turbine inlet manifold loading conditions are manifold pressure_ external line loads_ and temperature gradients.

The existence of the manifold inlet ducts complicates a simple axi- symmetric solution for the manifold housing; therefore an approximate solution was used to predict the stresses at the inlet-to-manifold junction° The approximate method consisted of taking a section of the inlet and treating it as a shell of revolution. Two sections were taken_ one along the longitudinal axis and the other in the circumferential direction.

I!o S_Y OF RESULTS The critical stresses and Margins of Safety (M.So) for the Oxidizer Turbine Manifold and Backplate assembly are summarized below. The minimum Margin of Safety was determined for two stress conditions: the primary stress condition only_ and the primary plus secondary or peak stress condition. The MoS° for the latter takes into account thermal stress cycling° The critical stress condition for the backplate assembly is the result of a large temperature gradient at the manifold-to-backplate flange° in the manifold section_ the critical stress occur at the inlet-to-manifold juncture along the circumferential direction° The resulting stress in this area is largely the result of the inlet line loads° Page C-3 TABLE C-I CRITICAL STRESS IN BACKPLATE AND MANIFOLD De sign Minimum Condition Total Stress @ A M.S.

Backplate Me ch Load 120,000 psi i. 5* (I) , _ + Thermal Load

T =650°F

F = + 70 Kips Mech Load _?149,000 0 ( 2 ) Manifold Section B P = 375 psi + The rmal 146,000 °85* (I) P : 375 psi

P T : 650°F

2Fty (i) MoS. - ...... Tot_l_e_s __ -i (See Section III Design Criteria) 2) M.S. = i -i _m _b + Fry _by Page C_4 Ill. DESIGN CRITERIA The design criteria for the M-I engine states that all structural components shall be capable of withstanding limit load conditions without suffering excessive permanent deformation and without experiencing deflections which will adversely affect the performance characteristics of the engine.

The design criteria for the Mod I! Oxidizer Turbine Manifold and Backplate Assembly as well as the calculated stress are described as follows: A. DEFINITION OF PRESSURE AND LOADING CRITERIA Nominal-Maximum pressure to which component is subjected under steady state conditions.

MEOP-The maximum expected operating pressure at any time including engine transient condition.

Proof-l.2 x MEOP Burst-l°6 x MEOP B. DEFINITION OF INERTIA AND THRUST LOADING TERMS AND CRITERIA Limit Load-The critical load or combination of loads and environment the occurrence of which is expected at least once during the life of the component.

Design Yield-l.O x Limit Load Design Ultimate 1.5 x Limit Load Co MARGIN OF SAFETY allowable stress The margin of safety is equal to calculated stress -I Do CALCULATED STRESS The calculated stress can be any_ or a combination_ of the following_ io PrimaryStress A stress developed by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moment° The basic characteristic of a primary stress is that it is not self-limitingo If a primary stress exceeds the yield strength of the material through the entire thickness_ the prevention of failure is entirely dependent upon the strain- hardening properties of the material.

Page C_5 i ii_i_i i!_ _iI ¸ ...........................................

2. Secondary Stress A stress developed by the self-constraint of a structure. It must satisfy an imposed strain rather than being in equilibrium with an external loado The basic characteristic of a secondary stress is that it is self-limiting because minor distortioms can satisfy the discontinuity conditions or thermal expansions which cause the stress to occur° 3o Peak Stress The highest stress in the region under consideration° The basic characteristics of a peak stress is that it causes a significant distortion and is objectionable mostly as a possible source of fatigue failure° Eo MARGIN OF SAFETY FOR PRIMARY STRESS i. Margin of Safety Based Upon Yield Allowable stress is minimum yield tensile_ compressive or shear stress; yield bending modulus; or_ yield torsional modulus at op%_ating temperature.

Calculated stress is derived from proof pressure or design yield loado The MoSo for the combined membrane and bending stress condition at yield is: MoSo 1 (1) yield = -i _m _b Fry Fby where_ Fry = yield strength at 0o2_ offset Fby = yield bending modulus (see MatVl Property Section for calculated values) = calculated mem0rame stress m _b = calculated bending stress 2o Margin of Safet_Based Upon Ultimate Allowable stress is minim_multimate tensile_ compressive or shear stress; ultimate bending modulus; ultimate torsional modulus; or endurance limit at operating temperature° Calculated stress is derived from burst pressure and design ultimate loado (i) MIL-HDBKS___gf Metal Aircraft E!ement_ ANC_5 Bulletin Page C_6 __ _ ' _ , .... iI. i '_ e _,,i ,:_ii' I_i:'i17 '_'_,_'_' i, ....... _' '"_i",; , ') The M.S . for the combined membrane and bending stress condition at ultimate is: M. S.ult imate = i -i Cm Cb + Ftu Fbu where: Ftu : ultimate tensile strength Fbu : ultimate bending modulus Fo DETERMINING THE MARGIN OF SAFETY FOR THE COMBINED PRIMARY, SECONDARY, AND PEAK STRESS Upon satisfying the conditions of yield and ultimate for the primary stresses (general membrane, local membrane, and bending), a second condition is investigated. This condition combines the primary, secondary, and peak stresses and compares the summation to twice the yield value (2 Fry ) of the material to establish a second yield margin of safety.

