FINAL REPORT TO WRIGHT AIR DEVELOPMENT CENTER, MATERIALS LABORATORY ON A SURVEY OF THE EFFECT OF AUSTENITIZING TEMPERATURE AND RATE OF CONTINUOUS COOLING ON THE STRUCTURE AND 7000 TO 12000F PROPERTIES OF THREE LOW-ALLOYED STEELS By K. P. MacKay A. P. Coldren A. I. Rush J. W. Freeman Engineering Research Institute, University of Michigan Project Number: 1903 AIR FORCE CONTRACT NUMBER AF 33(038)-13496 EXPENDITURE ORDER NUMBER 73512 September 15, 1955

FOREWARD This report was prepared by the Engineering Research Institute of the University of Michigan under Contract No. AF 33(038)-13496. The work authorized under this contract involved a survey of the effect of austenitizing temperature and rate of continuous cooling on the structure and 700' to 12000F properties of three low-alloyed steels. Mr. C. B. Hartley was the Wright Air Development Center acting project engineer for this project.

A SURVEY OF THE EFFECT OF AUSTENITIZING TEMPERATURE AND RATE OF CONTINUOUS COOLING ON THE STRUCTURE AND 700~ TO 1200OF PROPERTIES OF THREE LOW-ALLOYED STEELS SUMMARY The relationships between microstructures formed with various cooling rates and austenitizing temperatures and properties at 700~ to 1200~F were surveyed for three low-alloyed steels. The steels were Ni-Cr-Mo (SAE 4340), 1. 25Cr-Mo-V ("17-22-A"S), and 3Cr-Mo-W-V (H-40). Martensitic, martensitic-bainitic, bainitic, and bainitic-ferritic structures were produced by oil quenching 1-inch rounds and air cooling 1-inch and: simulated 3- and 6-inch rounds. Modifications of the oil-quenched and normalized structures of the 1inch rounds resulted when the austenitizing temperature was raised from 17500 to 21000F. The hardness level was maintained at 280-320 BHN by tempering the structures which had higher hardnesses in the as-transformed condition. The results indicated that the fully bainitic structures which were predominantly upper bainite had maximum strength over the range of testing temperature used. In general, such structures were found in the larger, normalized sections and with the higher austenitizing temperature. Since low values of ductility resulted from the higher temperature treatments, however, the best combinations of strength and ductility were obtained when the largest sections were normalized from the lower austenitizing temperatures (1750~F for SAE 4340 and "17-22-A"S, and 1950~F for H-40). Regarding the effect of varying the cooling rates of normalized bars, it was found that increases in strength occurred for all three steels as the effective bar diameter was increased from 1 inch to 6 inches. The effect of raising the austenitizing temperature from 17500 to 21000F was to increase the strength, w~ith the H-40 steel being affected the most. Ductility was lowered for all three steels as the heat-treating temperature was raised. A correlation between the structures and properties of the continuously cooled bars of this investigation and the structures and properties of the continuously cooled turbine wheels studied previously, ranged from poor for the SAE 4340 and'117-22-A'S steels to fair for the H-40 steel. The lack of better correlation was assumed to be due to (1) differences between heats and (2) variations in heattreating conditions. 2111

TABLE OF CONTENTS Page INTRODUCTION 1 TEST MATERIALS 2 PROCEDUR E 2 Continuous Cooling Transformation Conditions 2 Austenitizing Temperature 3 Hardness Level 4 Basis of Evaluation of High Temperature Properties 4 RESULTS 5 Cr-Ni-Mo (SAE 4340) Steel 6 Influence of Rate of Continuous Cooling 6 Austenitizing Temperature 7 1. 25Cr-Mo-V (" 17-22-A"S) Steel 8 Influence of Rate of Continuous Cooling 8 Influence of Austenitizing Temperature 9 3Cr-Mo-W-V (H-40) Steel 11 Influence of Rate of Continuous Cooling 11 Influence of Increased Austenitizing Temperature 11 DISCUSSION 12 Correlation of the Structures and Properties of the Continuously Cooled Bars and Turbine Wheels 12 Cr-Ni-Mo (SAE 4340) Wheels and Continuously Cooled Bar Stock 13 1. 25Cr-Mo-V (" 17-22-A"S) Wheels and Continuously Cooled Bar Stock 13 3Cr-Mo-W-V (H-40) Wheels and Continuously Cooled Bar Stock 14 General Comments on Correlation of Structures and Properties of Wheels and Continuously Cooled Bars 14 Structures of Continuously Cooled Bars 15 Cr-Ni-Mo (SAE 4340) Steel 15 1.25-Cr-Mo-V ("17-22-A"S) Steel 15 3Cr-Mo-W-V (H-40) Steel 15 CONCLUSIONS 16 BIBLIOGRAPHY 17 iv

LIST OF TABLES Table I Influence of Austenitizing Temperature, Section Size, and Cooling Medium on the Microstructure and Hardness of SAE 4340 Steel II Influence of Austenitizing Temperature, Section Size, and Cooling Medium on the Microstructure and Hardness of " 17 -22-A"S Steel III Influence of Austenitizing Temperature, Section Size, and Cooling Medium on the Microstructure and Hardness of H-40 Steel IV Rupture, Total Deformation, and Creep Data at 700~, 900~, 10000, and 1100~F for SAE 4340 Steel for Several Austenitizing Temperatures and Cooling Cycles V Rupture, Total Deformation, and Creep Data at 700, 900q, 11000, and 1200OF for "17-22-A"S Steel for Several Austenitizing Temperatures and Cooling Cycles VI Rupture, Total Deformation, and Creep Data at 700', 900~, 11000, and 12000F for H-40 Steel for Several Austenitizing Temperatures and Cooling Cycles VII Rupture, Total Deformation, and Creep Strengths at 10000 and 1100~F for SAE 4340 Steel as Influenced by Heat Treatment VIII Rupture, Total Deformation, and Creep Strengths at 11000 and 1200~F for "17-22-A"S Steel as Influenced by Heat Treatment IX Rupture, Total Deformation, and Creep Strengths at 11000 and 1200~F for H-40 Steel as Influenced by Heat Treatment v

LIST OF ILLUSTRATIONS Figure 1 Cooling Curves for the Centers of 1-inch, Simulated 3-inch, and Simulated 6-inch Rounds of "17-22-A"S Steel Cooled from 1750~F in Air. 2 Cooling Curves for the Centers of 3/4-inch, Simulated 3-inch, and Simulated 6-inch Rounds of H-40 Steel Cooled from 1950~F in Air. 3 Effect of Austenitizing Temperature on the Bainitic Grain Size of Normalized 1-inch Diameter Bars of SAE 4340, "17-22-A"S and H-40 Steels. 4 SAE 4340 Steel. One-inch Diameter Bar Stock (a) As Normalized from 1750~F, (b) As Tempered to 300 BHN, and (c) After CreepRupture Testing at 10000F. 5 SAE 4340 Steel. Simulated 3-inch Diameter Bar Stock (a) As Normalized from 1750~F and (b) After Creep-Rupture Testing at 10000F. 6 SAE 4340 Steel. Simulated 6-inch Diameter Bar Stock (a) As Normalized from 1750~F and (b) After Creep-Rupture Testing at 1 000~F. 7 SAE 4340 Steel. One-inch Diameter Bar Stock (a) As Normalized from 1950~F and (b) After Creep-Rupture Testing at 10000F. 8 SAE 4340 Steel. One-inch Diameter Bar Stock (a) As Normalized from 2100~F and (b) After Creep Testing at 1000~F. 9 SAE 4340 Steel. One-inch Diameter Bar Stock (a) As Oil-Quenched from 1750~F (b) As Tempered to 300 BHN, and (c) After Creep Testing at 1000~F. 10 SAE 4340 StecJl. One-inch Diameter Bar Stock (a) As Oil-Quenched from 1950~F (b) As Tempered to 300 BHN, and (c) After CreepRupture Testing at 1000~F. 11 SAE 4340 Steel. One-inch Diameter Bar Stock (a) As Oil-Quenched from 2100~F, (b) As Tempered to 300 BHN, and (c) After CreepRupture Testing at 10000F. 12 "17-22-A"S Steel. One-inch Diameter Bar Stock (a) As Normalized from 1750~F, (b) As Tempered to 300 BHN, and (c) After CreepRupture Testing at 1100~F. 13 "17-22-A"S Steel. Simulated 3-inch Diameter Bar Stock (a) As Normalized from 1750~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 1100~F. vi

LIST OF ILLUSTRATIONS (continued) Figure 14 "17-22-A "S Steel. Simulated 6-inch Diameter Bar Stock (a) As Normalized from 1750~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 11000F. 15 "17-22-A"'S Steel. One-inch Diameter Bar Stock (a) As Normalized from 1950~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 11000F. 16 "'17-22-A'S Steel. One-inch Diameter Bar Stock (a) As Normalized from 2100~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 1100'F. 17 "17-22-A"S Steel. One-inch Diameter Bar Stock (a) As OilQuenched from 1750~F, (b) As Tempered to 300 BHN, and (c) After Creep Rupture Testing at 1100~F. 18 "17-22-A"S Steel. One-inch Diameter Bar Stock (a) As OilQuenched from 1950~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 11000F. 19 "17-22-A"S Steel. One-inch Diameter Bar Stock (a) As OilQuenched from 2100~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 1100~F. 20 H-40 Steel. Three-quarter-inch Diameter Bar Stock (a) As Normalized from 1950'F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 11000F. 21 H-40 Steel. Simulated 3-inch Diameter Bar Stock (a) As Normalized from 1950~F, (b) As Tempered to 300 BHN, and (c) After Creep Testing at 1100~F. 22 H-40 Steel. Simulated 6-inch Diameter Bar Stock (a) As Normalized from 1950~F, (b) As Tempered to 300 BHN, and (c) After Creep Testing at 1100~F. 23 H-40 Steel. Three-quarter-inch Diameter Bar Stock (a) As Normalized from 1750~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 11000F. 24 H-40 Steel. Three-quarter-inch Diameter Bar Stock (a) As Normalized from 2100~F, (b) As Tempered to 300 BHN, and (c) After Creep Testing at 11000F. 25 H-40 Steel. Three-quarter-inch Diameter Bar Stock (a) As OilQuenched from 1950~F, (b) As Tempered to 300 BHN, and (c) After Creep-Rupture Testing at 1100~F. vii

LIST OF ILLUSTRATIONS (continued) Fig ure 26 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of SAE 4340 Steel at 900~, 10000, and 1100~F. Curves for SAE 4340 Turbine Wheels Included for Comparison. 27 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 0. 5% Total Deformation Data for SAE 4340 Steel at 900~, 10000, and 1100~F. Curves for SAE 4340 Turbine Wheels Included for Comparison. 28 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 1, 0% Total Deformation Data for SAE 4340 Steel at 900~, 10000, and 11000F. Curves for SAE 4340 Turbine Wheels Included for Comparison. 29 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for SAE 4340 Steel at 900~, 1000~, and 1100~Fo Curves for SAE 4340 Turbine Wheels Included for Comparison. 30 Influence of Cooling Rate as Controlled by Section Size and Quenching Medium on the High Temperature Properties of SAE 4340 Steel at 7000 to 11000F. 31 Influence of Normalizing Temperature on the High Temperature Properties of SAE 4340 Steel at 7000 to 1100~F. 32 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of "17-22-A"S Steel at 11000 and 1200~F. Curves for "17-22-A"S Turbine Wheels Included for Comparison. 33 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 0o 5% Total Deformation Data for "17-22-A"S Steel at 11000 and 12000Fo Curves for "17-22-A"S Turbine Wheels Included for Comparison. 34 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 1. 0% Total Deformation Data for "17-22-A"S Steel at 1100~ and 1200~F. Curves for "17-22-A"S Turbine Wheels Included for Comparison. 35 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for "17-22-A"S Steel at 11000 and 12000F. Curves for "17-22-A"S Turbine Wheels Included for Comparison. 36 Influence of Cooling Rate as Controlled by Section Size and Quenching Medium on the High Temperature Properties of "17-22-A"S Steel at 7000 to 12000F. viii

LIST OF ILLUSTRATIONS (continued) Figure 37 Influence of Normalizing Temperature on the Elevated Temperature Properties of "17-22-A"S Steel at 700' to 1200~F. 38 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of H-40 Steel at 11000 and 1200'F. Curves for H-40 Turbine Wheels Included for Comparison. 39 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 0. 5%o Total Deformation Data for H-40 Steel at 1100~ and 1200~F. Curves for H-40 Turbine Wheels Included for Compari<r -n. 40 Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 1. 0%o Total Deformation Data for H-40 Steel at 11000 and 1200~F. Curves for H-40 Turbine Wheels Included for Comparison. 41 Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for H-40 Steel at 900~ and 1100~F. Curves for H-40 Turbine Wheels Included for Comparison. 42 Influence of Cooling Rate as Controlled by Section Size and Quenching Medium on the High Temperature Properties of H-40 Steel at 700 to 1200~F. 43 Influence of Normalizing Temperature on the Elevated Temperature Properties of H-40 Steel at 700~ to 12000F. ix

