THE UNIVERSITY OF MICHIGAN INDUSTRY PROGRAM OF THE COL EGE OF ENGINEERING SLCTION OF ALLOYS FOR EXTROEMEPRESSURE APPLICATIONS AT ELEVATED TEMPERATURES James W. Freeman Howard R. loorhees This paper was presented before a meeting of the American Chemical Society'in Minneapolis, Minnesota., September 12-16, 1955* October, 1955 P-r13-'II7

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ACKNOWLEDGEMENT Program of the College of Engineering..

SElECTION OF ALLOYS FOR EXTREME PRESSURE APPLICATIONS AT ELEVATED TEM\PERATURES James W. Freeman and Howard R. Voorhees University of Michigan, Ann Arbor., Michigan ABSTRACT Design and selection of materials for extreme pressure applications is discussed from the metallurgical viewpoint, with emphasis on hightemperature applications where creep governs design considerations. Major factors covered include properties of~ materials and their relations to service requirements. Effects of metallurgical factors on properties are reviewed, together with anticipated behavior under complex stresses, stress concentrations., and stresses which vary during service. Applications for the so-called "tsuper-alloys11, as related' to their properties., are considered.

INTRODUC]TTON Selection of the proper alloy for any application requires matching material properties to service requirements at a minimum cost, Invariably some compromises in design and. operating conditions are necessary to obtain adequate agreement. Extreme pressures'Introduce nothing new into basic design principles, but the thick sections needed to keep stresses within acceptable limits intensify man~y of the usual problems. V14hen temperatures are increased many or all of the usual materials may become so weak that it is'impossible to build a practical vessel for high pressures from them. The materials engineer must then turn to special alloys which have been developed for retention of strength at elevated temperature. Quite exact performance of alloys can be predicted if the design and materials engineers, working together, can correctly anticipate the exact stresses, temperatures and corrosion effects, This'is however not a simple matter in extreme pressure applications. The~ heavy sections introduce complex stresses and stress gradients, and rigidity of these heavy sections intensifys'Inevitable stress concentrations, Temperature gradients give rise to thermal stresses which complicate stress analysis. Finally, most of the data available for the strength of alloys were determined for static uniaxial tension at constant temperature. In practice pure tensile stresses are rare and the simple tension properties must be adapted to complex bi'axial'and triaxial loads, stress gradients over the section, varying stresses and varying temperatures. Most operations at extreme pressures involve use of alloys under conditions where little or no prior experience exists, In moderate-pressure applications a wealth of experience is available on which to base expected performance of materials. Because it is difficult to anticipate service stresses and temperatures exactly, the absence of experience is a real handicap. The primary requisite of any material is adequate strength. The first step in selection of materials is therefore to check the strengths of available alloys. Basic requirements of design and selection of materials are ei'ther to avoid actual fracture or to prevent any part from becoming inoperative due to excessive deformation within the expected service life, At ordinary temperatures this is accomplished by limiting the stress to the yield strength or by using some fraction of the ultimate tensile strength or yield strength which experience has'Indicated to allow adequately for unpredictable stress concentrations and variation'in alloy properties, As is well known, the ASMK Unfired Pressure Vessel Code uses one-quarter of the ultimate tensile strength and either 5/5 or 2/3 of the yield strength for this purpose, At high temperatures a new phonomenon is encountered'in addition to reduced strength from temperature alone: continuous plastic deformation or "1creep"' at constant stress, When a typical alloy'is first loaded at elevated temperature, plastic deformation for a given stress level begins

still later times the creep accelerates ("third stage") until fracture of ture under stress at elevated temperatures requires modificatior of the

HIGH-TEWPEIRATUftE TEST DATA Tests on high-temperature alloys are set up to obtain data the engineer feels are related to particular service requirements. These are most commonly tests'in simple tension under constant load and temperature. In different applications the amount of creep which can be tolerated varies. Also, expected service time ranges all the way from a few minutes to design lives of 25 to 40 years for central power stations. Published information tends to vary considerably and seldom covers all possible ranges of stress, time, temperature and amount of creep. Unfortunately, the metallurgist has no sure way to establish performance at conditions remote from those tested. Stress-Rupture Strength The time until fracture under creep conditions is dependent on the stress level. At constant temperature the relationship between the'Initial applied stress and fracture time can usually be shown as a straight line,, or intersecting straight-line segments, when plotted to logarithmic coordinates. The resulting "~stress-rupture time" curves (See Figure 1) are a timedependent measure of the ultimate strength. On these plots the elongations measured on the fractured specimens are often'Included as additional information of value. Common practice is to conduct actual tests out to about 1000 hours and then to extrapolate to 10,000 and. 100.,000 hours. These extrapolations can be very reliable if a family of curves is established at several temperatures including at least 2000 F. above the temperature of interest, Reasonable similarity of slope for the stress-rupture time curves is evidence of reliability at the lower temperature. A number of curves covering the range of useful temperature for an alloy describes the time dependent ultimate strengths quite completely. Rupture strengths are usually presented as the stresses for fracture in 1, 10, 100,~ 1000, 10,000 and/or 100,000 hours. Design_ Dta Curves Rupture strengths have the same limitation as does tensile strength at ordinary temperatures: they do not define the stress-dependency of limited amounts of strain. In many cases only a limited amount of deformation can be tolerated before a part becomes inoperative. At high temperature strain is limited through use of data establishing the relationships between stress and creep as a function of time. A series of different constant-load tests is run and the creep measured. The results are conveniently presented as curves of stress versus log time for total strains of 0.1, 0.2, 0.5 and 1.0 per cent, the strains which seem to most useful to designers (See Figure 2). These total strains are defined as the elastic strain on loading Plus any

Creep Strengths Direct determination of limited-deformation design curves is usually carried only to about 1000 hours and there is at present no reliable way of extrapolating to longer times. The strength for limited deformations in prolonged time periods'is usually estimated from the stresses for secondstage creep rates of GOoOQOl1 and 0.00001 per cent per hour determined from tests of about 1000 hours duration. Three or four tests are conducted in this range of rates and, a stress-creep rate curve plotted for'Interpolation (Figure 3). It is customary to extrapolate these rates as follows: 0'.0001%/hr 0.1%/bOG hours = l%/l0,000 hours,, 0.00001%/hr 0.01%6/bOG hours = 0.l%6/l0,000 hours = l%/l00,000 hours. Even assuming that the second-stage creep rate remains constant, it is evidn tathtoasrinis greater than indicated by that rate alone, due to deformation on loading and the higher rates of first-stage creep. Whether or not corrections are made for the added. strains prior to the start of second-stage creep, creep strengths are basically a measure of load-carrying ability for limited deformations and prolonged time periods. The two values quoted. are those most widely used by engineers. Intermediate values or values for other time periods can be established by similar treatment'if sufficient test data are available. Existing Data A set of design curves., together with creep data., for one temperature requires from 9,000 to 12,000 hours of testing time, more than a year's use of a test unit. Only when an alloy has well-established potential use., and well-standardized production techniques,, are design curves over a temperature range justified. Consequently, most Qf the data available represent some portion of a complete evaluation. Rupture data are quite commonly available. Creep strengths of considerable reliability are available chiefly for alloys which have been employed in large tonnages for many years, as in the power-boiler and petroleum-refining fields, Data for stress versus time for total deformation a-re relatively rare. When test data from several sources and for different lots of a given alloy have been obtained, a range in properties is always found.. For a spcfc treatment this stems from unavoidable variations In prduction procedures and inherent variables'in testing. Often a given alloy will be heat-treated in different ways for different uses, with resultant variation in reported strength properties.* For these reasons data compilations are being prepared. to provide an indication of expected spread in properties (1). There'is growing'Interest in surveys of this type and in statistical evaluation of alloys produced within commercial specifications. USE OF HIGH-TE4PERATURE DATA