M.S. 2 Fty yield = -i q p + Cs where C = calculated primary stress P = calculated secondary stress s The allowable stress value of 2 Fry is justified as follows: With reference to Figure C-I, the calculated elastic stress (primary + secondary + peak) would be S = SI = E gl o Considering the case of the secondary stress, it shall be assumed that the nature of the loading is such as to cycle the strain from zero to g! and back to zero, rather than cycling the stress from zero to SI and back to zero° When the structure is unloaded and returned to its undeflected position_ 0, the maximum stressed fibers have a residual stress (opposite to the tension or compression stress in the loaded state) of magnitude SI - Syo Page C-7 . .h. i ip- J..l

f I I

Secondary + Peak II / I

Y

t s _'_ sy_ __ / ID

Primary Stress I I

I

t o

i : > _y _l Strain g _ Strain g Re sidu@l Stress C -S -_S:.y s <sl<2 s Y _F S1>2 S Y Y Y

(a) (b)

Figure C-I On subsequent loading_ the residual stress must be removed before the stress goes into tension and thus_ the elastic range has been increased by the quantity S7 - S .

If SI = 2S_ the elastic range becomes 2S_ but if SI > 2S _ the fiber yields_as Y s_o_n by '_' i_Figure C-1(b) andallsubsequent cycles _roduce plastic strain, Therefore_ 2Sv, is the maximum value of calculated secondary elastic stress which • " 11s _ It - • .

will hake down to purely e±astmc actlono Go L OW CYCLE FATIGUE L IFE i For ' the condition of plastic strain cycling_ the structure must be investigated for fatigue° The criteria for a low cycle fatigue life under repetitive plastic action is as follows_(2) '(2) Langer, B.

Fo_ Design of Pressure Vessels for Low-Cycle Fatigue_ ASME Trans. jou_ _f Ba_ic _n_ _eptem-_ pp 389-402 Page C_8 i The plastic strain is determined by where : CT = Total Stress Thus the number of cycles to failure is given by N_ = i in 2 _ i00 - RA P where: N = No. of cycles to failure RA = Reduction of Area (i00 x _) Page C-9 IV. MATERIAL PROPERTIES The mechanical properties of INCONEL 718 as a function of temperature are presented in Figure C-3. The minimum yield and ultimate strength properties at room temperature and elevated temperature are specified per AGC 44151_ Con- dition A.

In addition to the tensile strength properties_ an allowable bending strength in the plastic range has been determined for INCONEL 718. The strengths were calculated by the following procedure.(3} Starting with the basic bending equation for a beam section (see Figure C-2) the allowable bending moment is defined as: Figure C_2 M a = Ba_ k (I / C) Eq. (i) Where: Ba, k is the bending modulus of rupture Me Eq. (2) Ba = _-- = _a + (k-l) _o,a Thus the bending modulus B for the ultimate condition is: u_k

+ _ _q.(3)

Bu,k: %,u. (k-l) o,u

where section factor k is k : 2Q__c Eq° (4) I For a rectangular cross-section, k = 1.5, thus the ultimate bending modulus is_

Bull.5 = _t,u + 0.5_ Eq. (5) o_u

Gavalis_ Ro, How to Determine Bending Strength in the Plastic Range, Machine Design Data Sheet, July 19-6_ Page C-IO The yield bending modulus is:

• sq.(6)

By, l . 5 : _t,y + 0.5 _o,y The stress C and C are plastic limits_ and are obtained from Figure 2 o,u o_y of the Machine Design Data Sheet for any material. (4) A plot of the ultimate and yield bending strength as a function of tempera- ture for INCONEL 718 is shown in Figure C-3.

V. METHOD OF ANALYSIS AND LOADING CONDITION The Mod II Oxidizer Turbine Manifold and Backplate assembly were analyzed for the following loading conditions: turbine manifold pressure; external inlet line loads; pump cavity pressure; bearing loads; and temperature gradients.

The pressure and line loads for the manifold and backplate assembly are shown in Figure C-4. The line loads at the inlet stack induce a bending moment on the inlet in the circumferential as well as longitudinal direction.

The temperature gradients for the manifold and backplate assembly were obtained from heat transfer data. The temperature gradients at the manifold-to- backplate flange junction were predicted at various operating time intervals up to 400 sec. From the available heat transfer data, the maximum temperature gradient in the manifold juncture occurs at 400 sec. This was assumed to be the most critical thermal condition. A more realistic approach to determine the most critical thermal condition would require a parametric study of the temperature gradient at different operating time; however_ because of the lack of time as a result of the M-I program phaseout, this was not accomplished.