INTR ODU CTION There are a considerable number of actual and potential uses at elevated temperatures for heat-treatable low alloyed medium carbon steels in jet engines and air frames. The information covering the basic principles of heat treatment for such service is incomplete. A rather wide range of microstructures and hardness levels are possible in such steels. Control of heat treatment can develop structures with a range of pearlitic and bainitic structures as well as martensite. In many cases, practical considerations dictate a continuous cooling type of heat treatment. The result is a range of structures for a given heat treatment, depending on the actual cooling rates. For instance, normalizing (air cooling) could result in structures ranging from martensite through bainite to pearlite depending on the section size being heat treated. Yet, normalizing is often catagorically stated to give the highest strength at high temperatures. The basic information relating the type of microstructure to the temperature, time and allowable deformation during service was rather meager. The indefinite statement often made is that the maximum strength shifts from martensite to bainite to pearlite as the temperature and time period is increased. Accordingly, an investigation has been in progress to establish the basic principles of relating type of microstructure to creep-rupture properties at 700~ to 12000F, the potentially useful temperature range where creep-rupture properties could govern performance, The properties of nearly "pure structures" obtained by isothermal transformation in temperature ranges for transformation to the pearlites and bainites have been surveyed (Ref. 1). The results have been correlated with the properties of structures obtained by oil quenching, interrupted quenching and normalizing forged turbine wheels (Ref. 2). The research reported herein covers two phases extending the knowledge of the relationships of microstructures to creep-rupture properties: o1 The ranges in microstructure resulting from direct cooling over a range of cooling rates down to and including air cooling of a 6-inch round bar. This continuous cooling transformation represents the most widely used method of heat treatment. Transformation under such conditions can occur over a range of temperatures with the possibility of forming mixed microstructures that influence properties. 2. The influence of, microstructural changes resulting from increasing the heat-treating temperature for austenitizing during hardening, Creep-rupture strengths are generally considered to increase with austenitizing temperature while ductility deteriorates0 However, many apparent anomalies are encountered in practice so that it would be useful to develop a better understanding of the interrelationships of heat-treating temperature and microstructure to service conditions of temperature, stress, time and amount of creep. Three types of medium carbon, low alloy, heat-treatable steels were used to check the generality of the findings: Type of Steel De signation Cr-Ni-Mo SAE 4340 1 o 25Cr-Mo-V "i 17-22-A"S 3Cr-Mo-W-V H-40

Tempering was used to reduce hardnesses to the range of 280 to 320 Brinell. Limited survey creep-rupture tests in the range of 700~ to 12000F were used to develop general trends. The results are correlated with those previously reported in References 1 and 2. TEST MATERIALS The chemical compositions of the bar stock material used for this investigation were as follows: Steel Heat C Mn Si Cr Ni Mo V W Cu... II.J _ _ __.- -.,.,., 4340 19053 0. 40 0. 70 0. 30 0, 78 1, 75 0, 26 -- 0. 12 " 17-22-A"S 10420 0.29 0. 61 0. 67 1. 30 0. 18 0. 47 0o26.. _ H-40 K2509 0.29 0.48 0.26 3.05 0.49 0.49 0. 85 0.55 0.15 PROCEDURE mne inirtial steps in Lhis phase of the program were to establish the range.of cooling conditions during which continuous cooling transformations of interest would occur, to establish the range of austenitizing temperature of interest, to prepare specimens heat treated under the selected conditions, and to select suitable survey test conditions. Continuous Cooling Transformation Conditions Consideration of the possible section sizes and quenching media involved indicated that the cooling cycles obtained by the oil quenching of 1-inch diameter bars and the air cooling of 1-, 3-, and 6-inch diameter bars should cover adequately the range of structures of major interest; that is, from martensite formed by oil quenching to bainite formed by relatively slow cooling. Since survey data were available from the previously reported work for the oil-quenched and normalized 1-inch bars for "17-22-A"S and SAE 4340 and for 3/4-inch bars of H-40, it was necessary to obtain data only for the 3- and 6-inch diameter bars. However, since only 3/4- and 1-inch diameter bar stock were available, the desired cooling cycles were obtained by the retarded cooling of the 3/4- and 1-inch round bars in a cylinder made from a low heat duty fireclay insulating brick. The general procedure was as follows: 1. To establish the cooling cycles for 3- and 6-inch rounds, the cooling cycles for 1-inch bars of "17-22-A"S and 1- and 3-inch bars of plain carbon steel were obtained during air cooling by means of a thermocouple inserted axially to the center of the bar. Cooling curves were not obtained for 6-inch rounds because of the handling difficulties and the capacity of the laboratory furnaces.

2. From the cooling curves obtained for the 1- and 3-inch rounds, the heat transfer coefficient (h) to the surrounding medium, air in this case, was calculated for the conditions existing in the laboratory. From the value obtained, the cooling curve for the 6-inch round was calculated (Ref. 3). 3. Cooling curves were obtained at the center and surface of 1-inch rounds of SAE 4340 and "17-22-A"S materials and at the same locations for 3/4-inch rounds of H-40 steel during cooling in insulating firebrick cylinders estimated to produce cooling cycles similar to the 3- and 6-inch round bars. After several trials, the thickness of the firebrick cylinder was adjusted to produce cooling cycles equivalent to the experimentally determined cycle for the 3-inch round and the calculated cycle for the 6-inch round. It was found that the cooling rate at the surface of the enclosed 1- and 3/4inch diameter bars was essentially the same as that at the center, This observation was confirmed by the uniform microstructure throughout the cross-section of the heat-treated bars, 4. After the necessary thickness of firebrick to produce the desired cooling cycle had been determined, a number of specimens were heat treated by the following procedure: a. Bar stock of 3/4- or I-inch diameter was enclosed in the firebrick cyclinder with a thermocouple attached to the surface of the bar at the midpoint. b. The assembly was inserted in a furnace at the desired temperature and held until the thermocouple indicated that the specimen had been at temperature for 1 hour, co The assembly was removed from the furnace and the test bars allowed to cool to room temperature in the insulating firebrick. 5, Samples of the actual cooling curves obtained are shown by Figures 1 and 2. Austenitizing Temperature Experience with alloys of the same type as those being studied has indicated that, at least in some instances, better than usual combinations of high temperature strength and ductility may be obtained if the austenitizing temperature is just below the temperature at which coarse-grained, bainitic structures develop, However, this generalization seems to be most applicable to those steels containing carbides which are difficult to dissolve. Heat treatments above the coarsening temperature may or may not improve the strength properties of such steels, but it almost always results in considerable loss in ductility at rupture. Consideration of these facts lead to the belief that the relationships between properties and the structures developed for several austenitizing temperatures should be established for the subject steels~ The method used to evaluate the influence of austenitizing temperature on microstructure was to determine the coarseness of the bainitic structure resulting from normalizing the 1-inch and 3/4-inch round bar stocks between 1750~ and

4 22000Fe The results of this evaluation have been expressed in terms of a "bainitic" grain size, The bainitic grain size coarsened gradually over the range of temperatures as is shown by Figure 3. It was, therefore, decided to evaluate the high temperature strengths after normalizing over a wide range of temperatures: a, 4340 and "17-22-A"S at 1950~ and 2100~F, supplementing previous data for 1750~F. bo H-40 at 1750' and 2100'F, supplementing previous data for 1950~F. The influence of austenitizing temperature on martensitic structures obtained by oil quenching was partially evaluated for 4340 and "17-22-A"S steels. Due to a shortage of stock, oil quenching of the H-40 was not included, Hardness Level For the work done on the effect of continuous cooling transformation, and all other previous work, all test bars were tempered to a hardness range of 280 to 320 Brinell when the as-transformed hardness was at a sufficiently high level. In so far as possible, the tempering times and temperatures were the same as, or similar to, those previously employed. However, since the hardness of the 4340 steel as-normalized from 1950~ and 2100~F was only slightly above the 320 Brinell maximum and even slight tempering resulted in hardness below the 280 minimum, this steel was tested in the as-normalized condition. Basis of Evaluation of High Temperature Properties The general basis of evaluation of the effects of continuous cooling transformation and of austenitizing temperature was the same as previously employed for the isothermally transformed structures (Ref. 2). Only limited surveytests were selected to indicate general trends of the structural variables rather than accumulating complete design data, The basis for selection was as follows: a, The properties were evaluated for the temperature range over which creep and rupture properties could be the controlling factor. For 4340, this range was set at 7000 to 1100~F, while 700~ to 1200oF was employed for the "17-22-A"S and H-40 steels, The upper temperature was based on the probable maximum temperature at which hardened structures would have useful properties. bo The structures were evaluated on the basis of the property which was the controlling factor at the temperature of interesto Thus, at 7000 and 900~F, the criteria of comparison were mainly creep rate and total deformation characteristics, At thes e temperatures, the stresses required to cause rupture in reasonable times would be well above the yield strength, and thus, service stresses would be limited to those below which rupture would occur,

5 However, approximately 100-hour rupture tests were conducted for 4340 steel at 900~F because the relatively low strength of this material at that temperature indicated that a knowledge of its stress-rupture properties was desirable. c. At 1000~F for 4340 and at 1100~F for "17-22-A"S and H-40, the temperatures of major interest for these steels, stressrupture, creep, and total deformation data from relatively long time rupture tests were obtained. d. The property considered to be of most interest at 1100~F for the 4340 steel and at 1200'F for the "17-22-A"'S and H-40 Steels was the 100-hour rupture strength. However, 1000-hour creep tests were conducted for the 4340 and "17-22-A"S steels to permit better correlation between structural variations and temperature of tes ting. RESULTS The interrelations between steel composition, microstructure and properare fairly complex. Hence, it is necessary to use several test temperatures measures of strength at high temperatures to cover the effects of temperaturetime-amount of creep. Two main variables were studied: (1) Influence of Rate of Continuous Cooling; and (2) Influence of Austenitizing Temperature. In analyzing the results, it seemed best to consider the two together for each steel. Actually, microstructures were the real variable and it is easiest to grasp the data when composition is not a variable. For this reason, all the photomicrographs illustrating the structures are presented together as Figures 4 through 25. The graphical presentation of the creep and rupture data are also kept together as Figures 26 through 43. The results are presented in the following sections as general trends. In most cases, the details should also be considered by those attempting to apply the results. It is also important to recognize that the data are sparse and are only intended for general survey purposes. To aid in the evaluation of the influence of stress from the sparse data, complete curves showing the stress dependency of creep rate, rupture time, and total deformation time for the turbine wheels(Ref. 1) have been included in the appropriate figures. It is also to be noted that specific hardness values are given with the photomicrographs, whereas, average hardness values are used in some of the tables.

Cr-Ni-Mo (SAE 4340) Steel Influence of Rate of Continuous Cooling Reducing the cooling rate changed the as-transformed structures from martensite to bainites plus 35 to 20% martensite (Table I and Figures 9,4,5, and 6). The hardness also decreased. In the normalized 1-inch round the lower hardness was apparently due to the appearance of large amounts of relatively soft bainite. In the 3- and 6-inch rounds, both lesser amounts of martensite and tempering of the martensite present were involved. Other than the degree of tempering, there seemed to be little to distinguish the three normalized structures. To meet the requirement of approximately 300 BHN for high temperature testing, it was necessary to temper the martensite (oil-quenched) structure for 10 hours at 1100'F and the mixed martensitic plus bainitic structure formed after normalizing the 1-inch round for 1 hour at 11000F. The resulting microstructures (Figures 9 and 4) were quite similar. The normalized structures for the larger sizes (Figures 5 and 6) could not be tempered without reducing hardness below the desired range. The structures tested, therefore, represent the spheroidization and agglomeration accompanying tempering during reheating in the first two cases and the structures resulting from direct transformation in the latter two cases. It should be recognized that there is probably a gradation in the degree of tempering of martensite in the structures. The martensite in the mixed bainitic-martensitic structures was probably harder than in the more drastically tempered, originally oil-quenched material. The mixture of hard martensite and softer bainite gave a hardness equal to that of the tempered martensite. A review of the high temperature test data (Table IV and Figures 26 through 30)leads to the following generalities: 1. The tempered martensite had the lowest strength at all temperatures in all the tests used to evaluate the strengths at high temperatures. 2, In most cases, there was very little difference in strength between the martensitic plus bainitic structures produced at the centers of bars having diameters of 1, 3 and 6 inches. There was some slight tendency for the short-time rupture strengths to increase with decreasing cooling rate. 3. Data obtained at 700~F were somewhat erratic, There was possibly a slight tendency at 7000F, and even at some of the higher temperatures, which suggested that the less initial self-tempering of the 3-inch round, in comparison to the normalized 1-inch and 6-inch rounds, was detrimental to strength. 4. The data on ductility in the rupture tests is difficult to generalize because increasing rupture time generally reduces ductility and the data are inadequate to evaluate this effect in relation to initial structure due to the variations in rupture time. It appears from Table IV, however, that the tempered martensite was most ductile at 9000F and probably the larger 3- and 6-inch sizes were least ductile. The tempered, mixed bainites plus martensite of the normalized 1-inch round were least ductile at the higher temperatures, Long duration creep tests at 10000F showed relatively little change in structure and hardness for the tempered martensite (Figure 9) and tempered, mixed bainite and martensite (Figure 4). However, considerable tempering to