conditions.* Several common problems will be discussed in terms of existing knowledge. Many fundamental questions will be seen to still await answers, and some of the procedures to be suggested are approximate. Complex Stresses By complex stressing is meant the simultaneous action on a small element in a structure of stresses in different directions. When any one force is exerted on a body'it may act to pull out or push in on the surface, or I't may tend to produce sliding ("shearing") parallel-to the surface. When the force meets the surface at any angle other than 900 or 00, both types of action are produced and the force may be resolved into a normal (tensile or compressive) component and a shearing component.' Regardless of how many forces act at a point, or in what direction, it is always possible to find three mutually perpendicular directions'in which the shear stress is zero so that the only stress'in these three directions are pure tension or compression. These three stresses (Sl >3S2> S ) include the maximum and minimum normal stresses at the point and are calted "principal stresses"1. By convention., tensile stresses are denoted as positive and compressions as negative stresses, Pressure inside a cylindrical vessel exerts its major effect as a tension in the circumferential or hoop direction, tending to split the wall along its length. For a thin wall this stress (S) equals the internal pressure (p) multiplied by the ratio of the cylinder diameter (D) to twice the wall thickness (t):- S = plD/2t. When the ends are closed by movable pistons, the pressure exerts no axial pull so that the intermediate princi'pal stress is zero. The principal stress in the radial direction is zero at the outer surface and. equals the pressure stress -p at the'inner surface. Under this condition and with a veryr thin wall only the hoop stress is significant and the effect on the metal is the same as'in a sheet stressed by a straight pull along a pair of opposite edges. If the pistons are replaced by caps fastened to the cylinder ends, an axial pull on the walls is added., equal to half the hoop stress for the particular case of a very thin circular cylinder. Such a stress pattern., corresponding to simultaneous pulling along the length of a sheet and across its wdthrepresents biaxial loading. In thickerm vessels the axial stress becomes less important and the compressive stress of the pressure against the inside wall assumes a magnitude of the same order as the axial stress, The most general pattern -- triaxial stressing -- is thus obtained near the bore of a thick cylinder at high pressure. This corresponds in simultaneous action of forces on all six sides of a block. Theories of Failure Under COin lex Stresses A number of relationships have been proposed to derive from the three principal stresses a single equivalent stress giving the same effect as an equal stres~-in pure tension, This enables one to use data from pure tension to determine -yield, fracture, fatigue, creep and stress-rupture under complex

evidence. In historical order of their appearance,, the three theories are: (1) Maximum Principal Stress: The largest of the principal stresses (a)IS postulated to be the only- one of importance., with failure or creep independent of stresses normal to the direction ofS. (2) Maximum Shear Stress:- The critical stress combination is taken to be the largest shearing stress present', (S 83 )72. (3) Shear-Stress Invariant: The equivalent stress ~ is computed from the sepa-rate principal stresses by- the relation: V= 1/2 [(Sil S2 )2 + (Sil S 3)2 + (S S3)2] This theory, generally- attributed to von Mises, goes by- many- names including "Octahedral Shear Stress",9 "Distortion Energy-" and "'Maximum Shear Strain Energy-". Results of Coaiiplex-Stress Studies For ductile alloy-s y-ielding, fracture and fatigue at room temperature all seem to depend on a shear mechanism with generally —better correlation by- the shear-stress'Invariant theory than by- the maximum shear stress. Brittle materials, such as cast iron., follow the criterion of maximum principal stress for fracture and fatigue (2-7). The technical literature also reports studies (8-11) on a variety- of alloy-s under different patterns of complex stress at temperatures where creep occurs. For these tests the elastic limit., creep behavior,, and shorttime stress-strain relations (some carried to fracture) were all definitelynot a direct function of the largest principal stress. Satisfactory corre. lation could usually- be obtained'in terms of either the shear-stress'invariant or the maximum shear stress, with slight favor for the former. Published findings for rupture after creep under combined stress are limited to a series of five. tests on a O.5% Mo steel at 55Oo C (10210 F) and three test points for commercial-purity- copper at 2500 C (4820 F). (See Ref. 12). Tn these particular tests the criterion of failure in stressrupture appeared to be closely- that of maxi.mum principal stress. This finding was quite unexpected'in view of past observations that the shear-stress invariant theory applied during loading and for both first and second stages of creep. Further studies are required before one may- set down a final conclusion as to the general criterion for stress-rupture under complex stresses. It is tentatively recommended that both creep and stress-ru ture- pr erties.I 2 -111 of ucil alo -s unde comple stresses_ beclcuate fro sml

Non-Uniform Stress Distributions Not only are the stresses imultiaxial in pattern in a thick pressure vessel, but they are also non-uniformly distributed, In fact, the chief characteristic of extreme-pressure equipment which distinguishes it from most other'is the steep stress gradient through the walls. Stress gradients are even steeper and more localized in the vicinity of notch or a sudden change in cross section. If a notch is introduced into a member submitted to tension perpendicular to the notch, the stress'in the direction of the tension will rise sharply near the notch root, Local stresses'in other directions will also develop, but of smaller magnitude. Dluring elevated.-temperature service., stress gradients of the nature described tend to level out by creep., with the result that the actual fiber stress at any portion of the vessel wall changes with time, even in service at constant applied pressure and temperature, The higher the temperature, the faster redistribution of stress should occur due to the rapid'increase in creep rates for a. given stress with temperature, Notched.-Bar Rupture Tests Response of a heat-resistant alloy to the presence of a stress concentration is being evaluated increasingly by rupture tests on notched specimens, In some cases., notches of varying sharpness may uniformly raise the stressrupture time curve, (See Figure 4*) For other conditions,, notches have been reported to increase rupture life at high stresses and shorten life at lower stresses, Some published data'Indicate that at still lower stresses the notched specimens may once more have a longer rupture life than smooth. The curves of Figure 5'Indicate the latter possibility. For many borderline cases,, rupture life of the same alloy may be lowered by some notches and raised by others of different geometry, Results shown in Figures 4 and 5 were accumulated during studies still in progress at the University of Michigan and supported by the Materials Laboratory, Wright Air Development Center (13), For the S-816 alloy a combination of low yield strength and rapid. relaxation by creep at 13500 F makes it virtually impossible to retain a. stress concentration, At this same temperature the higher yield. point and high resistance to relaxation shorten life for Inconel X-550 under the rather severe stress concentrations studied. This difference in behavior is also evident from the times required for relaxation without elastic follow-up from the yield strength to the 1000hour rupture stress -- less than one hour for 3-816., nearly 100 hours for Inconel X-550, Vessels to operate at extreme pressures., especially on a commercial scl, are of necessity massive. Non-uni'formities of stress arising from a large ratio of wall thickness to diameter a-re present whatever the scale of equipment size, In addition., a heavy section'in a massive structure provides a greater degree of restraint, reducing its ability to flex and adjust to a localized stress concentration, Moreover, a thicker wall offers