A o BACKPLATE ASSEMB L Y, P / N 286506 The stresses in the backplate, under a pressure load, bearing loads, and thermal gradients were predicted using a finite element computer program for axi-symmetrically loaded shells of revolution.

The loading applications for the critical conditions is shown in Figure C-5. The temperature gradient represents the maximum thermal condition.

A thrust bearing load of + 70,000 !b acting upon the backplate was applied to the inner external surfac_ of the assembly. In addition_ pressures from the impeller induce a pressure distribution on the forward side of the backplate.

_7_ ib id.

Page C-ll • i_•_ • ••• i_•_ iii_il i_• I _•_ • •_ (i_ q i

B. _NIFOLD_ P / N 2865O1

The Mod II Oxidizer Turbine Manifold was analyzed in two parts; the inlet stack-to-manifold area and the manifold housing° The stress analysis of the inlet stack-to-manifold housing area cannot be accomplished without simplifying assumptions because of the asymmetry of the structure. Therefore_ an approximate_ but conservative method of analysis was used. This approximate analysis is described in the following paragraphs.

Because the line loads at the inlet are acting primarily in the longitudinal and circumferential directions_ it was necessary to investigate two stations on the inlet stack. The two stations (see Section A and B of Figure C-6(a) were investigated for the external loading conditions as well as internal pressure condition.

The shell section shown in section A must be designed to resist not only internal pressure and thermal gradient but also the longitudinal bending caused by line loads. This section is complicated by the backplate which was assumed to be "fixed _ from rotation and deflection.

The shell section of section B was investigated for the circum- ferential loading as well as pressure and temperature.

The stresses in the two sections were determined by treating the section as an axi-symmetric shell of revolution loaded by internal pressure and axial membrane load caused by the line loads (see Figure C-6(b). The line loads used in this analysis were the axial and bending loads. The bending load distribution was replaced by a uniform tension load equal to the maximum bending load / in; this is quite conservative. The shear loads were neglected.

A boundary condition for section A was applied at the backplate-to- manifold junction. _t was assumed that the stiffness of the backplate was suffi- cient to resist deflection and rotation; thus_ a fixed end boundary condition was used. This condition will provide conservative stress results.

At section B_ the membrane section of the manifold was not fixed as in Section A_ instead a resultant uniform membrane load is applied° The stresses in the manifold housing away from the inlet stacks are essentially the result of internal pressure and temperature gradient. Section C of Figure C_6(a) was used to determine the membrane stress and bending stress.

The backplate was assumed fixed_ whereas a membrane load was applied at the free end° Page C=12 The membrane, bending, and thermal stresses for the three sections were predicted by utilizing a computer program for analysis of axi-symmetrically loaded shells of revolution. The bending stresses, particularly at the discon- tinuity, are primarily the result of line loads; however, some secondary stresses are included° The temperature gradients for the three sections were for an operating time of 400 SeCo Their distribution is shown along with their respective loading conditions.

VI° STRESS ANALYSIS The predicted stress distributions for the Mod II Oxidizer Turbine back- plate inlet stack and manifold housing are summarized in Figures C-7 through C-16.

The design conditions investigated were limit load, ultimate load, and limit load with thermal gradients.

Ao BACKPLATEASSEMBLY The tangential (hoop) stress and the meridional stress distribution for the backplate assembly is shown in Figures C-7 and C-8, respectively. The resultant stresses are the result of pressure loads and temperature gradient.

The loading condition is shown in Figure C-5.

A maximum peak stress of 120 ksi, which is the meridional stress, occurs at the discontinuity between the manifold and backplate. This area also has the largest thermal gradients. The stresses (both meridional and tangential) throughout the backplate are quite small and therefore, do not present a struc- tural problem° B o MANIFOLD AND INLET STACK i. Section A The inlet stack-to-manifold was analyzed for the limit and ultimate condition as well as limit condition with temperature gradient. In this manner, the primary stress can be separated from the thermal stress. The tangential stresses, which were the most critical, are shown in Figures C-9, C_I0 and C-ll for the above loading condition.

At the discontinuity near the backplate, large bending stresses (144_500 psi) resulting from the line loads were developed for the mechanical and thermal loading condition (see Figure C-12). The maximum bending stress includes a thermal hoop stress of 86,000 psi_ which is self-limiting and is considered a secondary type of stress. Thus, the primary bending stress at a critical point is not too severe.

Page C-13

•I 2. Section B The tangential stresses at section B are more critical than in Section A mainly because the line loads are greater. The maximum stress occurs at the discontinuity between the inlet stack and the manifold housing (See Figures C-12 and C-13). The thermal stresses at this section are negligible because the temperature throughout the section is nearly uniform. Thus_ the critical condition is the primary bending and membrane stress condition. A high membrane stress of 62_500 psi was the result of the bending caused by the membrane line load.