7 a coarser carbide and ferrite structure occurred during testing of the larger, normalized sections (Figures 5 and 6) with a large drop in hardness. The relatively sparse data in Figures 26 through 30 are, of course, not conclusive but they do indicate that there is relatively little change in the relative slopes of the log-log relationships between the different structures. Thus, the superiority of the normalized structures is maintained to relatively low stresses and long time periods, even though the absolute differences decrease with decreasing stress. From these results, it can be concluded that the following general relationships between structure and properties at high temperature exist for SAE 4340 steel transformed during continuous cooling over rates ranging from those producing martensite to those producing the mixed, tempered martensite and bainite of air cooling a 6-inch round: 1. At equal, initial hardness levels of 300 BHN, the controlling factor in strength is the presence of considerable bainite (or the presence of bainite indicates some other structural condition controlling strength). Tempered martensite is inferior in strength. 2. The structure seems to be the controlling factor. Large structural changes during testing, as evidenced by large hardness changes of the 3- and 6inch rounds, did not apparently influence the strength criteria when considered in relation to the initially slightly tempered 1-inch round structure which was more stable. (This should not be applied to very long time period strengths without further tests for proof or disproof). 3, Differences or similarities in strength do not seem to be clearly reflected in microstructures before or after testing. The martensitic structure of the oil-quenched material and the martensitic plus bainitic structiures of the normalized 1-inch round were similar as tempered and after testing. They were, however, quite different from the structures of the 3- and 6-inch rounds. Yet, the properties of the normalized 1-inch round were nearly the same as for the 3- and 6inch rounds while the oil-quenched, tempered martensite was consistently weaker. Austenitizing Temperature Increasing the austenitizing temperature before oil quenching increased the coarseness of the martensite and background apparent grain size (Figures 9, 10 and 11). There was also the suggestion that small amounts of bainite were produced. After tempering,there was little difference in structure beyond the size of the background grains (Figures 9, 10 and 11). Increasing the normalizing temperature apparently reduced the amount of martensite and reduced overall hardness (Figures 4, 7 and 8). When normalized from 1950~ and 2100~F, the hardness was too low to allow tempering and still maintain the desired hardness level, Increasing the heat-treating temperature did not produce the apparent coarsening of the grain structure evident in the oilquenched samples. Apparently, this was due to variation in the orientation effects within the grains so that there was no great contrast from grain to grain. The increase in austenitizing temperature before oil quenching reduced rupture strength and ductility somewhat (Table IV and Figure 26). There were

8 smaller, similar reductions in strength and ductility for limited deformations and creep rates (Table IV and Figures 27, 28 and 29). For normalized bars, the increase in temperature of treatment generally raised rupture strengths (Figures 26 and 31). Short-time strengths at 1000~F were an exception. Elongation and reduction of area were also reduced, particularly as the time period for rupture and the testing temperature increased (Table IV). Limited deformations and creep rates generally followed the same trend. The major exception was at 700~F where deformation on loading increased and reduced the time for 1% deformation. The oil-quenched and tempered martensitic structure retained higher hardness and underwent little change in microstructure during testing at 10000F. The untempered, normalized structures softened considerably more than the 1-inch round treated at 1750~F and tempered for 1 hour at 11000F, It seems that even the 1-hour tempering treatment, when it can be applied without too much reduction in hardness, considerably stabilizes the structure. The results indicate that: 1. For the strength criteria considered, increasing the heat-treating temperature slightly reduced strength at elevated temperatures when quenching avas used to develop an originally martensitic structure, The opposite effect was observed for normalizing. In both cases, elongation and reduction of area in the rupture tests were reduced somewhat. 2. The results again indicate that bainites (or treatments which produce bainites) are necessary for highest strength. 3, Any coarsening of the structure resulting from increased heat-treating temperature had opposite effects in that tempered martensitic was weakened while bainitic structures were strengthened. In considering these results, it should be recognized that the lowest temperature of treatment, 1750~F, is the highest temperature recommended on the basis of experience with the alloy, It is possible that the lower temperatures commonly used might have different effects. 1.25Cr-Mo-V (" 17-22-A"S) Steel Influence of Rate of Continuous Cooling The structures obtained ranged from martensite, bainites plus martensite, bainites, to bainites plus a small amount of ferrite (Table II and Figures 17, 12, 13 and 14) as the cooling rate was reduced, The bainites became coarser as the cooling rate was reduced, Hardness values decreased with decreased cooling rate as would be expected from the changes in structure, The severity of the temper had to be reduced with reduced cooling rate in order to maintain the hardness level at about 300 BHN. It should be noted, however

that secondary hardening permitted a 6-hour temper at 1200~F after air cooling a 6-inch round to an as-normalized hardness of 325 BHNo The structures after tempering varied (Figures 17, 12, 13 and 14). The martensite from oil quenching and the bainitic plus martensitic structure of the normalized 1-inch round resulted in a heavy dispersion of carbides with the carbides being coarser in the latter, The bainites of the larger rounds underwent considerable alteration of the ferrite as well as spheroidization of the carbides, The tempered structures all had strengths at high temperatures which showed the tempered martensite from oil quenching to be weakest (Table V and Figures 32 through 37), The strengths of the normalized and tempered structures tended to increase as the section size increased, although this effect definitely decreased with increasing test temperature. The evaluations on the basis of time and creep rate at a specific stress at 7000 and 900~F exaggerate the effect since the probable stress-time for deformation and stress-creep rate curves have very little slope at these temperatures. In other wor'ds, wide differences in these values at a specific stress represent very little effect on the stress for a given deformation or creep rate. Elongation and reduction of area values in the rupture tests were low at 11000 and 1200~F for all initially normalized structures and apparently tended to be slightly lower for the slower initial cooling rates. Even the tempered martensitic structure had quite low ductility for the longer times at 1100'F, Some softening and spheroidization occurred during creep testing at 1100~Fo However, there was very little difference in this respect between any of the four initial structures. The data are essentially similar to those for SAE 4340 steel, Originally martensitic structures are weaker than originally bainitic structures at 7000 to 1200~F. The strengths generally increase slightly with decreased cooling rate so long as essentially bainitic structures form initiall.y. This relationship existed after drastic tempering. Influence of Austenitizing Temperature Oil quenching from 1950~F produced a somewhat coarser martensitic structure than from 1750~F (Table II, Figures 17 and 18), When the heat-treating temperature was raised to 2100~F, however, the structure contained about 50% of what appeared to be bainite in very coarse grains of martensite (Figure 19). The as-transformed hardness decreased with increasing temperature of heat treatmento After tempering 1 hour at 1300~F, the degree of carbide spheroidization decreased with increasing hardening temperature (Figures 17, 18 and 19), The hardness differences were, however, small in comparison to the initial hardness values, The strengths at high temperature (Table V and Figures 32 through 35) were generally higher for those samples originally quenched from the higher temperatures. The 1950~F treatment generally had the highest strength, except for the lower stress tests at 1200~F. The 1950~F treatment produced structures which were probably inferior in strength at 700uF to those produced by the 17500F treatment, Increasing the austenitizing temperature before oil

10 quenching reduced ductility in the rupture tests (Table V and Figure 32), particularly at the shorter time periods. It appears that increasing the austenitizing temperature did increase the strength of originally martensitic structures somewhat. The appearance of considerable, fine bainite-appearing structure in the material quenched from the highest temperature, however, did not result in a substantial increase in strength. Increasing the normalizing temperature to 19500 and 2100~F prevented the formation of any martensite and gave increasingly coarse 100% bainitic structures (Table, II and Figures 12, 15 and 16). The hardness decreased accordingly. Less tempering was necessary for the structures formed from the higher temperatures to obtain 300 BHN. The spheroidized carbides were less evident and the background ferritic structure more prominent after tempering the structures formed from the higher normalizing temperatures. The strength properties increased with normalizing temperature at 700~, 900~, 1100~ and 1200~F, except for short-time rupture strength at the latter two temperatures (Table V, Figures 32, 33, 34, 35 and 37). The changes were not particularly large in any case. The apparent, large changes at 700~ and 900'F would not be nearly so prominent on the basis of stress for a given deformation or creep rate. Ductility dropped off in the rupture tests as a result of increasing the normalizing temperature (Table V). The structures were somewhat less stable with increasing normalizing temperatures as judged by hardness after testing at 11000F (Figures 12, 15 and 16). Some spheroidization occurred during testing, possibly being least in the material normalized from 1950~F. Increasingly coarse bainite resulting from increasing the normalizing temperature was accompanied by moderately increased strengths at high temperatures. The rather low ductility in rupture tests accompanying a normalize from 1750~F was lowered still further. The strengths were substantially higher than those obtained by oil quenching from the same temperatures. Raising the austenitizing temperature, in some cases, produced strengths in oil-quenched, martensitic structures slightly higher than those obtained by normalizing at 17500F. It thus appears that some effects accompanying increased austenitizing temperature are effective in increasing strength regardless of whether martensite or bainites are formed on cooling. A bainitic structure will, however, generally be superior to a martensitic structure provided both are formed from the same austenitizing temperature.

11 3Cr-Mo-W-V (H-40) Steel Influence of Rate of Continuous Cooling Reducing the cooling rate altered microstructure mostly between oil quenching and normalizing of 3/4-inch rounds by changing from martensite to a predominately bainitic structure (Table III and Figures 25, 20, 21 and 22). Completely bainitic structures were formed in the 3- and 6- inch sections. The bainitic structures in the normalized samples were not greatly different, possibly becoming somewhat coarser with decreasing cooling rate. There was a substantial increase in the background apparent grain size with decreasing cooling rate. There was a small decrease in as-transformed hardness of the bainitic structures with decreased cooling rate. The as-transformed structures varied considerably in their resistance to tempering (Table III and Figures 25, 20, 21 and 22), The normalized 3/4-inch round was the most resistant, while the 3- and 6-inch sections were least resistant, The structures after tempering were not greatly different as viewed at high magnification. The difference in background grain size was quite prominent at low magnification. Changing from an originally martensitic to predominately bainitic structure by changing from oil quenching to normalizing 3/4-inch rounds had little benefit in the tests at 700~ and 900~F (Table VI and Figure 42). The 3- and 6-inch rounds were, however,. substantially stronger at these two temperatures. At 1100~F, improvement was obtained by normalizing. The structures giving improvement, however, varied between the 3/4- and 3-inch normalized rounds (Table VI and Figures 38 through 42). Apparently, at 1200~F, there would be a steady improvement with decreasing cooling rate. Ductility in the rupture tests decreased as the result of normalizing to form bainitic structures (Table VI). The 3-inch section samples were least ductile with some evidence that the 6-inch round was not reduced as much in ductility. The alteration of structure during creep testing at 11000F was relatively small microscopically (Figures 25, 20, 21 and 22). The 3- and 6-inch sections, however, underwent considerably more softening than the two smaller section sizes. Influence of Increased Austenitizing Temperatures Reducing the normalizing temperature to 1750'F produced a 100% fine bainitic structure. The two higher temperatures resulted in a small amount of martensite plus fine acicular bainite with somewhat increased hardness (Table III and Figures 24, 13 and 25). The resistance to tempering increased with normalizing temperature (Table III and Figures 23, 20 and 24). Tempering produced spheroidized carbides and alteration of the ferrite. The strengths at high temperatures of the tempered structures generally increased with normalizing temperature to a pronounced extent (Table VI and Figures 38, 39, 40, 41 and 43). The only exception might be 700~F although additional tests should be made to verify the apparent strength of the material normalized at 1750~F. In most cases, the increases in strength were large in comparison to the other two steels considered.

12 Ductility in the rupture tests decreased markedly when the austenitizing temperature was raised from 1750~ to 19500F (Table VI). A further decrease resulted from raising the temperature to 2100~F. The structure obtained by normalizing from 1950'F was far more stable than those of either the 17500 or 2100~F treatments (Figures 23, 20 and 24) as judged by hardness drops during creep testing at 11000F. The material treated at 1750~F may have undergone recrystallization during testing. The low strength after a 1750'F normalize indicates that a bainitic structure alone is not sufficient to obtain high strength at high temperatures. There was also more effect of heat-treating temperature in this more highly alloyed steel. DISCUSSION From a practical heat treating viewpoint, the data obtained lead to the following important generalizations: 1. Increasing the section size (up to 6 inches) of bars normalized from the established austenitizing temperatures did not substantially weaken any of the steels. In most cases, it actually increased strengths even when the structures could not be tempered without reducing hardness below 300 BHN. The larger section sizes often softened more during testing, but this evidence of structural instability did not seem to adversely affect the strengths considered. 2. Increasing the austenitizing temperature above the established he.ttreating temperatures did not give marked improvement in strength for either SAE 4340 or "17-22-A"S. The influence was greater in the case of the H-40 steel which contained vanadium, suggesting that more complete solution and diffusion of vanadium compounds was somewhat beneficial. In the 1. 25Cr-0. 5Mo0. 25V (t117-22-A"S)steel both martensitic and bainitic structures were somewhat increased in strength by increasing the austenitizing temperature. This would probably have been true also for the 3Cr-Mo-W-V (H-40) steel although only the bainitic structures were tested. Martensitic structures in SAE 4340 steel were not improved but the bainitic structures obtained by normalizing were. 3. Too low a temperature of austenitizing for complete solution is very detrimental as judged by the results for H-40 steel. 4. Increased austenitizing temperature or decreased cooling rate generally reduced ductility in the rupture tests. Correlation of the Structures and Properties of the Continuously Cooled Bars and Turbine Wheels The figures showing the stress dependency of strength properties included curves for forged turbine wheels previously established (Ref. 2). One reason for this was to indicate probable stress dependency effects for the sparse survey data from this investigation. In addition, it permits, in conjunction with the micro