The notch-bar rupture test provides a very useful qualitative indication of the ability of an alloy to adjust to concentrated stresses. Whenever extreme service conditions demand use of a. new alloy., or even of familiar alloys in an untried application, careful consideration should' be given to notch-sensitivity characteristics. Demonstrated ability of a particular alloy to reduce local high stresses and prolong life in a notched-bar rupture test at the expected service terra,. emure suggests that the same alloy made into a pressure vessel should be able to relax initial high bore stresses, Perhaps of greater'importance, such an alloy provides a measure of insurance against disasterous effects of unpredictable stress concentrations and inadvertent operation at conditions more drastic than planned. Creep strengths and rupture strengths generally rise and fall together. Desire for rapid creep to eliminate initial stress concentrations must be balanced against need for high rupture strength for super'-pressure appli-W cations, This dilemna serves to illustrate the constant compromises required in successful selection of materials, All other things being equal, an alloy with proven notch strengthening for a wide range of notch acuities and test conditions'is always to be pre — ferred. Should superiority'in creep-rupture strength or other property require the use of an alloy at temperatures where a notch lowers the rupture strength, special precautions must be taken'in the design to avoid the presence of sudden section changes or other stress raisers, and case must be exercised during operation to minimize thermal stresses and shock loading. When better data are absent., ductility in the rupture test can be used as a rough measure of resistance to stress raisers, Rapid decrease in ductility with increasing time for rupture is particularly indicative of probable notch sensitivity, There is a range between about 2 and 10 per cent elongation where notch sensitivity may or may not be present. But even two per cent is more than would be required to relieve even the most severe stress concentration, It. therefore, appears that ductility in rupture tests'is only qualitative in its prediction of relaxation resistance, Life of a part under an'Initial stress gradient at high temperature is intimately connected with relative creep and rupture strengths at service conditions. If an alloy is very resistant to creep the peak stress remains to control the time and point of rupture. On the other hand, for metals with low creep strength, rapid leveling of stress gradients occurs without excessive use of life during the initial period. of localized high stresses, Restraint and triaxiality at the point of stress concentration also exert a major influence, If a metal fiber is free of restraint against elongation the plastic strain of creep results in a change in length. (This is the situation'in the usual constant-load creep test.) However under complete restraint against change in overall dimensions the same fiber would only experience a drop in stress level., called tirelaxationti, All plastic creep strains

Resistance of alloys to such relaxation of stresses can be measured in a modified form of creep test'in which portions of the load are removed as creep occurs., thereby lowering the stress acting and keeping the length of the specimen constant, Complete relaxation curves showing. residual stress as a function of the elapsed time are seldom presented. Instead the residual stress at, say, 100, 1000, and 10.,000 hours are reported for a given'initial stress, Most of the relaxation data currently available have been accumulated by the ASTM-ASME Joint Committee on Effect of Temperature on Properties of Metals for publication in the near future, Relaxation data have direct application'in design of bolting and. some gaskets and cone-joint fittings where the conditions of simple loading and high degree of restraint are fulfilled, For vessel design, relaxation data give useful comparisons between materials in their ability to redistribute localized high stresses and can be used to estimate the rate of change of stress concentrations. A little consideration will show that only stress differences are amenable to relief by relaxation, Assuming the von Milses law to apply, the stress gradient subject to relaxation'in a pressure vessel free from extraneous stresses would be the shear-stress'Invariant at the bore less that at the outer surface. In triaxial stressing, plastic yielding and creep should lower the largest principal stress component fastest, reducing differences between pairs of principal stresses even faster. When a stress concentration'is characterized by a favorable degree of triaxiality it is even possible to lower the effective stress sufficiently to obtain an increase in life above that for' the case of uniform stress. The govern-Ing factor'is the amount of rupture life used up in reaching a sufficiently low effective stress to prolong life, If the alloy has high creep resistance at high stress levels, the rate of relaxation of stress concentrations will be so slow that life Will be shortened even though the internal creep strains may reduce the effective stress to quite low a value, On the other hand, a material which creeps readily and has low relaxation strength at high stress levels will allow the concentrated stresses to be reduced before an appreciable amount of rupture life has expired., and life will be prolonged by the lowered effective stress present during the large fraction of total service life, In work performed at the University of Michigan under an Air Force contract, the variable behavior of different alloys in rupture tests of notched bars under axial loading could be satisfactorily explained by such reasoning (13), In extensive tests withithree heat-resistant alloys,'it was also established that in the absence.of large metallurgical alterations fractions of rupture life are additive; ioe.*, the portion of life used up by a given period of time at a particular stress is simply the actual time at that stress divided by the rupture life in a test with that same stress throughout. Rupture will occur when the sum of all the fractions of live consumed reaches unity (140, These results have been confirmed for alloys of other types (l5)o For materials and temperatures where the addibility of life fractions

gradient can be calculated, limited only by how closely one can determine the initial stress patter. and follow the changes with time of this pattern at critical points of the vessel. Effect of Variable Temperature Pressure equipment is frequently operated' batch-.wise., with'Intermittent heating to service temperature for an on-stream period and then cooling before the next cycle. *Long-time experience indicates that for metallurgically stable materials under such operation only the period at working temperature need, be considered in estimating service life., provided the stress continues to act while the temperature is being brought up or reduced. Almost the same result is to be expected when the stress andtemperature rise and. fall simultaneously. If the stress'i's removed before cooling and not re-applied until after the temperature is brought back up, the result is usually quite different0 In this new case, metallurgical changes ("recovery") occur which reduce the strengthening that caused the declining rates of first-stage creep. This results in more creep and earlier fracture than would. be anticipated from constant temperature and stress, It is possible for very frequent temperature cycles to keep the metal always in the primary creep period, considerably reducing life. In advertent overheating'in service under stress to excessive tern.perature, even briefly, can reduce life of a pressure vessel very sharply. This results partly from the much shorter rupture life at higher temperatures where a few minutest operation can be as harmful as many hours at normal operating conditions. In addition, temperature-induced structural changes can be very damaging. Overheating can also have other deleterious effects. If restraint to expansion is present, thermal shock could easily be the worst damage. Protective oxides may be destroyed, with resultant acceleration of corrosion. Thermal Stresses and Thermal Shock In a heavy wall pressure vessel, it is difficult to avoid uneven heating and cooling. The restraint to expansion or contraction imposed by the parts of the vessel which have expanded less (heated to a lower temperature) or contracted less (not cooled as much) results in a localized stress. Restrained expansion can be a source of very high stress. Steels subjected to temperature gradients of 2000 to 3000 F per inch under complete restraint would develop stresses'in excess of the fracture strength. On the other hand., heating without restraint to expansion causes no thermal stress no matter how large the temperature change. A uniform steady temperature gradient through a vessel wall would be suibject to stress redistribution by yielding and creep relaxation at a sufficiently high temperature. Such stresses and their effects would be subec to the prnilspeiusydsusd hipratpitt