3. Section C The critical stress condition at section C is caused by the tangential stress pressure and temperature gradients (see Figure C_16) o The maximum tangential stress occurs in the vicinity of the backplate assembly.

Because of the large thermal gradient at the discontinuity_ large thermal bending stresses (124_000 psi) were developed.

The primary membrane and bending stresses at this section are shown in Figures C-14 and C-15 for the limit and ultimate loading condition.

C. S_ESS SU_NARY A_D MARGIN OF SAFETY Table C-If summarizes the critical stress and Margin of Safety for the Mod ii Oxidizer Turbine Backplate and Manifold housing° The stresses includes membrane_ bending_ and thermal stresses. For •thermal conditions_ the stress includes thermal membrane and bending stress.

TABLE C-I!

MOD II 0TPA BACKPLATE AND MAN_FOLD STRESS SUMMARY Design Stresses @ Point A Minimum Condition Membrane Bending Total Stress MoSo Backp!ate Limit i0_000 Ii0_000 120_000 1.50" _eeh + thermal P = 375 psi _ loads _0000 ib Page C=14

r ¸ Design Stresses @ Point A Minimu_ Condition Membrane Bending Total Stress M.S.

Manifold @ Limit 9,500 58,500 68,000 1.30 Section A P = 375 psi p = 9oo lb / in.

P

A P = 495 psi n p = lO90lb / in.

_p Ultimate 16,500 67,500 84,000 1.66 Limit and 1,500 144,500 146,000 .85_ Thermal T = 650°F Manifold @ Limit 62,500 86,500 149,000 0 Section B P = 375 psi p : 27oo lb / in.

_P Ultimate 88,000 121,000 209,000 .08 p = 3830 ib / in p = 495 psi

Limit and 63,500 82,500 146,000 °85 P

Thermal

T = 650°F

Limit 16,000 54,000 72,000 I..18 C P = 375 psi Ultimate 19,500 72,500 92,000 1.4

_ p = 495 psi

Limit and 21_000 179,000 200,000 .35 Thermal T = 650°F Fry i otal Page C_15 VII. CONCLUSION The Mod II Oxidizer Turbine Backplate and Manifold housing and inlet has been investigated for the most critical operating condition; that is, mechanical load combined with thermal loads.

Using the design criteria established for the Margin of Safety, two margins of safety were reported herein. One for primary membrane combined with primary bending, and the other for the total stress which includes primary and secondary stress. The secondary stresses in this case is thermal bending.

The most critical section investigated is the inlet stack-to-manifold juncture in the circumferential direction° The M.S. for the primary stress con- dition is MoS. = Oo The biggest factor affecting the stress is the line loads which in this analysis was quite conservative.

Based upon the analysis presented herein, the M-I Mod II Oxidizer Turbine Inlet Manifold and Backplate Assembly design has adequate structural integrity to withstand the applied loads and design operating condition.

Page C=16 300 .

U

f_b (K = io5)

2 50 _ 2 OO Ftu _ 150 _ , _ F_y lO0 0 • , [ _--7------"' Z __ 200 400 600 800 I000 1.200 TEMPERATURE9 OF FIGURE C-3 MECHANICAL PROPERTIES OF INCONEL 718 Page C-17

T:_= 49,oo01b

_ = 40,000 in-lb H Line Loads _ M = 144,000 in-lb H C x o VT = 800 lb H T _c = 2,010 lb

Ma

_ M H Support c ----Manifold Inlet Line Pump Housing o ii k Pressure Distribution \ _ _ On Backplate _ Y = 1490 Housing _'f -Manifold MEOP = 205 psi _ ! _. MEOP=310 psi _ - P=O o P = 290 Axis of Revolution ckplate o Assembly Bearing Load + 70,000 ibs.

NOTE_ DESIGN PRESS (LIMIT) H P = MEOP X 1°2 TEMPERATURE DISTRIBUTION @ 400 S_Co (TEMP = °F) T=650°F

55o

34o

Pressure Dist r ibution on Backplate p = 1490ps P = 375 psi

16o

Bearing Load + 70_000 lbso p = 290 psi FIGURE C_5 MECHANICAL AND THERMAL LOADS - BACKPLATE ASSEMBLY Page C-19 T=49,O00 lb

_=4o _O00_b tO_G_ _

MC=144,000 in®Ib __0_ VT=800 '_ CIRCUMFERENTIAL DI RECT ION

I

!