13 structures, an appraisal of the agreement in properties for the continuously cooled samples of this investigation with those which existed in the wheels examined for Reference 2. Cr-Ni-Mo (SAE 4340) Wheels and Continuously Cooled Bar Stock The oil-quenched wheel was treated at 15500F and tempered at 10500F to 280/320 BHN. This produced a uniformly tempered, martensitic structure which appeared to be similar to the bar stock oil quenched from 1750~F for this investigation. References to Figures 26 through 29 show that the two had approximately the same properties, except that the bar stock had lower strength at short time periods, and higher strength at longer time periods. Apparently, the bar stock material did not exhibit the breaks in the curves that the wheel did. There are a number of possible reasons for this, One is that the wheel was actually cooled at a somewhat slower rate than the bar. It will be noted that the slower cooled, normalized disk did not show the breaks.. It is, however, equally possible, in view of the lack of confirming data, that the lower austenitizing temperature of the wheel was responsible. The normalized wheel was treated at 1750~F and was not tempered. The section size of the disk ranged from about 3-inches to 6-inches. The structures of the disk did not perfectly match those of the samples simulating the centers of 3- and 6-inch rounds. The restriction of cooling rate to more closely simulate that of the wheel did not improve the agreement in strengths over that obtained with the 1-inch normalized bar stock (Figures 26 through 29). It appears that either differences between the heats of the wheel and bar stock materials or the double normalize of the wheel resulted in the different response to heat treatme nt. 1. 25Cr-Mo-V ("17-22-A"S) Wheels and Continuously Cooled Bar Stock The oil-quenched wheel had properties which were in general superior to the normalized wheel, In Reference 2, it was concluded that the retarded cooling rate from the mass of the forging gave lower bainite plus martensite on oil quenching. The normalized wheel had varying tempered bainitic-ferritic structures near the rim, and bainitic-pearlitic structures near the hub. Again, the mass of the forging was presumed to have retarded cooling rate and allowed higher temperature transformation products to form. Unfortunately, as-transformed structures for the wheel were not available, However, the structures resulting from the simulated rates of continuous cooling were compared with the tempered wheel structures. The simulated center of the 6-inch round gave a structure resembling the rim of the normalized wheel. In no case was the predominantly peailitic type structure of the wheel found. The structure of the oilquenched wheel was not reproduced in the bar stock experiments, The.strengths of all three normalized section sizes were superior to the oil-quenched wheel in most cases, as is shown by Figures 32 through 35. The oil-quenched bar stock was generally inferior, It is therefore presumed that the oil-quenched wheel had structures in between the martensite formed by oil quenching and the bainitic-martensitic structure formed by normalizing the 1

14 inch diameter bar stock. The large amount of pearlite in the normalized wheel has always been a mystery because the alloy should not have formed pearlite on air cooling a 6inch section. Secondly, the oil-quenched wheel should have been largely martensite. Differences between the two heats could be involved. Prior history effects also could have been a factor. The most probable explanation, however, is in unrecorded variations in the thermal conditions during heating and cooling of the wheels. 3Cr-Mo-W-V (H-40) Wheels and Continuously Cooled Bar Stock In Reference 2, it was observed that there was little difference in structure between the oil-quenched and normalized wheels of H-40 steel after tempering. The properties were also similar. In so far as could be determined, the wheels apparently had a coarse-grained, tempered bainitic structure, Good agreement in rupture strength was found between the bainitic bar stock structures and the wheels in spite of the apparent difference in grain size. Total deformations, however, were less, Normalized and tempered 3/4-inch bar stock, even though it did not have a similar structure, came closer to the wheels for limited deformations. The structures of the continuously cooled bar stock from this investigation were compared with those of the wheels even though it was difficult to draw conclusions from the tempered structures of the wheels, The structures of the 3and 6-inch rounds were quite close to those of the wheels. The 3-inch round normalized specimens matched the properties of the oil-quenched wheel somewhat closer than the normalized 3/4-inch round. It seems somewhat surprising that the center of a normalized 3-inch round should match the 3-inch section of an oil-quenched wheel. The wheel apparently had a coarse grain size. The results of the present investigation indicated that the retarded cooling rate developed the appearance of.the coarse grains. Since the wheels and bars were from the same heat, this points to the wheels, even when oil quenched, having been cooled at rates similar to those of the simulated 3- and 6-inch rounds. General Comments on Correlation of Structures and Properties of Wheels and Continuously Cooled Bars One of the reasons the correlation between bar stock and wheels of the 4340 and "117-22-Al'S steels was not better could be that different heats were involved, There was also a difference in austenitizing temperature for 4340 ste el The questionable aspects of the structures of the "17-22-A"'S and H-40 wheels previously mentioned leaves questions regarding the conditions of treatment of the wheels, It must be recognized, however, that there were differences in heating rates, and therefore, times at temperature, between the bar stock and the wheels

15 as well as differences in prior history. The results of the tests involving variation in austenitizing temperature showed some sensitivity to this factor for both the "17-22-A"S and H-40 steels. It is, therefore, entirely possible that the heating condition during forging combined with the final heat-treating conditions were sufficiently different to cause the observed differences in structure and prope rtie s. Structures of Continuously Cooled Bars In an attempt to better define the structures of the continuously cooled specimens, they were compared with those obtained by isothermal transformation in Reference 2. Cr-Ni-Mo (SAE 4340) Steel As far as could be judged, the three normalizing conditions produced mainly upper bainite and martensite, with the greatest effect being the increasing amount of tempering of the martensite as the cooling rate decreased. There was some gradation of the fine mass of the bainite, with the 1-inch stock having the greatest predominance of fine bainite. This follows the general trend of properties as set forth in Figure 30 and the shift in relative properties towards superior properties for upper bainite with increasing test temperatures (Ref. 2). Increasing the normalizing temperature apparently altered transformation conditions so that the bainite which formed more nearly approached the softer middle bainatic structures, This also is consistent with the correlation of properties of Ref. 2 where increasing amounts of middle bainite gave superior properties in the 9000 to 1000~F range., e25Cr-Mo-V ("17-22-A"S) Steel As far as could be judged, normalizing the I-inch section gaved mixed middle bainite and martensite. Increasing the section size resulted in a shift towards upper bainiter Apparently, increasing the normalizing temperature had the same effect. This is not too generally consistent with the properties versus structure relationship of Reference 2 where the middle bainitic structure generally had the highest strengths. It would appear that a more thorough knowledge of the factors involved in continuous cooling transformation is needed. 3Cr-Mo-W-V (H-40) Steel Only one temperature of isothermal transformation to bainite was reported in Reference 2 due to the narrow temperature range in which bainite would form in a reasonable time, The three slower cooling rates of this investigation all gave similar structures, except for the apparent grain size of the bainite, and the properties were not too far different, The major difference was decreased resistance to tempering with attendant decreased retention of hardness during testing as the cooling rate was decreased, The "bainite" formed isothermally for Reference 2 did not closely resemble the structures formed in this investigation. It is concluded that there are differences between structures formed iso

16 thermally and structures formed over a range of temperature on continuous cooling which must be studied in greater detail if a completely satisfying correlation is to be made between the present work and the work reported in Reference 2. CONCLUSIONS 1i, All of the data showed predominantly bainitic structures to have the highest strength in the range from 7000 to 12000F. 2. Sufficiently rapid cooling rates to produce essentially martensitic structures result in lower strength. When the cooling rate was reduced from oilquenching to normalizing of 1-inch rounds, changing the structure from martensitic to predominantly bainitic gave a substantial increase in strength with loss in rupture-test ductility. Further reductions in cooling rate to those equivalent to the centers of air-cooled round bars 3 and 6 inches in diameter still resulted in batinitic structures-9 with a further increase in strength and some. loss.- in ductility. 3. The increase in strength from reducing the cooling rate from that of the centers of air-cooled 1-inch rounds to that of the centers of 6-inch rounds appeared to be due to the development of complete bainitic structures similar to middle to upper bainite, The strengths approached those of the strongest structures developed by isothermal treatment. The correlation between observed structures of this phase of the investigation and previous studies for isothermally transformed structures was fairly good for 4340 steel, not as good for "17-22-A"S and poor for H-40 steel. The assumption is that there is need for more detailed information concerning the relationships between structures formed on continuous cooling and structures formed isothermally, as well as the relationship between these two types of structures and high-temperature properties. The estimation of properties of structures formed by continuous cooling is still uncertain for this reason. 4. Increasing heat-treating temperature for 1-inch rounds generally increased strength and reduced ductility. The largest effect was obtained from H-40 steel, This was true for martensitic as well as bainitic structures, except for 4340 steel where the martensitic structures were not increased in strength. 5. The range of austenitizing temperature (1750~ to 2100'F) was at or above the usual temperature of treatment of 1750~F for 4340 and "17-22-A"S while 1750'F was below the established temperature of 1950F for H1 —40. This probably accounts for the much greater apparent effect for H-40 steel. 6. Increasing the austenitizing temperature generally increased the coarseness of the background grain size. The effect on the basic structure was less, although there was a tendency on normalizing for increasing coarseness of the bainite with increased austenitizing temperature,

17 7. The correlation of the structures of the continuously cooled structures and properties with those of forged wheels previously investigated ranged from poor for 4340 and "17-22-A"S to fair for H-40 steel. The lack of better correlation was assumed to be due to the materials being from different heats and differences in prior thermal history. 8. The trends shown by the data are believed to be quite reliable. However, it should be recognized that survey-type tests were used - necessarily limiting the dependability of the exactness of the levels of the strength criteria. BIBLIOGRAPHY 1. A. Zonder, A. I. Rush, and J. W. Freeman, "High-Temperature Properties of Four Low-Alloy Steels for Jet-Engine Turbine Wheels", Wright Air Development Center Technical Report 53-277, Part I (November, 1953) 2. A. I. Rush and J. W. Freeman, "High-Temperature Properties of Four Low-Alloy Steels for Jet-Engine Turbine Wheels", Wright Air Development Center Technical Report 53-277, Part II (February, 1955) 3. T. F. Russell, "Some Mathematical Considerations on the Heating and Cooling of Steel", First Report of the Alloy Steels Research Committee by a Joint Committee of the Iron and Steel Institute, 1936, p. 149.

TABLE I Influence of Austenitizing Temperature, SectionSize, and Cooling Medium on the Microstructure and Hardness of SAE 4340 Steel Austenitizing Quenching Bar Diameter Microstructure Obtained Average Tempering Conditions Average Temp. ( 0F) Medium (inches) BHN Temp (0F) Time (hrs) BHN 1750 Oil (a) 1 100%6 martensite 585 1100 10 307 Air (a) 1 35% martensite + 65%o bainites 374 1100 1 306 Air 3 (b) 25% martensite + 75% bainites 329 None -- 329 Air 6 (b) 20%1o martensite + 80% bainites 322 None -- 322 1950 Oil 1 100% martensite 560 1100 10 302 Air 1 15% martensite + 85% bainites 328 None — 328 2100 Oil 1 100% martensite 534 1100 10 298 Air 1 15% martensite + 85% bainites 333 None -- 333 (a) Previously reported in WADC TR 53-277 Part II (b) 1-inch diameter bars were air cooled in insulating cylinders of firebrick to simulate the cooling cycles of normalized 3- and 6-inch diameter bar stock.

TABLE II Influence of Austenitizing Temperature, Section Size, and Cooling Medium on the Microstructure and Hardness of "17-22-A'S Steel Austeniti z ing Quenching Bar Diameter Microstructure Obtained Average Tempering Conditions Average Temp ('F) Medium (inc he s) BHN T emp- ('F) Time (hrs) BHN 1750 Oil (a) 1 10% martensite 525 1300 1 291 Air (a) 1 15% martensite + 85% coarse 355 1200 10 307 bainites Air 3 (b) 100% bainite 325 1200 6 302 Air 6 (b) 98% bainite + 2% ferrite 325 1200 6 306 1950 Oil 1 100% martensite 470 1300 1 300 Air 1 100% bainite 335 1200 6 298 2100 Oil 1 50% martensite + 50% lower 412 1300 1 290 bainite Air 1 100% bainite 313 1200 6 306 (a) Previously reported in WADC TR 53-277 Part II (b) 1-inch diameter bars were air cooled in insulating cylinders of firebrick to simulate the cooling cycles of normalized 3- and 6-inch diameter bar stock.

TABLE III Influence of Austenitizing Temperature, Section Size, and Cooling Medium on the Microstructure and Hardness of H-40 Steel Austenitizing Quenching Bar Diameter Microstructure Obtained Average Tempering Conditions Average Temp ('F) Medium (inches) BHN Temp ('F) Time (hrs) BHN 1750 Air 3/4 100% fine bainites 409 1200 4 310 1950 Oil (a) 3/4 100% martensite 523 1200 12 306 Air (a) 3/4 20% martensite + 80% bainites 435 1200 18 316 Air 3 (b) 100% bainites 400 1200 4 308 Air 6 (b) 100% bainites 390 1200 4 309 2100 Air 3/4 20% martensite + 80% bainites 429 1250 4 + 2 319 (a) Previously reported in WADC TR 53-277 Part II (b) 1-inch diameter bars were air cooled in insulating cylinders of firebrick to simulate cooling cycles of normalized 3- and 6-inch diameter bars.