In many cases high thermal stresses are only temporarily present during rapid rates of heating or cooling before the temperature reaches a final uniform level. Such rapid nom-uniform heating or cooling in the presence of a high degree of restraint'is commonly called thermal shock. Only rarely does failure occur in one cycle. Repeated heating and cooling induces a fatigue effect, eventually causing cracks or undue warping. Thermal shock stresses are a function of the temperature gradients and the degree of restraint. Both vary over wide ranges and usually cannot be determined sufficiently well to calculate effects. A further complication is the continual change in stress and temperature as the part heats or cools, so that'it is difficult to relate these factors to metal properties. The usual recourse is to conduct thermal shock tests. The results of such tests are highly empirical and generally cover only one condition of stress and restraint. Specimens are repeatedly subjected to a specifi condition of non-uniform heating and cooling until cracks appear. Unless such tests are closely correlated to actual service conditions the results may not predict metal behavior in the proposed service, The lower the coefficient of expansion, the less the difficulty from thermal stresses, Likewise, the higher the thermal conductivity the less severe will be the temperature gradients. These two factors are the major reasons why thermal stress problems are more severe in austenitic stainless steels than'in low alloy ferritic steels. It can also be postulated that low yield strength and high ductility reduce the peak stresses by stress adjustment. Thus., as alloys become more refractory and maintain strengths to higher temperature the more severe the thermal stress problems. If ductility'is sacrificed for strength, the problem will be intensified, When equipment operates intermittently,, the period of holding at high temperature may allow recovery from the plastic deformation of thermal shock which occurred during the heating period or during cooling on a pre-~ ceding cycle. This may allow further plastic flow during succeeding cycles without fracture. It is possible to repeat such processes and finally obtain total deformations much larger than would occur in a tensile test, thus leading to more warpage than can be tolerated. Thermal stresses can be quite high at any temperature in the temperature range involved in heating and cooling. Thus,'it is important that an alloy not be unduly brittle at any temperature'in this range. *Often thermal shock failures occur at quite low temperatures because the alloy'is brittle at those temperatures or became brittle as the result of prior service at higher temperatures. In addition to the damage arising from the stress present in thermal shock, there is a largely-unknown field'Involving combined effects of temperature and plastic flow on the properties of the metal afterwards. In many cases there'is reason to suspect that loss of strength or embrittlement'introduced by the thermal shock may be of more importance than the

Fatigue Under dynamic loading fracture can occur by "fatigue" when the maximu stress is only 40 to 60 per cent of the static ultimate strength. At low temperatures superimposing dynamic stresses on large static stresses can reduce the load carrying ability as'is commonly shown by Goodman diagrams. The intermittant, strokes of the piston pumps employed for extreme-pressure applications could lower the load-carrying ability below that predicted by calculations based on a steady nominal pressure. As temperatures are increased the ratio between static ultimate tensi'le strength and fatigue strength under completely reversed stresses does not change much., However, when temperature'is sufficiently high creep makes the ultimate strength time dependent. Temperatures and time periods for rupture are eventually attained for all materials where the rupture strength is less than the fatigue strength. These temperatures become lower as the time for rupture increases. Fatigue strengths in steels for completely reversed stresses do not reach a fatigue limit at elevated temperatures as they do at ordinary temperatures. The fatigue strength continues to decrease with'increasing numbers of cycles of stress, similar to the behavior of non-ferrous metals. Most fatigue strengths a~ high temperatures are evaluated'in terms of the stress for failure in 100 cycles. Combined static and dynamic stress effects at high temperatures can best be appreciated by considering the effects of superimposing alternating stress on steady stresses. The general effect is illustrated by Figure 6. Alternating stresses superimposed on the stresses for rupture in several time periods have less and less effect on the stress for rupture in a given time period as the temperature and time period for rupture increases. At l5O0o F the alternating stress has little effect until it approaches the l5O0o F fatigue strength. Small superimposed stresses frequently'increase the steady stress required for fracture in a given time period. Apparently similar'influences occur for limited amounts of creep. Various materials differ in respect to the temperature and time periodsat which superimposed alternating stress ceases to reduce strength. Figure 6'Illustrates only the general effect but not the specific effect for all alloys. Fatigue'i's very sensitive to stress concentrations, A stress concentration does not deteriorate as much with time as'it can when static creep occurs. Increasing'temperature, therefore, does not reduce the damaging effect of a notch nearly as much as it does for static loading. On te oherhan, the beneficial effects of residual stresses'in offsetting fatigue loading is less helpful at elevated temperatures. Operations frequently used at low temperatures to increase fatigue strength which depend on residual stresses., such as shot-peening surfaces exposed to bending fatigue., lose their advantage due to creep-relaxation removing

intensify fatigue effects. Corrosion Effects Chemical reactions of a deleterious nature between the environment and a pressure vessel are difficult to predict. They range from simple ai r ox'idation through reaction with a large variety of chemical compounds which may be introduced'Into or formed by reactions taking place in pressure vessels. Specific knowledge of probable behavior'is generally required for each case, Hydrogen introduced into the metal could. cause serious embrittlement or even corrosion effects under certain circumstances. High pressures or high stresses in the metal of the vessel present additional hazards by possi'ble alterations of the corrosion from expe cted normal behavior. Each case requires careful analysis of available experience and test data for the expected corrosion effects. It'is difficult to devise laboratory corrosion tests which faithfully anticipate service conditions. In extreme-pressure applications considerable caution'is needed in applying availbecroindt nIl exeience is obtained. Certainly it'is to be anticipated that the higher pressures and high stresses will alter known corrosion characteristics, Unfortunately, relatively little test data or experience is available even under less severe service conditions for most of the high-strength alloys developed only recently. Corrosion effects may be far more damaging in extreme pressure appli-" cations than in less severe types of service. Previous discussion has emphasized the reduction in safety margins when stress concentrations are present. Thus, if corrosion causes pitting or intergranular corrosion the resulting stress concentrations can be very dangerous. Stress corrosion cracking problems certainly would be expected. to increase in extreme pressure applications. Relatively little is known about such reactions where the combination of a critical stress and a corrosive medium can cause rapid and spectacular failures with little or no external evdence of corrosion, Such effects can occur with surprisingly mild corrosive media, For example, it has recently been found that austenitic, stainless steel can be peculiarly susceptible to stress-corrosion cracking at temperatures around 6000 F where there is a water-steam interface with a slight chloride contamination in the water, Repeated stressing well within the fatigue limit of an alloy may be perfectly safe, If, however, the stress accelerates corrosion and. the corrosion introduces further stress concentrations, the ability of the material to withstand fatigue loads may be drastically reduced, Other Tps of Data Elongation and reduction of area values in rupture tests are valuable indicatorsP ofpoal efrac of an) aloy Prev Diously 11 I, th relatM..Ions - ~