I SECTION A SECTION B / FIGURE C®6a LINE LOADS AT MANIFOLD INLET Page C-20 1300 lb / in 1300 lb / in ___ _ AXIAL LOAD __

+

_- 14OO ib / in

<

2 700 lb / in 900 lb / in SECTION A SECTION B (LONGITUDINAL (CIRCUMFERENTIAL DIRECTION) DIRECTION) FIGURE C-6b LINE LOAD DISTRIBUTION AT MANIFOLD INLET (LIMIT CONDITION) Page C-21 Note: Stress in KS1 Figure Noo C_7 Tangential Stress Distribution - Backplate Assembly Page C - 22 Note_ Stress in KSI

,o

Figure No° C_8 Meridional Stress Distribution - Backplate Assembly Page C_+23

__ <in, e, Lio , // l__°n° _ ¸ ¸_

Loading Condution F lOO _- __ Inner StLrfac e M e mbrane v 50 _ Outer Surfac e °4 -i00 -150 ..........................................................................

Stress Distribut f _on Figure N0. .C_-9 i : :: C : _<,_ Manifold Inl e t at iSeStion _A : : (Limit Condition) Page C-24 I090 , / / \ / "Deflection = 0 // "po ==495495 psinsi.

L o ading C o ndition

15o

i00 H

5o

_ / _ Outer Surface _ _50 o_ Inner Surface o_ o _I00 =150 Stress Distrfbution FiguroeNoo C=lO Manifold Inlet at Section A (Ultimate Condition)

" PageC_25

Loading C o nd i ti on lO0 b 4 4 ._ ._ ® 5C -lOC -20C Stress Distribution Figure Noo C=ll Manifold Inlet at Section A Limit Load + Thermal L o ad (Temperat u re at 400 se0) Page C -- 26 2700 lb / in _ Manifold _Line _ _ p = 375psi.

Loading Condition / t---Outer Surface

1oo

_ H _ _ Membrane -_ 0

_o-50

'-d o_ nner Surface _ _lO0 _ _150 Stress Distribution Figure Noo C-12 Manifold Inlet at Section B Limit Load + Thermal Load (Temperature at 400 sec) Page C_27 3830 ib / in _nlet Line " _ ' p = 495 psi Loading Condition Sur face 100 Membrane 50 _,_ • _ 0

o -50

.N % _ -lO0 I_er Surface

-15o

-200 Stress Distribution Figure No o C_13 Ma_ifoid _nlet at Section B (Ultimate Condition)

_8

Page C_ _ ._ < _ Backplate _._._ f-_Manifold p - 75 p Loading Condition i00 Inner Surface

_ 5o

_ 0 _Membrane o o_ Outer Surface _I00 =15o ..............

Stress Distribution _ Figure Noo C_14 Manifold at Section C (Limit Condition) Pag e C_29 __ Manifold Condition 12g495 psi __Inner Surface Membrane . I00 __ "_ 50 _ m 4 ._ _ 0 r -_ o -50 °_ " _ m _ "_-_ Outer Surface _i00 -15o • Stress Distribution Figuz_e N<.., C-l_ Manifoid at Section C (Ultimate Condition) Page C-30 Backplat e - -- Manifold P = 375 psi Loading Condition

2® I

1501 Sur face i00 Membrane H 5C • m m (3 r-t r_ o o c t N q0 -I00 Outer Surface _15o -20Oj Stress Distribution Figure No o C-16 Manifold at Section C Lim._t Load + Thermal Load (Temperature a t_ 400 sec) Page C=3 1

APPENDIX D TECN_OIOGY USED !N FABRICATING THE TURBINE INLET MA_NIFOLD_ PUMP BACKPIATE_ AND _TURBINE SUPPORT STRUCTURE ASSEMBLY FOR THE FLIGHT TYPE M-I OXIDIZER TURBOPUMP Page D-I TABLE OF CONTENTS I. INTRODUCTION D-2 Iio TECHNICAL DISCUSSION D-2 A o MATERIALS D-2 B. DETAIL FORMING D-2 C. WELDING AND SHEET METAL DETAILS D-2 D. MACHINING OF DETAIL PARTS D-3 E. DETERMINATION OF WELD PARAMETERS D-3 F. ASSEMBLY WELDING OF BACKPI A TE (286506-9) D-4 G. ASSEMBLY WELDING OF MANIFOLD (286503-9) D-5 Ho WELDING OF MANIFOLD / BACKPLATE ASSEMBLY (286502-9) D-5 I. HEAT TREATMENT OF ASSEMB L Y D-6 J. HYDROSTATIC TEST AND HE L IUM LEAK CHECK D-6 Ko MACHINING OF ASSEMBLY (286501-9) D-7 LIST OF FIGURES Figure No. Title D-I Manifold Inlet - Rough D-IO AGC Drawing No. 286503 (2 sheets) D-2 Backplate_ Oxidizer Pump - Rough D-12 AGC Drawing No. 286506 (9 sheets) D-3 Manifold_ Inlet-Weldment D-21 AGC Drawing No. 286502 (2 sheets) D-4 Technical Sheet - Hydrotest D-23 AGC Drawing No. 286507 D-5 Manifold Inlet D-24 AGC Drawing No. 286501 (6 sheets) Page D-2 I. INTRODUCTION The turbine inlet manifold_ pump backplate_ and turbine support structure for the flight-weight turbopump are made as one unitized component. This component is a composite weldment made chiefly from Inconel 718 material and involves fabri- cation complexity_ precision machining_ and details requiring high surface finishes.