TABLE IV Deformation Time to Reach Specified Total Austenitizing Quenching Bar Diam Stress Rupture Time Elongation Reduction of Area on Loading Deformation (hours) Minimum Greed Rate Temp. ('F) Medium (inches) BHN (psi) (hours) (% in 2 in) (%) (%) 0~-.1% 0.2% 0.5% 1.0% (%/hour) 700~F 1750 Oil(b) I 304 90,000 1350(d) 0.430 a a g 675 0. 00027 Air(b) I 300 90,000 129 t(d) 0. 467 a a I 1000 0. 00016 Air 3(c) 327 90,000 134g(d) 0.463 a a ~3 668 0. 000gl Air 6(c) 315 90,000 1342(d) 0.450 a a ~a5 1000 0. 000ZI 1950 Air I 328 90,000 1293(d) 0.54g a a a 440 0.00014 2100 0il I 294 90,000 1584(d) 0.511: a 215 0.00032 Air I 336: 90,000 1293(d) 0.547 a 565 0.00015 900'F 1750 Oil(b) I 306 55,000 381 19.5 39.5 0.269: 2 13 0.01480 Air(b) I 300: 55,000 842 12.0 22.3 0.260 8 64 0. 00414 Air 3(c) 3Z7 55,000 1536 6.0 6.3 0. 305 a a 3 30 0. 00130 Air 6(c) 315 55,000 1951 5.0 5.5 0.290 a a 3 45 0.00085 1950 Air I 336 55,000 1660 4.0 4.7 0.307 a a 5 263 0.00056 2100 Oil I 296 55,000 324 10.0 12.? 0. 300 a ~3 12 0. 01000 Air I 328: 55,000 1975 2.0 4.? 0. 285 a 5 277 0. 00065 1750 Oil(b) I 302 40,000 3338 4.0 5.5 0. 175 a 50 355 0. 00111 Air(b) I 300 40,000 19 l')(d).... 0. 164 a; 1160 > 3000(e) 0.00015 Air 3(c) 331 40,000 1404(d).... 0. 185 Ha 82 1800(e) 0. 00024 Air: 6(c) 315 40,000 1483(d).... 0. 175 ~4 430 > 3000(e) 0.00015 1950 Air I 336 40,000 l-t61(d) o- 0. 163 a 2.5 475 (f) 0.00013 2100 Air I 328 40,000 1650(d) -- 0. 194 a a 425 (f) 0.00012 1000'F 1750 Oil(b) I 310 31,000(g) 160 11.0 15.0 0. 149 ~a I,~3.5 16 0.02500: Air(b) I 290 31,000 371 ~. 5 7.-t 0. 126 N5 50 145 0.00505 Air 3(c) 331 31,000 259 1,'.5 0.148 a NI 16 39 0.01140 Air 15.71 6(c) 315 31,000 362 11.0 11 0. 157 a NI 19 94 0.00650!950 Oil 1 302 31,000 111 9.0 0. 173 a <1 3 18 0. 02800 Air q.% I 321 31 000 283 o.0 4. 0. 148 a N2 35 142 0.00420 Oil I 296 31, 000(~) 130 10.0 I 1. t 0. 158 a < I 3 18 -- Air I 328 31,000 490 -i. 5 4.0 0. 148 a Ng 50 220 0.'00280 t?50 Oil(b) I 309 20,000(g) 780 12.0 15.0 0.099 a? 47 190 0.00380 Air(b) I 300 20,000 1392 5.0 4 0 0. 090.v 1 20 228 650 0. 00114 Air 3(c) 327 20,000 1310 8.5 o.0 0. 101 a 5 110 38? 0. 00170 Air 6(c) 315 20 000 1488 o.0 6.0 0.092 NI 12 145 534 0. 00120 1950 Oil I 302 20,000(g) 480 7. ~ 9.5 0. 110 a 3 33 130 0.00520 Air I 325 20,000 1425 Z. 5 3.0 0. 093 < I ~ 2 120 768 0. 00073 2100 Oil 1 301 20,000 548 8.5 9.8 0.091 <1 10 58 166 O. 00440 Air I 328 20 000 2257(d).... 0. 116 a 18 285 1300 0.00043! 100~F 1750 Oil(b) 1 293 18,000 43. 5 Z0.0 25.2 0. 148 a........ Air(b) I 293 18,000 69.6 7.0 11.7 0.116 a Air 3(c) 328 18,000 78.6 17.5 32.5 0. 107 a <1 2.5..... Air 6(c) 321 18,000 77.2 18.0 21.8 0.112 a <1 3..... 1950 Oil I 302 18,000 36. 3 12.0 14.8 0. 106 a........ Air I 321 18,000 105.0 5.5 7.0 0.093 <1 Z 12 45 --- 2100 Oil I 284 18,000 67.2 19.5 18.9 0.215 a a....... Air 1 328 18,000 123.0 5.5 4.9 0.105 a 2.5 14 40 --- 1750 Oil(b) I 309 4,500 1080(d) 0. 027 5 22 104 258 0. 00316 Air(b) I 311 4,500(g) 1100(d) 0.020 12 56 300 900 0.00150 Air 3(c) 327 4,500 815(d) 0. 022 15 60 268 725 0. 00108 Air 6(c) 315 4,500 839(d) 0.030 15 60 305 822 0. 00100 1950 Air I 325 4,500 2740(d) -- 0.040 17 165 1148 (f) 0.00019 2100 Air I 330 4,500 1897(d) -- 0.035 50 250 1810 (f) 0.00015 {a) Specimen reached indicated deformationon loading (b) Data previously reported in WADC TR 53-277 Part II (c) l-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycles of normalized 3- and 6-inch diameter bars. (d) Test discontinued at indicated time (e) Extrapolated value (f) Deformation not obtained during testing period and to indicate time would have required excessive extrapolation (g) Interpolated values

TABLE V Time to Reach Specified Total Rupture Reduction of Deformation on Deformation Minimum Austenitizing Quenching Bar Diam. Stress Time Elongation Area Loading (hours~ Creep Rate Temp. ('FI Medium (inches) BHN ( s p_~._ (hours) (% in 2 in) (%) 0. 1% 0.2% 0~ 5% ~ l'.(J% (%/hr) (%1 700'F 1750 Oil(b) I 298 107 000 1145(d).... 0. 550 a a a 1000 0. 00010 Air(b) I 307 102,000 1194(d).... 0. 465 a a ~'1 >2000(e) 0. 00007 Air 3(c) 291 102,000 1793(d).... 0. 620 a a a 150 0. 00016 Air 6(c) 294 102,000 1324(d).... 0. 660 a a a f 0. 00001 1950 Oil I 294 102,000 1733(d) -- 0. 715 a a 75 0. 00008 Air: I 302 102,000 1757(d) -- 0. 550 a a f 0. 00003 2100 Air 1 102,000 1897(d) -- 0.653 a a a f 0.000016 900'F 1750 Oil(b) I 272 70,000 756 30.0 64.0 0. 378 a a 3 50 0. 00384 Air(b) 1 303 70,000 1482(d).... 0. 335 a a 24 1400 0. 00030 Air 3(c) 302 70,000 1223(d).... 0. 288 a 235 >2000 0. 00014: Air 6(c) 297 70,000 1152(d).... 0. 285 a 525 >2000 0. 00008 1950 Oil I 300 70,000 714 6.0 17.5 0. 370 a a 22 215 0. 00180 Air I 300 70,000 1277(d).... 0. 355 a a 60 f 0. 00008 Zi00 Air I 70,000 2020(d).... 0. 350 a a 263 f 0.00004 iI00'F i750 Oil(b) I 293 41,000 23.4 28.0 27.5 0. 173 a - 0. 0065 Air(b) I 309 41,000 112 2.5 3. I 0.212 a 26 0.00614 Air 3(c) 291 41,000 110 2.0 h 0.206 a: 34 0.00523 Air 6(c) 293 41,000 115 2.0 3.~ 0. ZZl a a 29 0.00420 i950 Oil I 294 41,000 29 5.5 10.0 0.213 a a - - Air I 298 41,000 75 2.5 1.2 0.172 a <1 16.5 - - a100 Oil I 290 41,000 29 8.0 14.5 0.215 m a - - - Air I 317 41,000 79 I. 0 1.0 0. 205 a a 54 - -!750 Oil(b) I 306 19,000(g) 850 4.0 -- 0. 105 17 170 420 0. 00152 Air(b); I 311 19,000(g) 900 2.0 -- 0. 085 80 580 800 0. 00063 Air 3(c) 295 19,000 1220 2.0 h 0. 090 15 193 1045 1500(e) 0. 00026 Air 6(c) 291 19,000 1309 1.0 /.0 0.091 5 183 994 1400(e) 0.00028 ~950 Oil I 298 19,000 1081 2.5 Z. 0 0. 090 ~Z 62 597 ~'1070 0. 00047 Air I 294 19,000 1205 1.5 1.0 0. 080 10 150 940 - 0. 00028 Z100 Oil 1 290 19,000 1001 3.0 3.9 0. 105 40 328 888 0. 00068 Air 1 317 19,000 1653 0.5 h 0. 090 425 ~1650(i) ~ (i) 0. 00012 IZ00'F 1750 Oil(b) I 298 14,000 73 8.0 14.9 0. 096 a 4 14 - Air(b) I 304 14,000 167 -1.0 5.0 0. 066 5 22 65,~, 140 0. 0064 Air 3(c) 290 14,000 158 2.0,'. 8 0. 072 5 22 9I,"130 0. 0040 Air 6(c) 295 14,000 15Z 1.5 h 0.070 5 22 93 aPlZ7 0.0041 1950 Oil I 294 14,000 157 4.5 8. I 0. 080 ~ 11 42 92 0.0084 Air I 302 14,000 1,'-9 3.5 3.5 0. 073 2 16 87 - 0. 0052 2100 Oil I 302 l'i, 000 129 6.0 9. 5 0.089,~,1 7 24 57 0.0150 Air I 294 1.t, 000 214 1.5 3.9 0.080,~,1 61 213 - 0.00155 1750 Oil(b) I 310 7,500 575 30.0 39.8 0.058 6 17 69 144 0.0066 Air(b) I 313 7,500 918 10.0 14.9 0.046 6 46 176 333 0. 0023 Air 3(c) 313 7,500 1079(d).... 0. 029 45 160 532 945 0. 00085 Air 6(c) 322 7,500 1125(d).... 0. 037 37 155 568 1030 0. 00068 1950 Oil I 286 7,500 937 5.0 14.4 0. 052 8 31 153 347 0. 00218 Air I 294 7,500 1757 2. b 7.5 0. 037 37 196 717 1320 0. 00056 2100 Oil I 286 7,500 1073 6.0 1.9 0.045 45 103 317 62O 0.00134 Air I 294 7,500 Z313 4.0 7.5 0. 035 82 360 1500 A, 2300 0. 00022 (a) Specimen reached indicated deformationon loading (b) Data previously reported in WADC TR 53-277 Part II (c) l-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycles of normalized 3- and 6-inch diameter bars. (d) Test discontinued at given time (e) Extrapolated value (f) Deformation not obtained during testing period, and to indicate time would have required excessive extrapolation (g; Interpolated values (h) Fractured in shoulder radius (i) Failed with only 0.5% elongation

TABLE VI Rupture, Total Deformation, and Creep Data at 700', 900', 1100., and 1200'F for H-40 Steel for Several Austenitizing Temperatures and Cooling Cycles Time to Reach Specified Total Rupture Reduction of Deformation on Deformation Minimum Austenitizing Quenching Bar Diam. Stress Time Elongation Area Loading (hours) Creep Rate Temp. ('F) Medium (inches) BHN (psi) (hours) (% in 2 in ) (%) (%) I% 0.zo. 5o 1-. -0.o (%/hour) 700'F 1750 Air 3/4 311 90,000 1369(d) -- -- 0.375 a a 100 f 0.00006 1950 Oil(b) 3/4 290 90,000 1514(d) -- -- 0.410 a a 15 "2000(e) 0.00019 Air(b) 3/4 310 90,000 1292(d) -- -- 0.416 a a 13 1770(e) 0. 00017 Air 3(c) 313 90,000 1272(d) -- -0. 390 a a 135 f 0. 00008 Air 6(c) 322 90, 000 1124(d) -- -- 0.410 a a 60 f 0.00008 2100 Air 3/4 319 90,000 1897(d) -- -- 0.458 a a 20 f 0.00005 900'F 1750 Air 3/4 323 65,000 699 24.0 70.0 0.355 a a 4 77 0. 0556 1950 Oil(b) 3/4 290 65,000 917 31.0 68.0 0.359 a a 1 50 0.00463 Air(b) 3/4 320 65,000 1052 18.0 36.0 0.301 a a 10 85 0.00328 Air 3(c) 313 65,000 1131(d) -- -- 0.390 a a 20 1500(e) 0.00011 Air 6(c) 322 65,000 1131(d) -- -- 0.414 a a 15 62200(e) 0.00008 2100 Air 3/4 319 65,000 1897(d) --. - 0. 283 a a 625 f 0. 00005 1 100'F 1750 Air 3/4 321 40,000 34.7 39.0 77.0 0.253 a a -- -- -- 1950 Oil(b) 3/4 320 40,000 136 12. 5 46. 0 0.232 a a 10 39 0.0160 Air(b) 3/4 310 40,000(g) 110 5.0 13.6 0.231 a a 30 89 0.0054 Air 3(c) 304 40,000 213 4.0 9.5 0.230 a a 15 107 0.0047 Air 6(c) 297 40,000 149 10.5 31.0 0.230 a a 9 44 0.0134 2100 Air 3/4 319 40,000 427 5.0 5. 5 0.238 a a 125 #"400 0.0014 1750 Air 3/4 301 30,000 83.7 54.0 79.5 0.187 a l 21 0.0336 1950 Oil(b) 3/4 321 30,000 865 20.0 26.0 0.150 a 2 94 279 0.00237 Air(b) 3/4 320 30,000(g) 850 h -- -- a 10 290 720 0. 00090 Air 3(c) 313 30,000 1317(d) -- -- 0.201 a a 160 725 0.00079 Air 6(c) 322 30,000 1318(d) -- -0.205 a a 215 910 0.00068 2100 Air 3/4 319 30,000 1922(d) -- -- 0.215 a a 317 1600 0.00037 1200'F 1750 Air 3/4 298 25,000 13.8 56.0 84.0 0.276 a a -- _- 1950 Oil(b) 3/4 300 25, 000 62 45.0 74.0 0. 185 a <1 4 14 -- Air(b) 3/4 315 25,000 100 17.0 45.0 0. 142 a 1 9 38 -- Air 3(c) 304 25,000 188 10.5 11.7 0. 170 a 1 24 71 -- Air 6(c) 297 25,000 205 11.0 9.4 0. 172 a 1 29 72 -- 2100 Air 3/4 319 25,000 244 7.0 7.5 0. 197 a ov6 31 119 0.0064 (a) Specimen reached indicated deformation on loading (b) Data previously reported in WADC TR 53-277 Part II (c) 1-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycles of normalized 3- and 6-inch diameter bars (d) Test discontinued at indicated time (e) Extrapolated value (f) Deformation not obtained during testing period and to indicate time would have required excessive extrapolation (g) Interpolated values (h) Broken in shoulder radius