uncertainties of actual servi ce conditions make it desirable to have a ductile material so that deformation may warn of aimpending failure. Ductility values are also useful in determining'if metal quality is correct. Impact tests are another useful measure of metal performance. Most alloys have high impact strength at high temperatures. Certainly, low impact strength would. be reason to be very suspicious of material for a pressure vessel, Concern is not only over response to rapidly applied stresses but also over the accompanying indication of susceptibility to brittle failure under complex stresses. If impact tests are conducted over a range of temperatures, a temperature is found where the energy absorbed falls off very rapidly. At or below these impact transition temperatures the alloy will be liable to brittle failure under shock loads and'is susceptible to brittleness under stress concentrations. Thus, for low temperatures only materials with very low impact transition temperatures, such as 18-8 stainless steel or aluminum are favored for pressure applications. Most steels have transition temperatures not far from room temperature. If a vessel is to be exposed. to high stress at low temperature, even though'it operates most of the time at higher temperature, materials with low impact strength should be avoided. Exposure to temperature and stress may embrittle many materials, Consequently tensile and impact data which do not show embrittlement developing at lower temperatures after such exposure are reassuring. It is important to recognize that such changes generally raise transition temperatures. Material with a sufficiently-low transition temperature'initially could become brittle at temperatures somewhat above normal as a result of hightemperature service, Ductility and Impact tests, as well as bend testshvutlyascek on metal quality, though the tests themselves bear little direct relation to design. Results are useful'in establishing whether the material has sufficient freedom from flaws and the proper structure to respond correctly to design conditions. USE OF PROPERTIES IN SELECTING ALLOYS The preceding section presented the principles of relating the usual strength properties of alloys at high temperatures to the more complex conditions of actual service. Because such strength properties under creep conditions are both temperature and time dependent., it is necessary to establish which ones apply for a given application. Experience'indicates that the measurement of strength used for alloy selection and proportioning of parts should not exceed the yield strength in any case. Likewise, where creep governs strength., the deformation in service should be limited to between 1 and 2 per cent during the design life for pressure vessels, It would seem necessary to avoid actual failure, thus ruling out design to the full rupture strength. Larger deformations than 1 to 2 per cent might allow instability to develop in the highly stressed vessel with consequent unduly rapid, increases'in applied stress,

hour (1 per cent per 10.,000 hours) and 0.00001 per cent per hour (1 per cent per 100,000 hours) can be used for more prolonged service and meet the requirement of limiting the deformation to 1 to 2 per cent. In practice it is recommended that both creep and rapture data be considered. The stress for 10,000 hour rupture life ought to be no more than 70 to 80% of that for a creep rate of 0.0001% per hour. The ratio for 0,.00001% per hour and. the ldO,000 hour rupture strength should be no more than 60%. Higher ratios'Indicate that one or the other strength value is in error, or that failure Will occur with very little deformation, The latter is indicative of stress concentration sensitivity. Lower values of the -ratio usually indicate that minimum creep rates were not established'in the creep tests or that the rupture strength was incorrectly extrapolated. Those experienced in high temperature testing can usually explain such discrepancies in the data so that correct appraisals can be made. Elongation and reduction of area data from the rupture tests provide additional useful information about the characteristics of an alloy. The next step is to estimate the actual metal temperatures, If the temperature is not constant, the proportions of expected life at various temperatures and stresses should be estimated so that the fraction of life used-up at the various conditions can be estimated, It turns out that a relatively small proportion of the total life at a high temperature and stress governs total life due to the far more rapid rate of use of avail-~ able life with increasing temperature. Situations'involving temperature gradients may require somewhat more complex estimates to determine the governing temperature-stress conditions, The ASNE has established allowable stresses as a function of temperature for alloys commonly used for pressure vessels (16), Whenever code requirements must be met, these stresses apply regardless of the expected life, These stresses were established for safe operation including many nontangible factors, Where creep governs strength, the stresses a re based on 100% of the 0.00001% per hour creep strength or 100% of the 100',000 hour rupture strength whichever was lower, using conservative average values for each alloy, This procedure does not allow for expected service life other than that the engineer has the option of using lower stresses where long life requirements make such procedures desirable'in his opinion, He does not have the option of using higher stresses for shorter life applications, so long as the code applies, without obtaining special approval by the Code Committee, Furthermore these allowable stresses at temperatures below the point where creep governs do not permit use of the full yield strength. They are furthermore based on minimum tensile and yield strengths permitted under applicable specifications for alloys. No credit'is allowed for special heat treatments to enhance properties at either low or high temperatures, With the criteria of strength, expected service life and temperature conditions fixed., data for available alloys can be surveyed and estimates of the dimensions of the ve:5sel developed for various alloys, This generally willAttesaetme'1 greatly narrow down the possible choics Attesm ie stimates of corrosio efetssoudbemd, rbal frhe aroig h psi

At this point, it'is necessary to evaluate the uncertainties involved. If the service stresses and temperatures and the properties of the alloys are quite accurately known, the principles outlined justify design to the full limit of the properties of the alloys* However, if considerable uncertainty in these values exists,~ most engineers tend to be conservative. If the major uncertainty involves stress concentration effects, notchedbar rupture data or ductility data in rupture tests indicating notch ije0t116i~ tivity justify ful use of the strength data, If the alloy has known notch senstitivity, the wall must be proportioned using the maximum effective stress initially present at the bore, with no allowance for any stress redistribution by creep. The usual relaxation test data can be directly applied to bolting, shrink fits, and other applications where performance depends on mainteanan~ce of elastic pre-stressing. It is as yet not possibltoapyraxin test results to stress redistribution problems. Simple relaxation correlates qualitatively with stress-concentration sensitivity. However, quanititative relationships have not yet been developed. Uncertainties in the temperatures of operation probably lead, in practice., to the most conservative design. Strengths fall off so rapidly with temperature increase that engineers must be sure that the strength i.s adequate for the highest possible temperatures. Analysis of many pressure applications at high'temperatures has shown very conservative design stresses due to temperature uncertainties, In design for extreme pressure applications, with the need for efficient use of alloys, it is vital that temperatures not only be known but also accurately and reliably controlled. The choice of the temperatures for which the strength properties will be used must be guided by how well this can be done. Consideration should also be given to the previously mentioned fatigue, thermal shock,'impact and. effects of structural changes, Difficult as it may be, thermal stresses should be included'in estimates and., if thermal shock is to be expected., how well the material will withstand it. Most of these factors are very indefinite and difficult to include firmly in an analysis. Usually all the needed information is not available, Careful inspection during service to'Insure freedom from these difficulties'is therefore necessary. In addition to design problems, a number of metallurgical problems must be considered. All alloys have ranges in properties inherent to the normal variations in chemical composition, manufacturing conditions, heat treatments and certain less tangible effects generally classified as "theat-to-heat" variations. Thus, the high temperature strength is not a single fixed value but a range for each specific alloy with a specific treatment, The problem then arises as to which values to use. Usually, the required time and cost of creep-rupture tests prohibit determination of strength values for each heat, The value of such tests is also doubtful unless the test material would be subjected to the same conditions of manufacture as the pressure vessel'Itself.