The unit was manufactured by the Rohr Corporation of Chula Vista_ California_ under subcontract to the Aerojet-General Corporation of Sacramento_ California.

II. TECHNICAL DISCUSSION The techniques and processes used in the fabrication of this part were primarily conventional. However_ the experience gained_ particularly in the weld preparation_ welding_ and more difficult machining operations with the relatively new Inconel 718 material_ is of interest and not necessarily limited to this speci- fic component.

A. MATERIALS All raw materials_ such as forgings_ sheet_ plate_ bar_ and tubing were certified by the supplier for mechanical and chemical specification compliance and reinspected in Rohr Laboratories° All inconel 718 forgings were procured in the rough-machined condition with a minimum of 3 / 16-in. excess material for finish machining.

B. DETAIL FORMING The majority of all detail parts were formed by conventional methods° The 286503-21 skirt_ (Figure D-I)_ was made from a rolled and welded metal cone and finish formed by explosives° The part was fully annealed prior to finish forming. Three explosive forming steps (shots) were required and the maximum charge used was seven strands of i00 grain primer cord. The forming was accomplished by the open die_ open pit process.

The 286506-7 cover_ (Figure D-2) was formed on a drop hammer using a conventional punch and die. One full anneal was required prior to finish-forming.

The tubes for the 286506-9 assembly (Figure D-2) were custom-formed using conventional tube forming equipment. A mockup of the assembly was used extensively as a guide.

C. WELDING OF S_H_ET METAL DETAILS All rolled and welded rings for the 286503-9 inlet ducts (Figure D-I) were hand gas tungsten arc welded using a suspended bead. The cone for the 286503-21 skirt (Figure D-I) was automatic gas tungsten arc welded in the conventional man- ner. All welds were x-ray inspected.

Page D-3 D° MACHINING OF DETAIL PARTS The 286506-5 bearing hub (Figure D-2) was semi-finish machined prior to welding° Holes that become inaccessible after welding were pre-drilled. Key-ways_ thread recess_ and the 0.035-in. diameter holes were electric discharge machined in the hub prior to welding.

Forgings for the 286503-3 flanges and the 286503-5 transition rings (Figure D-I) were finish machined with excess allowed for weld shrinkage.

Machining of the remaining detail parts consisted mainly of weld joint preparation.

Eo DETERMINATION OF WELD PARAMETERS Welding tests were conducted to evaluate contamination of base and filler metal as well as to determine the contributing causes of weld porosity.

io Contamination of Base and Filler Metal One test plate was welded as-received and wire brushed using an air-driven rotary stainless steel brush. The plate was degreased with double- distilled acetone prior to welding. The filler metal was used in the as-received_ chemically-cleaned condition° Radiographic inspection revealed excessive porosity throughout the 12-ino weld bead; also_ there was a crater void approximately 3 / 16-in.

deep at the termination of the weld. This crater was attributed to insufficient tailing time; by increasing the tailing time the crater was eliminated.

The surface of four test plates was machined to remove all surface oxidation and degreased with double-distilled acetone prior to welding. Filler metal was used in the as-received_ chemically-cleaned condition. Radiographic inspection showed that the weld was within the required (MiL-STD-453) specification° Based upon the above results_ the following procedures were used for all welding of Inconel 718 material: a. All surface oxidation was removed prior to welding.

b. The weld area was hand cleaned with double_distilled acetone prior to welding° Co Tailing time was extended to eliminate crater cracks at the end of a weld or tackweld.

2. Preflow of Shielding Gas Three types of tests were conducted to evaluate the effects of pre- flow torch purging. These included no pre-flow gas shielding_ 15 sec pre-flow shielding_ and 3 minutes purging with 15 sec pre-flow prior to welding.

Page D-4 Test plates welded with no pre-flow shielding prior to welding exhibited excessive porosity at the beginning of each weld.

Test plates welded with a 15 sec pre-flow shielding exhibited porosity at the beginning of the weld for approximately l-i / 2-in.

Test plates welded with a 3 minute torch purge and 15 sec pre-flow were acceptable for production welding, i Based upon the above results, 3 minute purging of manual and auto- matic torches with inert gas prior to welding was made standard procedure.

3. Inert Gas Shielding Helium and Argon inert gas shielding were evaluated for automatic and manual welding of Inconel 718.