TABLE VII Rupture, Total Deformation, and Creep Strengths at 10000 and 1100~F for SAE 4340 Steel as Influenced by Heat Treatment One-Percent Total Bar Rupture Strength Deformation Strength 0. 001%o/Hour Diam. (psi) (psi) Creep Strength Heat Treatment (inches) 00r 100-h r 00-hr 100hr (psi) 1000~F 1750~F - Oil Q, 1 35,000 18,500 22,500 11,000 12,000 1750~F - Norm. 1 46,000 22,000 33,000 17,000 19,000 1750~F - Norm. 3(a) 40,000 21,500 26,500 15,500 17,500 1750~F - Norm. 6(a) 46,000 23,000 30,000 16,000 19,000 1950~F - Oil Qt 1 32,000 16,000 21,500 ( 8,000) (11,000) 1950~F - Norm. 1 42,000 22,000 33,000 18,500 21,500 2100~F - Oil Q. 1 33,500 17,000 22,500 (11,000) (11,000) 2100~F - Norm. 1 45,000 26,000 36,000 22,000 24,000 1100~F 1750~F - Oil Q. 1 (14,000) -- 6,000 1750 F - Norm. 1 16,500 -- 8,000 4,500 4,400 1750~F - Norm. 3(a) 17,000 -- ( 8,000) 4,500 4,500 1750~F - Norm. 6(a) 17,000 - ( 8,000) 4,500 4,500 1950~F - Oil Q. 1 (13,000)..... _ 1950~F - Norm. 1 18,000 -- 15,500 ( 8,000) ( 6,800) 2100OF - Oil Q. 1 16,000 -- -- 2100~F - Norm. 1 19,000 -- 15,500 ( 8,000) ( 7,100) (a) 1-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycles of normalized 3- and 6-inch diameter bars. ( ) Indicates values estimated from insufficient data.

TABLE VIII Rupture, Total Deformation, and Creep Strengths at 11000 and 1200'F for "17-22-A"'S Steel as Influenced by Heat Treatment One-Percent Total Rupture Strength Deformation Strength 0. 001%o/Hour Bar Diam. (psi) (psi) Creep Strength Heat Treatment (inches) 00-hr 1000-hr 100-hr 1000-hr (psi) 1100~F 1750~F - Oil Q. 1 30,000 18,000 26,000 15,000 15,000 1750~F - Norm. 1 44,000 18,000 (34,000) 18,000 22,000 1750~F - Norm. 3(a) 42,000 20,500 -- 21,000 27,000 1750~F - Norm. 6(a) 43, 000 21,000 -- 21,000 27,000 1950~F - Oil Q. 1 31,500 19,500 -- 19,500 24,000 1950~F - Norm. 1 38, 000 20,000 -- (20,000) (27, 000) 2100~F - Oil Q. 1 31,500 19,000 -- 18,500 (22,000) 2100~F - Norm. 1 38,000 21,500 -- -- (35,000) 1200~F 1750~F - Oil Q. 1 12,500 6,400 9,000 -- 1750~F - Norm. 1 17, 000 7,200 16, 000 -- 1750~F - Norm. 3(a) 16,000 ( 8,500) 15,000 7,500 1750~F - Norm. 6(a) 16,000 ( 8,500) 15,000 8,000 1950~F - Oil Q. 1 17,000 7,400 13,500 3,000 1950~F - Norm. 1 15,000 8,600 -- 8,500 2100~F - Oil Q. 1 15,000 7,600 12,000 6,000 21000F - Norm. 1 17,000 9,400 -- (10,000) (a) 1-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycle of 3- and 6-inch diameter bars ( ) Indicates values estimated from insufficient data.

TABLE IX Rupture, Total Deformation, and Creep Strengths at 11000 and 1200'F for H-40 Steel as Influenced by Heat Treatment One-Percent Total Rupture Strength Deformation Strength 0. 00 1% /Hour Bar Diam. (psi) (psi) Creep Strength Heat Treatment (inches) 100-hr 1000-hr 100-hr 1000-hr (psi) 1 100~F 1750~F - Norm. 3/4 29,000 (14,000) 24,000 (17,000) (17,000) 1950~F - Oil Q. 3/4 42,000 29,000 35,000 23,000 25,000 1950~F - Norm. 3/4 41,000 29,000 39,000 28,000 31,000 1950~F - Norm. 3(a) 44,000 34,000 40,500 29,000 31,000 1950~F - Norm. 6(a) 42,000 32,000 37,500 30,000 31,000 2100~F - Norm. 3/4 (49,000) 36,000 (48,000) 33,000 32,500 1200~F 1750~F - Norm. 3/4 (13,000) 1950~F - Oil Q. 3/4 23,000 -- 1950~F - Norm. 3/4 25,000 —. 1950~F - Norm. 3(a) 27,000 __ 1950~F - Norm. 6(a) 28,000 -... 2100~F - Norm. 3/4 29,000 -- (a) 1-inch diameter bars were air cooled in insulating firebrick cylinders to simulate cooling cycles of normalized 3- and 6-inch diameter bars. ( ) Indicates value estimated from insufficient data.

2000 1200 800 400 20 40 60 80 100 120 140 10 Time - minutes Figure 1. - Cooling Curves for the Centers of 1-inch, Simulated 3-inch, and Simulated 6-inch Rounds of "17-22-A"S Steel Cobled from 1750~F in Air.

2000 1600 1200 4.a 800 400.. 0 20 40 60 80 100 120 140 160 Time - minutes Figure 2. - Cooling Curves for the Centers of 3/4-inch, Sivnulated 3-inch, and Simulated 6-inch Rounds of H-40 Steel Cooled from 19500F in Air.

0 tn~~~~~~~~~~~o.re s~~*e H~~-40 8 101 1600 1700 1800 1900 2000 2100 2200 2300 Temperature ('F) Figure 3. - Effect of Austenitizing Temperature on the Bainitic Grain Size of Normal.. ized 1-inch Diameter Bars of SAE 4340, "17-22-A."S, and H-40 Steels. (Note:- Because of the difficulty in ascertaining the grain size of the "17-22-A"S and 4340 steels in the normalized condition, a band representing the predominant bainitic grain size is shown.)

X100D X1000D 44'A~~~~~~~~~~~~~~~~~ v\' <a' >ii ~.4 ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~v ~~~~:~~~~~~~isl~~~~~~~~~~~~~,~~ ~~~~~CY~~~~~',::- a~.,..'"':.',..,'.., c~~~~~~~~~~~~~~~~~~~~~~~~~~~.~~..:.... 6~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~a~~~~~~~~~~~~~~:~ ~ ~-~..:: ~~~~~~~~~~~~~~~~~~~~~~~~ ~~~~~~~~~~~~~~~.::!~."'":...: -'s;..i~~c di~ I~~~~~~~~~-;EI::~~~~~~~~~~~~~~~~~~~~~..',~:~ 1.17~ (a) Normalized from 1750~F as 1-inch diameter bar stock I1390 BHN tested* 13~ hrs at 00~Fan 2,00 si - 0BH gre r tize 4~~~~~~~~~~~~~4 (b Nrmliedfrom 1750FF +b AsTempered 1hato 100 300 BHNnd()AtrCep Rutretn t100~, 3;~-'Q_ -~'~*:i$ E~ b~~~i~~~ s,~. rl t,:~~~~~~o! 5 tested ~ ~ ~ ~ ~; 1392 hrs at 1000 an 0 0 s 6 H from 1750F, (b) As Tempered to 300 BHN, and (c) After Creep(R) Normalized from 175Tsi at temper0d00Fr.at 110 F + crep-ruptur tesed 392hrsat 000F ad 2, 00 pi -260BH

X100D X1000D R~~~ui~~~~:i~~~~j~~~c tTC~~~~~~~~t~*?: A * ~~~~~~~~C *~~~~~O y vs," It rr~\'\ rz %~~~~~~~~~~~~~~~~~~~~~~~~ RZ, (a) Simulated 3-inch Diameter Bar- Stock Normalized from 1750'F 327 BHN':"....~ ":...:.:"<:.....Q ve 1171 x * )t. 4* ~~~~~~~X~~~~~~~~~~~~t ~~ ~ r, &sr...........p....~,%V A......-, ". v 3~~~~~~; (b) tNrmalized frmP70. crep-ubtretesed110+r*t100 n 20&."...p, 14444'W 20,000 psi 198 BEN Figure 5. - SAE 4'340 Steel. Simulated 3-inch Diameter Bar Stock (a) As Normalized from 1750"F and (b) After Creep-Rupture Testing at 1000~F.

XIOOID X IO000) "'~~~~~~~~~~~~~~.r ~ ~ ~ ~ ~ ~ ~ /.~~' y ~,,$S.t...'....~:,;,i.t.....'i~.,,4 A.' ~ c*'.:'~:~.~ -"-',":~ "Y'c';oil'.:"' 4~~~~~~~~~~~~~~~ and,2.tc;.t4.0 0 0r0 psyi ~ v -,.W- Z0'::)I'"'('144~~~~~~~~~~~~4214 (a)Fimue6 A: 34 te.~i.ulated 6-inch Dia.meter' B:::ar' Stockk Noralze frm150F 35B x~~'.1 A't x' 4,% "' (b) N ~ox mali4'ed hem ~ n 1 750%1; +n cree r f c'C'c.iupture' Te~ed 48hsati100g at 1000FIi,

XI00D XI000D;K V t~-7 — I~~~~~~~~~~~. ~~~~~~~~~~~~~~~ ~~~~~~~~~M nii.Iil P11,~' ~~~~~~~~1~~~~~4 W sA NrA ZY V ~ ~ ~ ~ ~ ~ ~ ~ ~' ~~~~~~~~~~~~~~~~~~~~~~~~~~~~ ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~1 (a) Normalized at 1950F~~~~~~~~~~- As1cIDaee BrSok 2 H A: x \-~~ i ——', Tfl, ~ ~ ~ ~ ~ ~ ~ ~ ~ ~' 471f~~ A j 2) it~ ~ ~~~~ 5m "vP V'r. I~ SWIM', ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ /-~— 4-~~~~~~~~~~~~~~~~~~~R (b) Normalized at 19500F + Creep-rupture tested 1425 hrs at 10000F and~~~~~~~~~~~7 2000 piy-28 H Fiue7 SE44 tel n dc Daee arSok()AsNraie at150 n ()atrCep-utr etiga 00F

X10QOD X1000D $34 w~~~~~~~~ A~~~~~f t-.~~~Vc ~~~~ wtr jt "ny~~~~~~~~~~~~~~~~~~~~~~~~~~r ~ H 5',a S.~~~~~~~~~~~~~~~~~~~~.i K'~~~~~~~~~~~~~4............. Mtt >; *.t A~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~t Kv ~~y r $ - 4$ t$#S~~~~27 t'O (b) Normalized at 2100W + creep tested 2257 hrs at 10000'F and 20, 000 psi - 222 BHN

XlO0OD X1000D (a) Oil-quenched from 1750'F as 1-inch diameter bar stock- 585 BEN',A li~~~~~~~A ~ % 4k~. &,~-~'. 4...~~~~A' ~ ~ *~ 4 —PtV... r. 77~~~~~~~~~~~~~~~~~~~~~~~~t;4-2 RWJ &t 4%na 14~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~4 K'.,. -.,.f -..?-...St.....~.. 4 t.4 r~~~~~~~~ c~~~~~~~~~~~..> -~~~~~~~~~~~~~~~~~~~~~. 2K' - -~~~~~:'.........~~:..;..~~. I 14~~~~~~~~~~~~~~~~~~~~~~0 ) Oil-quenched r K.S0' —t e. a 1 310 B~ tuo&'P, V&V' $ "y % ~' V.t$% ~ (c Ol-uncedfom170F tmerd 0 r a 10F cep ese 1025 hrs at lOQ~~~~~~tE' and 13,000 psi - 270 BEN~ ~ ~~~~~' Fiur 9.-SE444te.Oeic imee a tc a sOlQece frm15twbsTmee o 0 Ead()AtrCep Testing at 1000 F.~~~~~~~~~~~~~~~~~~~AJ