pressures, this can be dangerous. This, again., is a reason why engineers tend to be conservative in the choice of strength criteria. In many cases, parts are fabricated under production conditions and cut up for testing to obtain "typical values". The cost of such procedures., however., generally limits them to proposed high production items, In the absence of such assurance that the expected properties can be produced in a large, heavy walled-pressure vessel, it seems necessary to use conservative values. Quality problems in themselves, as well as inspection problems, increase with section size and lack of experience. The alloy producers and fabricators, therefore., have a far more difficult task in providing material known to be fre~e from defects. This is especially true where only one or at most, a limited number of parts are made, as is liable to be the case for vessels for extreme pressures. In the case of high strength heat-resistant alloys, where there may be little experience with large sized, heavy wall vessels, it may be necessary to carry on development work to obtain the necessary "know-how", Fabrication problems may become controlling factors. Many quite widelyused alloys which give no problems in ordinary use may become troublesome in larger sizes. For instance, 15-8 Gb (Type 347) stainless steel has been used for many years and enjoyed a good reputation for weldability. Yet, at the present time, welding problems'in heavy wall pipe for the newer high temperature steam plants are very serious. The alloy'is apparently subject to a hot-shortness just under the melting point which makes'it difficult to produce crack-free welds in heavy sections, Secondly., sections of the heataffected zones adjacent to the welds tend to crack in service. Apparently, these zones are deficient'in ductility and. ability to withstand thermal stresses as a result of being heated to temperatures'in excess of 20000 F during welding. Many of the so-called new "Superalloys", which may have attractive properties for highly loaded pressure vessels, can have weak brittle welds, This is particularly true for alloys which develop high strength through precipitation hardening from titanium plus aluminum additions. Care must be exercised'in considering the possibility of using many alloys to be sure that they can be made and fabricated in the section size needed. There may also be pactical restrictions in production conditions or response to heat-treatment which will result in properties considerably different from those reported in the literature for the usual tests on small-size bar stock, For'instance, it might be entirely impractical to hot-cold work a pressure vessel to enhance'Its properties although the procedure might be routine in bar stock or gas turbine rotor disks, Often it is necessary to limit solution heat treating temperatures to lower temperatures than were used to obtain test values on bar stock in order to obtain f ine grained ductile metal structures. It may be impossible to make a large size casting with adequate properties even though the alloy may be very successful as a small casting,. This is poarticularly true for those alloys

PROPERTIES OF ALLOYS A large number of possible metals and alloys might be used depending on the temperature and stress. Figure 7 has been included as a means of orientation of the possible useful temperature range for the various base metal systems. The criterion of stress for rupture in 1000 hours was used because it was the measure of strength most universally available. The comparison'is far from complete but it does indicate the temperature and stress ranges of usefulness for most base alloy systems and types of alloys. Creep strengths would shift the stress levels to lower temperatures or, at a given temperature, to lower stress levels, The comparison shows the progression of useful temperatures for base alloy systems from magnesium, aluminum., copper, titanium, carbon and low alloy steels, stainless steels, through the so-called Superalloys. Additional materials are shown as a matter of interest. Molybdenum and molybdenum base alloys maintain a very high level of strength to very high temperatures. It is now available in large sizes and a variety of forms from arc-cast ingots, However, it deteriorates very rapidly by oxidation in the temperature range where strengths are of most interest unless protected by an oxidation-resistant coating. Reliable coatings are not yet established although promising techniques are available, There are also many fabrication problems. The cermet data show the strength levels obtained'in products made from metal-refractory carbide mixtures, Cermets are highly experimental and suffer from stress-concentration and impact sensitivity. The material designated SAP (Swiss Aluminum Powder) shows the large increase in possible temperature at useful strength levels obtained by preoxidizing pure aluminum powder and then producing billets by powder metallurgy techniques.* The material can be forged, rolled or extruded-and yet retain high creep reistance, It Is still experimental with some shortcomins'h ehiu is interesting for, if comparable gains'in heat resistance could be obtained by a like technique in higher melting point metals, it would result in a major increase in useful temperature ranges, Titanium alloys are included in the comparison in view of their excellent corrosion resistance and high strength. The useful temperature range for titanium now extends to about 7000 F with indications that it may eventually be raised to 10000 F. Much has been learned about the metallurgy of titanium, It is, however, a relatively new material with a good deal remaining to be learned about its fabrication and reaction to servrice conditions, so that'Its use should be approached on an experimental basis, Titanium is particularly prone to embrit~tlement by hydrogen and any conditions exposing'it to hydroge should be carefully reviewed before it'is used. It is also very expensive, A number of the superalloys are only available as castings, In some caes these may be limited to relatively small investment castings. Most pressure vessels are made of steel, The usual materials range from carbon steel through increasing amounts of chromium up to 12 per cent combined with 065 to 1.0 per cent molybdenum. Typical creep data for these materials are included in Table 1. These alloys are usually used'in the

A number of medium-carbon low alloy steels can be heat treated to quite high levels of strength'in the temperature range up to l0000 to ll000 F. Such steels may have useful characteristics for some pressure vessel applications, The Cr-Mo-V steels maintain strength up to the highest temperatures for this group except for the complex 12 Cr steels. It should be noted that as the temperature and time period increases these steels lose their superiority to the softer annealed conditions or to the soft low-alloy steels of the previous group, The standard stainless steels represent the next step upward in ability to maintain strength with increasing temperature. Most of these steels have rather low yield strengths and this tends to limit applied stresses. There are a large number of relatively new alloys which have combinations of attractive properties for high-stress applications at high temperatures. Some of the available strength data are included'in Table 1, It should be recognized, as previously discussed, that'it may or may not be possible to fabricate pressure vessels from these alloys. To the authors' knowledge, some experience'is available with a number of these alloys such as 19-9DL, N-155% and Inconel-X. The newer most refractory alloys, such as M252, Waspaloy, Inco 700, Udimet 500, etc.,, may have severe size limitations at the present time. They are universally more difficult to machine than are the standard stainless steels. Such alloys as HS5 21 and HS 31 (X-40) have, mainly been used as relatively small investment castings. It'is to be emphasized that this list of alloys is by no means complete. The ramifications of production condition, heat treatments and properties have not been adequately covered. In those cases where properties are potentially useful, it'is strongly recommended that the possible applilcation be reviewed with the alloy producers. This should be done, in any case, for any alloy, Such groups have a background of experience and "know how" which'is essential to the success in special applications. CONCLUSIONS Principles of using available data to select alloys for vessels for extreme pressure at elevated temperature have been reviewed,. When the service stresses and temperatures are accurately known, together with properties of the specific lot of alloy used, performance of the alloy will be'in accord with predictions within practical engineering limits. Some of the properties which may be used to provide safeguards against some unknown stress conditions were discussed. A short listing of alloy properties is given. This was intended to indicate typical properties of some of the newer alloys with high strength at high temperatures, Metallurgical variables which are always present require a more thorough