Helium provided definite advantages over Argon in welding thick sec, tions (O.090-ino and over) because it eliminated a lack of fusion in multiple pass welding, minimized porosity_ and increased the welding speed. Arc stability was more difficult to maintain at lower currents with helium; consequently, all material under 0.090-in° thickness was welded with Argon. The transition from poor arc stability to good arc stability occurred at 65 amperes to 75 amperes.

The Linde Gas Lens was used for automatic and manual welding to elimi- nate turbulence of the gas_ which is a source of weld contamination.

4. Tack Welding In an attempt to eliminate cold spots on the root side of the weld in areas of tack weldimg, two methods of tack welding were investigated. Six weld test samples of 0o063-in. Inconel 718 material were prepared for butt welding. Three of the test samples were tack welded on the face side of the weld joint; the remaining were tack welded from the root side of the weld joint. The automatic inert gas shielded tungsten arc welding process was used at optimum machine settings for welding all specimens.

Samples tack welded on the face side of the weld joint exhibited cold spots throughout the root side of the weldo Specimens tack welded on the root side of the joint exhibited a uni- form bead free of cold spots. It was concluded that this method was to be used when- ever possible° F. ASSEMBLY WELDING OF BACKPLA_ 286_06-9 (Figure D-2) The problems encountered with this first main weldment were mainly warpage and distortion rather than metallurgical° Page D-5 flange (Figure D-I) followed by four equally spaced 3-in. long welds around the same diameter. As welding progressed, the 286503-21 shell (Figure D-I) warped inwardly.

Dry ice and shorter welds were used to complete the welding. Although the inward movement of the 286502-21 shell (Figure D-I) did not increase, there was a severe mis- match (.10-in.) of the 286503-21 shell and the 286503-5 flanges. This was corrected by depositing filler metal to the distorted area and then grinding to fair with adjacent parent metal without any sacrifice in structural strength.

Because of the inward movement of the 286503-21 shell (Figure D-I), the weld shrink allowance disappeared and the centerline-to-flange face dimension of the ducts was approximately O.150-in. short. The ducts were out-of-round and required straightening before the 286503-3 flanges (Figure D-I) could be attached, using weld buildup to compensate for the short dimension.

I. HEAT TREATMENT After completion of welding, the solution anneal and age operations were performed. The assembly was degreased prior to positioning in the heat treat fixture, which was made from Hastelloy "X" material. Hastelloy "X" has a coefficient of expansion equivalent to Inconel 718.

For temperature control, thermocouples were attached to both light and heavy sections of the assembly as well as to the fixture. The retort was of the sand seal type_ also constructed of Hastelloy "X" material. Three tubes were used to purge the inside; one routed inside the 286506-9 cavity (Figure D-2), the second inside the 286502-21 shell (Figure D-3) while the third was near the top center por%ion. The assembly was cold purged with argon for two hours prior to placing it into the furnace.

The retort was positioned in a preheated furnace which was then elevated to 1950°F_ as measured by the attached thermocouples, and held for one hour_ It was then removed from the furnace and force-cooled by fans to room temperature. Again, the retort and it contents were inserted into the furnace at 70°F and elevated to 1350°F and held for i0 hours. Then, the temperature was reduced to 1200°F and held for another i0 hours. The retort was removed from the furnace and cooled to room temperature before opening. The assembly was slightly discolored_ reflecting the presence of some impurities in the argon gas. Test specimens were included in the retort. These were analyzed and found to be within specification requirements.

Jo HYDROSTATIC TEST AnD HELIUM LEAK CHECK Successful hydrostatic test and helium leak check (per 286507, Figure D-4) of the turbine housing was accomplished in a heavy four-piece fixture.

With all elements of the test fixture in place, both cavities were pressur- ized with water to 245 psi and checked for leaks. The pressure in the upper cavity was then elevated to 375 psi and once again checked for leaks. Pressures held steady in both cavities for the specified time.

Page D-7 Automatic multiple pass welding was used in welding the 286506-3 plate to the 286506-5 inner ring and the 286506-1 outer ring (Figure D-2)° The initial welding sequence created cracks in both weld joints. This was corrected as subsequent weld passes were deposited, adding strength to the weld root° When the second unit was welded_ the condition of initial cracking was eliminated by depositing the first weld pass on the root side of the weld prior to completing the weld on the face side° Shrinkage from both weld joints moved the inner ring upward and out of position approximately 3 / 8-in. Four hundred tons were required to cold straighten the assembly on a hydropresso Welding and straightening was followed by rough machining all over_ finish machining inside the cavity_ and completion of drilling. Tubes were formed and trimmed to fit. Joints were trimmed short_ thus permitting burn-through welds to eliminate contamination pockets° Tubes were individually installed_ gas purged , tack welded_ and welded° All joints were x-rayed to determine the burn-through quality of welds and the tubes were then hydrostatically tested and leak checked with helium_ The inner cavity was cleaned and inspected prior to fitting and welding the 286506-7 cover plate (Figure D-2) o Shrinkage once again moved the inner ring upward° In the process of straightening_ the spherical radius of the cover changed to a conical contour° After straightening_ the assembly was stress relieved.