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X100D X1000D 71 y I~~~~~~~~~~~~~~~~~~~~~ (a) Oil quenched from 2~~~~~~~~~~~~100 As1icAimtrBa tc 3 H I O- ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~',',''~':.~.I-,.;.,..'.,.', ";'t' ~*....."''-':... ~:.'",":'~* ":' (c) Oil quenched from 2100'F +Aepee 10c himtrs Batr10F +tc cree-rutur Figure 11. - SA~~~Ep 430SelMn-nh imtrBrSok a sOlQece from 1000F, (b)'~ As'"'"'"':-.~' T m r to~ 300;., BHN,:' and" ".*.v.'.~ (c){' after.. Creep-m....':.::'.:...'!;Ruptur. Tstn at.00..... ~ ~,.~...~...:'-:'{"?;;~::":'~::'~"l ~'~"(~'*::';'?':i..7:~"::'.;'.,."i~ /-i~:::.2~....>.. ~..;:x.':".:~:..:,..~,..........::.,..,-~:......:..: ~,.{,. ~, V: ~ ~~~ ~ ~~~ ~ ~~~ ~ ~~~ ~ ~~~ ~~~~ ~~~~ ~~~~ ~~~~ ~~~ ~~~ ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ ""'~~~~"}I::N;: ".. (b) Oil quenche from" 20'F ""+ tepee;1" "is "'" I"* * 0'"" 300B".4:..::,.:.:.:.:..~:~::~,::~i.:::~, —~~~. ~:: t1AI:;.:~:.~~,:.?_.- /::.:..~ ~ ~ ~ W:t.'W.'::.'....';i,":.:;.:;:.i::*~:.'':~"..;~fe,~?:~,:::,;* ~.;:~:~..~';:~:,~';,::~:.:!:~:? ~'.., - ~<~:- 7::':'(''%.?:P!.:{I:%'N:;:;:~;,~'*~~~~:i..'":''*'" (hi~~~~~~~~~4 ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~',$"',."',..... Tested 548 hrs at 1000~F and 20,000 psi - 252 BHN

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X100D XlOOOD mete~~~~~~~~~~~~~~~~~~~~ 1r Ba (a) ~ ~ ~ ~ 5~ Siultemeed 36hDaee a tc Normas e fro 1700~F -32c5 BH N p'~~~ ~ A?~ ~~~~~o A A;@ j;~~~~~~~~ 4r~~~~~i~~~dY C;44 ~ ~ ~ ~ jti A, I~~~~~~ F4~~~~~~~~~~~ 74~~#1 R~~~~~~~~ (b) Normalized from 17500? + tempered 6 hrs at 1200 OF -295 BHN W, 4* A ~~R A 1p;~ ~ ~ ~ t". *.I,, >7; / ~~~~~~~~'A v~~~~~~r;v VIt:~;ji -%: /'0.,' tSJ ItT. 4T44t ~~~~~~~~~~~~~~~~~~~~~~~0.T (c) Normalized from 1750'F + tempered 6 hrs at 12000F+ creep-rupture Tested 1220 hrs at 1 100.? and 19, 000 psi - 257 BEN.

X1OOD X1OOOD.4,j,4 -W~~ "I' >s.*tA. Qs -. AA' t.'~~ ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~'a Ain~~~i (b)Siuae6inhDaeeBaStc Normalized from 1750 ~ eprd6 r t10 F - 329 5 BHN \ ~:':'_;'': v'i" "::'' ~~~~~~~~~~~~~:' at.:'..~;4 %7~~~~ IN N"'............ JW~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~." ~:'..' c~;~~...~~~"'.'."-' o''..,:,. ~ -:!!:: h~.,v- A, o.:!.? " k.:... ~~:,*::,~:.-~'~" ~t.-;-,t...~.... 2~ (c ormalized from 1750F2 + tempered 6 hrs at 12,00~'+ cre-ruptureH testeda 139 hrs at l00~F:an 19,000 p;s y- 25 BEN (c) Normalized from 1 F te750F ed ( Ars Tte' to...' 300 BENp-, an (c)e 10 afteratCreep-Rupture9Testingiat 1100FH

X100D XIOOOD 4N. ~ ~~~. R~~~~~~~ MJ~ ~ ~ ~ ~~~~~qo rmie 50~ Ia4 Normalized at 1950'F a ter Brs So 34 4 ~ ~ ~ ~ ~ ~~~~4'.~ ~-~..,:~:'~,.'.~~,'' 7j;''R'4,' 4'" (b) Normalized at 19500F + temerdI hrshaamte Bar00tock 302 BHN ~~~~~~~~~ ~~~~~~~~~~~~~~~~ T 4r A~~~~~ A (c) Normaized at 1500F + tmpered 6 rs at 1200F + crep-ruptur tete 125hs It10l Fad1,00pi-'5 H (b) ~Normalized at 1950'F epee (b Ars Tempered t 300 BHN, n (c fe re-utreTsiga 10F

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X100D XlI OQOD 5~ ~ N. (a) Oil quenched from 1750'F As 1-inch Diameter Bar Stock 524 BHN I: r~~~~~~, %'baia a V t4~~~~~~~~~~~~ad (b) Oil quenched from 1750'F + tempered 1Ihr at 1300'F 305 BEN N~~~~~ * 4. 4, V,42A~j 2~ gj~~~~ ~~44~~~V? nt.~~~~*A~il (c) Oil quenched from 1750'F + tempere-d 1Ihr at 1:300'F + creep rupture Tested 666 hrs at 1 1000F and 20, 000 psi - 265 BEN.

X100D X1000D S1494 4~~~~~~~~~~~~~~S~~~~~t~~~~ - t Stt.~~~~~~~~~~~~~0 4$.;4 45~~~~~~~~~~~~~~~~~ -s\ V.~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~::. - ft ~~~~~~~~~~~ ~~~~~~~~~~~~~~~~~~IN tZ_ Al. ~ ~ ~ ~ - ~t (a) Oil quenched from 19500F as 1-inch Diameter Bar Stock- 470 BHN -~~~ ~~~. ~ ~ ~ -fl ~~~~~~ *54,7~~~~~ XM,4 7 W A4 Aeeet (b) Oil quenched from 1950%' + tempered 1 hr at 1300%'F 300 BHN Z-A-~ ~ ~ ~~~SA& (c) Oil quenched from 1950%' + tempered 1 hr at 1300"F + creep-rupture tested 1081 hrs at 1 100%' and 19, 000 psi - 240 BHN

NH[ 0S7 - Tsd 000'6T Pue ei.oo0I ye s;tT ITOI p~aso aan~dn~-daaoa + elOO00I ye:it i pozodtuwa + fOOTI Luoij potpauonb ITO (3),.- ~ ~......4,.:'.~~~~~~~~~~~~~~~~~~..4''.::*4.'7..:.:.~~z:..:...:'.04.'/:"i.'4::,?:"::(' I~.' - V "J....5 N 6VIT~ ~~~~~V ~~~ ~~~L.~. NH9 06Z - o oo~i Tye a-t T paad-uai + a.0017 Lu0a p~aIz3u~nb ITO (q) 74, ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 4:4; U~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~int ~~ ~~~ ~~~Ar~~fe NH9 ZTT I U~ Ll3ut-T- Se 3o001Z ru o.IJ p~aM3u~nb ITO (e) sawMAi~ gr~~ooi Al~ oi

X100D XI OOOD (a Nomlie fo P90. s34,nhDaetrBrSocq 3 H ~~~~~..4~.~ I~~~~ At~ ~.;i.~, v b~,?7 "', _i t,4~~~~~~~~~~~~~~~~~~~~ (b) Normalized from 1950'F +aemeed1 hrsin aamte Ba00 Stc 316 BHN A {."%F " ~"~'" ~ " ~".~i~". ~%~'~, ~'.~. " "~'~~~, ~'~'~i~'q''~,'f"~'~,,'."t'~'......~'" ~". ~'""",'~'~' ~"" " J'.'C?",'.0'"' I.~'".,,. ~ "'~........'?%'~i~~'T','..,'-" j" ".. A, ~ ~ ~ ~ ~ ~ ~ ~ A teste 27 r a 10 ad3,00pi-8 H Figure 20"~''~'~!~i. — 40 Steel.. Three"%~".'' "i~i~'"~?~'-qat in Da e Ba So (a) ANormlize from,, (b) As T e to 300 BH, an (c) afte Cre-utr etnwt10F

X1OOD X1OOOD ~~~~~~~~~~~~~~~~~~~~~~~ ~ ~~~~~~~~~~~~ Mrv~,,.:.~~ -,:~:~~~....,.... * r~~~~~~~~~~~~~~~ -.,~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~A _ r~~~~~r W I P ~~ ~' 1444~ks& ~~ Itt & XI-N vt~ N,1 t'4 ~~~~~T, ~ ~ ~ ~ ~ ~ L *, ** A~~~~~~~~~~~~~~~~~~~~~~~ (c) Normalized from 1950'F +tempered 4 hrs at 12000? +9 creptete?*fr()afe Cep etigat100? W V,~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~-~/~ ~iii~~~~~~~~~~~~~~~~~~~~~~~~~~~~. s ":::'::~ =.'""...'V:. ".:,.'..;;j7,:'..'~..: ~~.....:,...:.'i'!,:.~;i(aed3ich imt) a t Normalized fromm1950~emeed4hs t100F+ re tse I[ iqn 00Fan3- 00 s-28 BHN

X100D XIOOOD "'& ~ 4 "'M:','~.....:~'".....:~!i' (a imulatedi 6-inch Diameter Bar Stck Normalie] rm150 9 H BeiAj'x<4Yg. 4. AP/'4' 1 * V'9 r..,.~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ ~~~~~~~ f~~~~w. t-t,,....~,.,'> t xt.' 4. tat j r t4....t.t.sn.........:,,~i~:., ~.'',.....r Qt"'S-''.'.,': - 44,~ ~ ~ ~ ~ ~ ~ ~~~~~4~ N.~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ L, fall~ ~ ~~~~~ 4,y - 41 ~~~~~~~~~~~~~~It IN~~~~~~~v 0 -1 -:V (c) Normalized from 1950"F + tempered 4 hrs at 1200"? +2cee teste 1318 hrs at 1100,? and 30,000 psi - 243 BH Figure 22. -11-40 Steel. Simulated 6-inch Diameter Bar Stock (a) As~~~~~~~~~~~~~~~~~~~~0 ol Normlize fro 195F, () AsTempred o 30 BHN an (c) after Creep Testing at 11000?.~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

XLOOD XlI OOOD 49~~~~~~~~~~~~~~~~~~~~~N V4,4 4 J j~~~~~~~~~~~~~#~~ts 1 14 s W A~~ 4 C, >t~~~~~~~~~~ w'C~~~~ (a) Nrmaie fro M70N s34-nhDaetrBrSoc, 0 H j~str. ~-~'~ -IC >~~~~~~~~~'.4< 4!~ ~ ~ ~~~~~ "S *.-1~4lC~ ~ ~ ~ VI (b) Normalized from 1750WF +aepees hrsinh aamte 1200 313c BHNB ~~:.n~~~'tk yts.1~~~~~~t~~t~~<;yiP $let~~~~~~~~~~~~~~~~~~~ ~~"~~r~~-'"-&''~~~~~":- ~ ~~~ itl RL ~ ~ ~ ~ ~ ~ ~ ~.~A RU ~ -. r —, 4i~~~t47~ "~~~~~'~~~"4-~ ~ ~ ~ ~~~' ~~~r 4>~~~~~~~~ — ~~~~~~~ I -~~~~4 (c) Normalized from 1750WF + Tempered 4 hrs at 1200WF +3cee rupur tested 83.7 hrs at 1100W and 30 000 psi - 222 BHN~~~~~~~~~~~~~~~~~~~~~~~~~~~~7. Figure 23. - 1-1-40 Stel. Three-quarter inch Dameter Bar Stock (a) A Normalized from 1750W, (b) As Tempered to 300 BHN, and (c)aferCrepRutur Tstngat110W

X1OOD XO1000D 4 4Ntt N't $~~~ ~~ ~~~~ r" ~''~..... $i ". "~i.' I ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~* (a) Normalized from Z100F as 3/4-inch Diameter Bar Stock -429 BHN i, c "~~~~~~~~~~~~~ ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~t~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~...... ti' iajhu~~~~~~~~~~~~~A t A.~~~~~~~~A I W~~~~~~~~~~~~~~~~~':M ~'.:Vr-X'z"~~': ~"' Ilk: 6:,,/ ~~>\ ~ $*w ~~'C t~q-%vt Y~ky4LX~W $ et' (b) Normalized from 2100'F + tempered 6 hrs at 1250 t'F 319 BHN 7r~~ ICA (c) Normalized at 2100tF' + tempered 6 hrs at 1250'F + creep tested 192?2 h rs at I 110 a''n- l'nd 3,00 psi- 240 BHN

X100D XOD C &.. 4~1... <'r i,~,ij'i:::/,. * ~!~:."~6. ft 4< ~ \v.~~: PC 4'rt, / " 4.?\! zt # I...... ",~~~~~~ #~,..~.': A.',-,'ft \ A.. ",,~.,%. A.._ ~ 4.,,"~-.; -~ —. — _. I. -4 _:,.-..,-,. ~.: < 1 4::,..:., f.;~~~~~~~:,... I., ~ ~ ~ ~ ~ ~ _, t 4 4 "v A~~~~~~~~~~ ft;..: K.S!~i~i:. (a) Oil quenched from 1950~F~ as 3/ nhDamtrBrStc 2 E,.,: — ~,-,~:, ~ ~ ~,'~~~... s.':.. &~~K >If 344 -....'....5#O 0~ ";~_J %#,~i;4. —` A_,',,!:'.,"',~!:~~,~,..~~":~~,:,.~. -1::,. t,,,f,+S ~ M~S e*",&. A: ~ -:` ~ "4 iS...4"e. 1,. f1. ~ i.~:" 7.',_..'%, ~~i, _: ~A' &lc -At.t'% jk4% " "'''~ A" 1j.;,nt..4. ~~~~~- ~~"4''n' -. t."., Io. I tk~~~,.....',.,~~'! ".~,', 4/r *. - ~Ths>&\M. Oil Quenched frm190,(b:s epre.o 0 EN n ~~." c)aferCrep-upur Tstig t 10~F

100, 000..... 80,000 Normalized Wheel - - -Oil Quenched Wheel I 1 900F 60,000 1750~F - O.Q. - I inch round A o A O'00~_., 40,000 0 1750~F - Norm. -I inch round - w 1~~~~~~~~~~~~~~~~~~~000'r. 40,000 H 1750~F - Norm. -3 inch round'" — Im )l | < 1750~F - Norm.-6 inch round I - I _1-_ _ 30,000 150F - O.Q - inchround - 20,000A 19500F - 0.0. -1 inch round En l |O 1950~F - Norm. -1 inch round - - | - 1. 20,000 2Z100~F - O.Q. - 1 inch round ~ A 2100~F - Norm. -1 inch round - 1 10, 000 Fgr 6 l~~ooo0~l I liliillo I I l!;o ('I l 1 11000 Time - hours Figure 26.- Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of SAE 4340 Steel at 900~, 1000~, and 1100~F. Curves for SAE 4340 Turbine Wheels Included for Comparison.