1. Cormpilations of Elev-ted Tepe r-ture -'roperties of Stainless Steels (Spe. oh. sb!. c.. 12. ) Throrriilul-L~ol:/ybcden-ll ueels (pec Tech. Publ. No. 1.1) elected.Ipe r Strength oys ( -oec. Tch-.,bl o. 10 Uif. Soc. Testing;atar as'....hicielphia' erina. 2 * ^ro c r v n k c)U;s on n T.'-9> &~ eLaL a 2. &rover,::. J.; G ordon,'S.!. anc.ckson, L. o.,....u....eteas ard t-rutcLre s, ashingbon. -'. Iepa rtoen t the Navy, uzeau of te ronati s r.1., j,_. 3 Guest, J. J., hil. l ag., 5 0(5th Series)..-13, (9) 4 urner, L, Engin eering, 92, 119:-7 l'33-:J 24x<-50 5-7, (9 11) Cook, G. and Robertson, A., Engineering, 92, 786-9, (1911). 6,. Taylor, u. I. and Quinney, H., Trans. Roy. Soc. (London), Series A 230, 323-62, (1931). 7. Griffis, L. V., Morikawa, G. K. and Fraenkel, S. J., J A. Welding Soc., 27, 161, (1948). 8$ Soderberg, C. R., Trans. Am. Soc..!ech. Engrs., 63, 737-48 (19a41). 9. Johnson, A. E., Proc. Inst. Mech. Engrs. (London), 161, 432-47, (1951). 10. Johnson, A. E. and Frost, N. E., J. Mechanics Physics of Solidss, 1, 37-52, (1952). 11. Johnson, A. E., Frost, N. E. and Henderson, J., Engineer, 199, 366-9, 402-5, 457-8, (1955). -. 12. Johnson, A. E and Frost, N. E, National Physical Laboratory Symposium on Creep and Fracture of Metals at High Temperatures, Paper No. 19, London, England, (1954). 13. Voorhees, H. R and Freeman, J. W., "Notch Sensitivity of HeatResistant Alloys at Elevated Temperatures", Wright Air Development Center, Tech. Rept. 54-175, ft. 2, (To be published). 14. Voorhees, H. R. and Freeman, J. W., "Notch Sensitivity of HeatResistant Alloys at Elevated Temperatures't" Wright Air Development Center, Tech. Rept. 54-175, Pt. 1, (August, 1954). 15. Guarnieri, G. J., "Intermittent Stressing and Heating Tests of Aircraft Structural Metals"'Wright Air levelopmient Center, Tech. Rept. 53-24, Pt. I, (Nay, 195a)4. 16. ASSI Boiler and Pressure Vessel Code, Section VIII, "Rules for Construction of Unfired Pressure Vessels"1, New York, Am. Soc. Mech Engrs., 1952. 20

TABLE 1 HIGH-TEMPERATURE PROPERTIES OF SOME REPRESENTATIVE ALLOYS Chemical Composition, Percent by Weight Alloy C Mn SI Cr Ni Co Mo W Cb Ti Al V Others Low carbon steel 0.15max 0.5 0.25max 1.25Cr, 0.5Mo steel 0.l5max 0.4 0.75 1.25 0.5 2.25Cr, 1 Mo steel 0.l5max 0.4 0.5 2.25 0.5 4140 steel 0.4 0.9 0.5 1 0.2 Cr-Mo-V steel 0.3 0.5 0.75 1.25 0.5 0.25 C422 0.2 0.75 0.55 15 0.75 1 1 0.5 Stainless, type 504 o.o8max 0.6 0.6 18 9 Stainless, type 547 0.O8max 1.5 0.5 18 12 0.8 19-9DL 0.5 1.1 0.6 19 9 1.2 1.2 0.4 0.5 16-25-6 0.1 1.25 0.7 16 25 6 N 0.15 r! 17-24 CuMo 0.12 0.75 0.5 16 14 2.5 0.45 0.25 Cu 5.0 Croloy 15-15N 0.10 1.5 0.5 15 15 1.5 1.5 1.0 N 0.12 A-286 0.05 1.35 0.95 15.5 26 1.25 2 0.2 0.5 N-155 0.12 1.6 0.4 20 20 20 5 2 1 N 0.12 s-816 o.4o 1.3 0.6 20 20 Bal. 4 4 4 Inconel X 0.05 0.5 0.4 14.5 Bal. 1.0 2.5 0.8 Fe 7.0 M-252 0.15 1.0 0.65 19 Bal. 10 10 2.5 0.9 Waspaloy 0.07 0.6 0.6 19.5 Bal. 15.9 5.3 2.8 1.1 Fe 1.1 Inco 700 0.1 15 50 27 5 2.25 Udimet 500 0.08 0.2 0.2 20 Bal. 15 5.5 5 HS31 (x-4o) o.48 25 9.7 55 7.2

TABLE 1 (continued) Stress (psi) for Minimum Creep Rate of 0.0001%/Hour = l%/10,000 Hours Alloy 8000F 8500F 9000F 10000F 1100OF 12000F 13000F 13500F 1400OF 15000F 16000F 17000F Low carbon steel 19,000 11,000 5,000 1,700 800 1.25Cr, 0.5 Mo steel 32,000 26,500 18,000 7,000 4,000 2.25Cr, 1 Mo steel 40,000 14,000 8,000 4,500 4140 steel Cr-Mo-V steel C422 Stainless, type 304 28,000 20,000 13,000 8,000 5,000 3,000 2,500 Stainless, type 347 32,000 23,000 16,000 10,000 5,000 1,750 19-9DL (Annealed) 22,000 12,500 5,000 19-9DL (Hot-cold worked) 34,000 16-25-6 22,000 15,000 10,000 6,000 17-24 CuMo 24,000 16,000 11,000 Croloy 15-5N 23,200 13,500 9,400 A-286 28,000 N-155 29,000 14,000 g,oo00 S-816 29,000 20,000 11,000 Inconel X 63,000 37,000 18,000 M-252 Waspaloy Inco 700 Udimet 500 HS 31 (X-40)

TABLE 1 (continued) Stress (psi) for Minimum Creep Rate of 0.00001%/Hour = 1%/100,000 Hours Alloy o800'F 85O F 900OF 10F 1100lOOUF 1200uF 1300~F 1350OF 1400UF 1500~ F 1600~F 17000F Low carbon steel 10,8oo 6,500 2,500 1,000 1.25Cr, 0.5 Mo steel 15,000 13,100 7,800 2,800 1,500 2.25Cr, 1 Mo steel 18,000 14,000 7,800 4,200 2,000 4140 steel 10,000 7,000 2,400 Cr-Mo-V steel 44,000 14,000 C422 Stainless, type 304 20,000 15,000 7,000 4,000 2,500 1,500 1,000 Stainless, type P) 347 27,500 16,500 10,000 5,000 2,000 1,000 19-9DL(Annealed) 12,000 7,000 2,000 19-9DL(Hot-cold worked) 15,000 16-25-6(Annealed) 13,000 9,000 6,000 4,000 17-24 CuMo 19,000 10,000 6,800 Croloy 15-15N 13,500 9,000 5,400 A-286 N-155 22,000 11,000 6,500 S-816 14,000 12,000 7,000 Inconel X 53,000 30,000 13,000 M-252 Waspaloy Inco 700 Udimet 500 RS 31 (x-40)

TABLE 1 (continued) Stress for Rupture in 1000 Hours (psi) Alloy 8500F 9000F 1000OF 10500F 1100OF 12000F 15000F 1350~F 14000F 15000F 16000F 17000F Low carbon steel 22,000 18,000 11,000 6,000 2,500 1.25Cr, 0.5 Mo steel 50,000 30,000 16,000 8,000 2.25Cr, 1 Mo steel 23,000 14,000 8,500 4,000 4140 steel Cr-Mo-V steel 62,000 20,500 C422 57,000 37,000 17,000 Stainless, type 304 49,000 36,500 25,000 16,500 10,000 6,000 4,000 Stainless, type 347 49,000 35,000 23,000 14,000 8,000 4,500 19-9DL(Annealed) 32,000 12,000 10,000. 19-9DL(Hot-cold worked) 40,000 12,000 16-25-6(Annealed) 33,000 22,000 12,000 8,000 16-25-6(Hot-cold worked) 40,000 22,000 17-24 CuMo 49,000 37,000 20,500 12,000 Croloy 15-15N 33,000 19,000 9,000 A-286 47,500 22,500 8,000 N-155(Annealed) 38,000 25,000 13,000 N-155(Hot-cold worked) 55,000 S-816 55,000 30,000 18,000 Inconel X 67,000 38,000 18,000 M-252 78,000 54,000 35,000 20,000 Waspaloy 37,500 20,000 Inco 700 87,000 59,000 30,000 17,000 8,000 Udimet 500 31,000 14,000 10,000 HS 31 (X-40) 35,000 23,000 16,000 10,000