Go ASSEMBLY WELDING OF MAN!FOLD_ 286503-9 (Figure D-I) The duct assemblies were fabricated by welding five rings together° The 286503-3 flange was omitted at this time as a joint was required to correct for weld shrinkage and tolerance build-up at final assembly° Weld porosity_ which was prevalent throughout the program_ was the only source of trouble with this part.

H. WELDING OF MANIFOLDA_ND BACKPLATE ASSEMBEY_ 286502-9 (Figure D-3) Problems _n connection with the final joining of the two major subassemblies were caused mainly by warpage and shrinkage. A subassembly of the 286502-11 flange and the 286502-5 cone (Figure D-3) when welded to the 286506-9 assembly_ (Figure D-3)9 resulted in a short dimension. Tb_is was corrected by adding an attached shim after final machiming.

The 286506-9 assembly_ (Figure D-3) was machined for fit-up of the shell 286502-1 (Figure D-3) allowing additional material for shrinkage° The inlet ducts were then attached to the other manifold components that were already welded to the backplate° To assure fit-up_ it was necessary to skip tack weld at increments of 3 / 4-in. around the 15.08-ino diameter of the 286503-5 Page D-6 The assembly was then drained and the test fixture removed. After drying_ the fixture was repositioned for the helium leak check. Helium was pumped into both cavities and pressurized to 125 psi_ showing no indications of leakage by mass spectro- meter check. The pressure in the upper cavity was increased to 185 psi and again showed no leaks.

Ko FINISH MA6_INING 286501-9 (Figure D-5) After inspecting the 286502-9 assembly (Figure D-3) to determine stock removal allowances and out-of-roundness_ the assembly was mounted with the open end down on the lathe fixture. With the assembly in place_ the fixture was mounted on a King Vertical Turret lathe and indicated for concentricities and flatness. By using hydraulic leveling jacks on the fixture, the assembly was positioned concentric and flat to within .001 TIR. For machining of the outside diameter, the assembly was clamped through the bore with a large washer.

During rough machining of the outside diameter / the flange face_ and the Conoseals_ the lathe fixture had to be reinforced to maintain concentricity° After rough machining_ the assembly was unclamped and allowed to stabilize before reclamping and finish-machining. The Conoseal contour was machined with a form tool to within a few thousandths of an inch of drawing dimensions_ to be finished by subsequent polish- ingo When the outside diameter was semi-finish machined_ fixture clamps were secured on the outside flange and the center clamping arrangement removed. All sur- faces were rough cut with the exception of the five degree angular face which is held to gauge point dimensions. A tracing attachment was installed and coordinated to previously machined surfaces. Single point tracing with a Valenite V-7 carbide tool with a positive rake angle proved satisfactory.

The root areas of the Conoseals were abrasive-polished to the specified 32 micro-ino finish. Other surfaces_ requiring 63 micro-in, finish were also polished.

Machining Inconel 718 in the aged condition is slightly more difficult than machining in the annealed condition° Very light cuts work-harden the material and cuts of no less than OoO!O-in. were used where possible. The material is very abrasive and requires regular cutter changes.

For the second machining operation_ the assembly was mounted on the lathe fixture with the open end up. A plug and a split ring were used to maintain round- mess of the 286503-1 ring (Figure D-I) during machining. The inside of the hub was finish-bored with exception of the extremely close tolerance 9.5460 / 9o5464-ino diameter bore which was ground on a separate set-up. The Conoseal grooves were plunge-cut with a form tool and abrasive polished to the required 32 micro-in, finish° *Registered Trademark of Aeroquip Corp.

Page D-8 _)_ C_ ;ili i / _ _il. ¸ i ) )_ii_ i_ i i _ i / ii_i'i _)_i_ _¸ _ (,_i_ ) Drilling was accomplished on a Burgmaster drilling machine. Previously conducted studies showed that the performance of Cobalt drills far surpassed high speed drills on hole sizes of I / 2-in. diameter and smaller. Broken drills were removed from the holes by electric discharge machining.

Finish boring of the venturies was accomplished on a Lucas Boring Mill, using a rotary index table for positioning. To bore the holes straight and accurate, a bushing was used as a pilot for a four-flute core drill. After core drilling, the lower end of the hole was reamed. Using this diameter as a pilot for the remaining steps_ it was possible to ream the holes and maintain concentricity. All tools were of high speed steel.

The remaining operations were of routine mode and were accomplished without incident, Because in-process inspection was practiced throughout, the final inspection was minimized. The completed part was cleaned by flushing with degreasing fluid and protective covers were installed prior to packaging and shipping° Page D-9 _/ _ --I _-_-°', ....... _ Ox ._l

i I 9o _9__I_F-_ 72 1- °

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