TABLE 1 (continued) Stress for Rupture in 10,000 Hours (psi) Alloy 8500F 9000F 1000~F 10500F 11000F 12000F 13000F 13500F 14000F 1500~F 1600~F 1700~F Low carbon steel 17,000 13,500 7,000 3,500 1, 000 1.25Cr, 0.5 Mo steel 39,000 24,000 11,000 4,000 2.25Cr, 1 Mo steel 17,500 10,500 5,500 2,000 4140 steel Cr-Mo-V steel 31,000 9,500 C422 49,ooo 23,000 Stainless, type 304 38,000 27,000 17,500 10,500 6,000 2,500 Stainless, type ro 347 39,000 27,000 17,500 10,500 2,500 19-9DL (Annealed) 27,500 19-9DL (Hot-cold worked) 30,000 16-25-6 (Annealed) 25,000 14,000 7,500 5,000 17-24 CuMo 45,000 32,000 16,500 7,600 Croloy 15-15N 26,500 16,000 6,500 A-286 31,000 N-155 26,000 s-816 42,000 22,000 12,000 Inconel X 50,000 30,000 12,000 M-252 Waspaloy Inco 700 Udimet 500 RS 31 (X-40)

TABLE 1 (continued) Stress for Rupture in 100,000 Hours (psi) Alloy 9000F 10000F 11000F 12000F 1300oF 13500F 14000F 15000F 16000F 1700 F Low carbon steel 1L25Cr, 0o5 Mo steel 30,000 16,000 7,000 2,000 2.25Cr, 1 Mo steel 14,000 7,500 4,000 1,500 4140 steel Cr-Mo-V steel C422 Stainless, type 304 21,000 13,000 7,000 4,000 2,500 1,750 Stainless, type 347 32,500 22,000 13,000 7,500 4,000 1,500 19-9DL (Annealed) 17,000 ro 16-25-6 (Annealed) 17,500 9,000 5,000 17-24 CuMo Croloy 15-15N 16,500 12,500 4,500 A-286 N-155 S-816 32,000 17,500 Inconel X M-252 Waspaloy Inco 700 Udimet 500 HS 31 (X-40)

80 20 29% Elongation 17% 11.5 I12,5% 8030125% 1000F 6.5% - 0 21% 2 196%5 I I I I/ tig.$-Sress- RupJtur. e Time Curves for Solution-Treated 16-25-6 Alloy. 80000- 1200 120F. 0000 - 14,, _ 130 20000 - - 1..0.. -— Solution -QuenchedI. 4 I I I2 4 6 I8 2 4 I I8 2 4 I III8 6 I0 20 50 102 103 104 1000 Rupture Time, Hours Fig. —Str- es ipt urae Time Curves for Solution-Treated 16-25-6 Alloy. 1200 F,0000 9x ELONGGdTI'ON I I I 1C11111RE I i " Il iacrEpTEST DATA jRESSRUPTURE TIMEJ 20 000 t 30 000 2 4 6 8 2 4 6 8 2 4 6 8 2 4 6 I0 100 1000 10 000 TIME,HR. rti, Pest3lyo7 DOlo t"IZoo"F ror N'/-15' RI9oY 9?,51

'" - "irrlfsi rr rir i ri |" "rr... INCONEL"X" STRESS vs MINIMUM CREEP RATE - | | | HEAT TREATEID 4 IRS. AT 2100 -F.. A24 RS. AT 1550 rF. 800010 2R0. AT 13-0' I1 I I ~=F~r MI II I I IEEP RATE R 1000 HOURS Figure 3. - Typical Stress - Creep Rate Curves

70,000 I r ---- - Notch Geometry (Inches) 60,000 Shank diam.,D 0.600 _ Diam. at notch root, d=0.424 50,000QQ I Noc Ro _ \ (a),_ Notch angle 60 ~ Notch Root How Code Type Spec. Radius,r(inches) Notched Ref. 0 -- s 40,000 o Smooth - I * Notched 0.004 Turned _ 0 Notched 0.005 Ground __ — _ _ 30,000 - Notched 0.010 Ground 2 --- * Notched 0.060 Ground 2 (a) Discontinued 1 20o000 ~0 0 I I I 6 1 1 14 s 6 t1 20,00 2 3 4 6 7 9 2 3 4 5 6 79 2 3 4 2 3 4 5 7 8 989 0.1 I 10 100 1)000 10 Rupture Life (Hr) FIG.4 - STRESS VERSUS RUPTURE LIFE AT 13500F FOR SMOOTH AND NOTCHED BARS OF S-816 WITH CONVENTIONAL HEAT TREATMENT.

100,0080,000, 60, 00in oc 50, 000 inc SMOOTH BARS 40,000 ~. 40,0001 1'''NOTCHED - EL, o " - 30,000 1 10 100 10005,000 Rupture Life - Hours Figure 5. - Stress versus Rupture Life at 1350'F for Smooth and Notched Bars of Inconel X-550 Notch Geometry Shank Diameter = 0. 600 inch Diameter of Notch = 0. 424 inch Notch Root Radius = 0. 005 inch Notch Angle = 600

~50 x 103 Testing machine A WestinghoUse c NACA-dlJestinghouse * Dynamic creep tester X Krouse + Rolls-Royce rotating cantilever be, UO 0~~~~~~~~~~~~~~~~~~~~~~~~~~ Sorintag SF-4 P4 V NEES rotating cantilever bean 0 551,50 F o Rupture tester (Univ. of Michigan) 4) ~3 - Rupture tester (Syracuse Univ.) +30 L 0 R~~~~~~~~~~~~~~~~~~~~~~~~Foorr tecmperature l,~~~ooO F(~~~~~h~ ~ 1,000" F ae ~20 1,200 0 F 00 ~10 __ 0 1_0 ________ 0 10 20 30 40 50 60 70 80 go 100 110 120x105 Mean stress, psi Figure 6.- Curves of alternating stress against mean stress for fracture in 50, 150, and 500 hours at room temperature, 1,0000, 1,2000, 1,5500, and 1,5000 F.(or N-is5boar6ock.

60 ickel -Bose Alloys (heot treoted) 1.25 0.5Mo Steel 50 - TI, e rtFe,, C, Ni, Co, Fe 4 I ioys m, (heof treoted)j:, I Iobo#%- IBos (Z0(O s worked) O ~~~~~~Tilonium~ CortrC~idge B~ross 4lloys 0 Cr, pype 304 05Mo toinlessMolybdenum ~30 ~olnsI I I \ lumrnum I\ IType316 r"4 f loy tloinless 20 Sintere n X AlProduc A/ /O Cre x Cos Mo gnesium Hoye 01 ( I`I (experimentol) [ Corbon Steel..r Type 300 500 I000 1500 2000 Temperoaure, deg Fohr Fig. 7 Stress-Temperature Curves for Rupture in 1000 Hours for Various Commercial and Experimental Alloys

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