70, 00 ~ ~ Normalized Wheel...... Oil Quenched Wheel 60, 000 0 1750F - 0. OQ. - Iinch rot n _ 1750~F - O.Qrm. - 1 inch round 1750~F - Norm.-3 inch round VL CIA 0 900':F 1750~F - Norm.-6 inch round 0 195 OF. Q. -Iinch round |* ) | |50 |O||900~ 0 1750~F - Norm.-6 inch round |1950~F - O.Q. - I inch round ~~~50,00q~~ l l l l l l l l l l l 5 e e -- | ~ 1950~ F - Norm.-I inch round 40, 00 - -- 5.0 0.10~100 -. 1inro000 0 OA ~ ~ ~ ~ 100Time - hours Figure 27. - Effect of Various Cooling Rates and Austenitzing Temperatures on the Time to 0. 5 Percent Total Deformation Data for SAE 4340 Steel at 900, 1000, and 1100. Curves for SAE 4340 Turbine Wheels Included for Comparison. 3~~~~~~ ~~~ ~~~~~~~~~~~~~~~~~~0, 00D c 1~~~~~~~~~~~ O, 000 -00 Time -hours Figure 27. -Effect of Various Cooling Rates and Austenit,.zing Temperatures on the Time to 0. 5 Percent Total DeformnationDt for SAE 4340 Steel at 900', I0000, and 1100~F. Curves for SAE 4340 Turbine Wheels Included for Comparison

70,000____Normalized Wheel -- - Oil Quenched Wheel * 1750'F - O.Q. - I inch round 60, 000 - o 1750F - Norm. - 1 inch round 900F | 1750F - Norm.-3 inch round A1 g O n | 4C 1750~F - Norm. -6 inch round * 19500F - O.Q. - I inch round 50,000- 0| 1950~F - Norm.-I inch round A 2100~F - 0.Q. - I inch round 2100F - Norm. -1 inch round 40,OO __ _ 00. -. - 10004'F C230, 000-_ 2 0,000 e L ~;j. -A 1 00- -4 0 10,000 oo N~~~,,~~~~ L~ 10~~~~~~~~~~~~~~~ O.- ~~~-~ 01 ~~~~~~~~10 100 1000 Time - hours Figure 28. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 1. 0 Percent Total Deformation Data for SAE 4340 Steel at 900-, 10000, and 1100l F. Curves for SAE 4340 Turbine Wheels Included for Comparison.

100,000 80,000 - - 900'F 60,000 13 A _ _ - -, 40,000 -AE -. O -IO__ 30, 000 - - i i - - i —,, 3,8000-n.o 6 000 100'F 0 750F - Norm.alized Wheel 4.,'~ ---- - Oil Quenched Wheel 2,000 -'- - 3 *| 1750-F - O.Q. - I inch round O 1750'F - Norm. -1 inch round 3, 000 ~C 1750'F - Norm. -3 inch round 1750~F - Norm. -6 inch round zooom 1950'F - O.Q. - I inch round 0 1950eF - Norm. -1 inch round A 2100~F - O.Q. - 1 inch round A 2100~F - Norm. -1 inch round 0. 0001 _ 091 0.01 0.1 Minimum Creep Rate - Percent per hour Figure 29. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for SAE 4340 Steel at 900~, 1000., and 1100F. Curves for SAE 4340 Turbine Wheels Included for Comparison.

1500 0.0003 700'F 0 0 1000,O0. 0002 Feor ~~~~~ t ~~~~~~~~~Time for 1 Percent Total Deform k ~~~~~~~~~~~~~~~ation at 90, 000 psi 0 500 0. 0001 0 0 80,000 900'F 1000-hr Rupture Strength 60,000.e 0~~~~~ 60,000 ~ 0 Stress for 1 Percent Total DefI~~ @~~j ~ormation in 1000 hours 40,000:a 20,000 60, 000 * I utr -1~000*F I. *... —--— e ~~~~~~~1000-hour RpueSrnt 40,000 - 0.2fi_ Ofifi tg s _________ i~~~~~~~~~Stress for 1 Percent Total Df 0- 1" iarn. 1" Di:. 3" am. jrmatonmaion 10hus Oi 0.10Norm.r Normre Normgt 0,000Meim nth ih epeaur roetes ofSA a34 teelf at000 to 1100F. ~ ~ 10hor utue tent

1500.0003'4 0 7 00 F 00 5000 9.oo: t90OOp 0 0. U 0 0 1750 1950 2100 80,000 9000F 60,000 - 1 0 1000-hour Rupture Strength 0. ~~0 * I I *. ~~~~~~~~~q tress for I Percent Total Deformation ________________ ~~in 1000 hours 40,000 _____ -..A..-.. A.Stress for Creep Rate of 0. 0001 Percent Iper hour 20,000 1750 1950 2100 60,000 1000*F 0 I 0~~~~~~~~ 1 00-hour Rupture Strength 40,000 0 * ~ ~............! _____ —--— 1-(Stress for 1 Percent Total Deformation 0 ~~~~~~~~~~in 100 hours *.~~~~~~~~~~~~~1000-hour Rupture Strength *k Stress foK Creep Rate of 0. 001 Per0- ~~~~~~~~~~cent per hour 20, ___________ ~Stress for 1 Percent Total Deforma0 I ~~~~~~~~~~~~~tion in Il000 hours 20,000 01 ______ 00-hour Rupture Strength,4 -Li~~~~~I 0 Stress for 1 Percent Total Deformation in 100 hours 10,000 ciJ{ 0 _ 1 _ _4_ ____ 0170 80 1000 20 2iNormliingTepeatue *

60,000 40,000 o. I-F __|| 30,000 Normalized Wheel -, -Oil Quenched Wheel 20,0000 1750F - O.Q. - I inch round' 0O 1750'F - Norm. -I inch round 1200-F [-_ * O 0 1750F - Norm. -3 inch round bm ~ |1750F - Norm, -6 inch round ~e 10,000 O 1950-F - O.Q. - 1 inch round 8,000 1950'F - Norm.-I inch round A R.10 0'F - O.Q. - I inch round -"" - - a 6,000 A&2 100*F - Norm.-I inch round __ 1 0 1000 Time - hours Figure 32.- Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of "17-22-A'S Steel at 1100- and 1200~F. Curves for "17-22-A"S Turbine Wheels Included for Comparison. 60, 000. Normalized Wheel ---- Oil Quenched Wheel 0 1750F - O.Q. - I inch roaund 50, 000 0 1750eF - Norm.-I inch round 5| 1750-F - Norm.-3 inch round 1750'F - Norm. -6 inch round I400 - - l I _ LL | o( | |_ | | 1950-F - O.Q,. - I inch round _ 40F000 - - a | | | z' l l l l l l l l ( | 0 1950-F - Norm.-l inch round 1100'F A 2100'F - O.Q. - l inch round eU~~~~~~ | l l | ~ T | || llOO-A 2100F -Norm.-l inch round ao ooo,,' ~ 30,000 -. —--- -. 20,000 - _ _ - - A - 10,000 I I - 1 10 lUU_ —L _ 0 Time - hours Figure 33. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 0. 5 Percent Total Deformation Data for'1t7-22-A"S Steel at 1100. and 1200F. Curves for "17-22-A'S Turbine Wheels Included for Comparison.

40, 000 ____ Normalized Whe - - - Oil-Quenched Wheel 1750 F - O. Q. - 1 inch round I___ 30, 000 -0 1750"F - Norm. -1 inch round 0 1750'F - Norm. -3 inch round 5~1750"F - Norm. -6 inch round I1I00'F s U ~~1950"F - O. Q. - 1 inch round k20,OO 13 1950'F - Norm. -l inch round*A 21 00 F - O. Q. - 1 inch round A. *m 0-' 2 100'F -Norm. -1inch round 10~ _ 10, 000 __ QE A 0 3 I 10 100 1000 Time -hour s Figure 34. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 1. 0 Percent Total Deformato Data for "117-22-Al'S Steel at 1100' and 1200'F. Curves for "117-22-A"'S Turbine Wheels Included for Comprisn

50,000 40, 000 - -0 _-_ ). —.30,000 20,000 O t') 10,000 Normalized Wheel - - - Oil Quenched Wheel A | 1750'F - O.Q. - I inch round 10,000 1750~F - Norm. - I inch round 8 1750-F -Norm.- 3 inch round - I - I *' 1750'F - Norm. - 6 inch round _ A _ |- b 6,000 *1950-F - O. -1 inch round E01950'F - Norm. - 1 inch round &2100'F - O.Q. - 1 inch round 4,000 A2100F - Norm. - 1 inch round 0. 000.0001 0.0001 0.001 0.01 Minimum CreepRate (Percent per hour) Figure 35. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for "17-22-A"S Steel at 1100. and 1200-F. Curves for "17-22-A"S Turbine Wheels Included for Comparison.

3,~000 0 0. 00031 2000|~~ 0003 ~ l\ -Time for I Percent Total Deformation,00 I J at 102,000 psi 4.ooo, I 1 1, 000 I 0.001 t. / Minimum Creep Rate - Percent per 0 0 1__ ~ hour at 70, 000 psi ~~~,0 0 v _ _ 4,000 0.0004 900"F 60.00384 14 0 ~~ )C ~ — 0 100-hour Rupture Strength 3,00000 / 000 0 2,00 Stress for 0 5 Percent Total Deformation 14 4. at 70, 000 psi 20000 O. I000I2to n10-or 1, 000 $4 0. 0001._____Stress for MinimumCreep Rate of 0001 Percentper 0 60,0000 _______ ~~~~~~~~~hour at 70, 000 psi _ _ I Ierou 0 B y l — _..~.-~~ 8100-hour Rupture Strength 40,000 -4 A-I/' B - Stress for 0.5 Percent Total Deforma~0.4 ~ ~ ation in 100 hours,0Stress for Creep Rate of 0.001 Percent 1 e ~00-0X i p 100-hour Rupture Strength ~0 20, 000 _____ _ 0-O 1000-hour Rupture Strength Oil f. N0 rcn Total Dorm1" Diam. 1" Diam. 3" Diam. 6" Diam. *HeatTreated Condition Figure 36.- Influence of Cooling Rate as Controlled by Section Size and Quenching Medium 0on the High Temperaupture Properties of "17-22-A"S Steel at 700h to 1200'F. A

3,000 140.O0003 o 700'F'4I 3, e~~~ 000'~,,I~ - O03L Time for I Percent Total Deformation.,2,000 0.0002 | ____ ____ at 102,000 psi 1,000 o 0.0001 o oo., 04 o::1 u _ 0 0~~~~~~~~~ - _ i 0 Minimum Creep R~te - Percent per 0 10 j hour at 10". 000 Rsi 1700 1800 1900 2000 2100 4,000 0.0004 9003F'4 0.0 3,000 0. 0003 O —--, U + RTime for 1 Percent Total Deformation e V~~~~~~~~~ ff | ~ourat 70,000 psi V~~~ 2,000 0.0002 90' I-' 3 o 0.0 o03 m o I1 1,000 U0.0001 I I _____I 0 Minimum Creep Rate - Percent per hour at 70, 000 psi 60,000 40,000 0 100-hour Rupture Strenth I } 2Stf~essifoy0%. StPercent Total DeformatD 00.~~ 0 t~S~ imtress for Creep Rate of 0.001 PerI| cent per hour A I | 1 1 ^ 1000-hour Rupture Strength ^ 20,000 - - I -, Stress for 0. 5 Percent Total Deforma20,00: 1'____ 1 tion in 1000 hours 20,000 0 IO [0 0 100-hour Rupture Strength 110. F 10,000 0 0 1000-hour Rupture Strength 700 1800 1900 2000 2100 Normalizing Temperature - *F Figure 37.1- Influence of Normalizing Temperature on the Elevated Temprature Properties of 17-22-A'S Steel at 700 to 1200F. ~~~~- 0 0 h orRpueSent ni ~ ~ O

40,000O O~I I I I 40, 000 oNormalized Wheel 0 -- - - Oil-Quenched Wheel E 030,000 H0 1750~F - Norm. -3/4 inch round O0 _ — 0 I 1950~F - O.Q. - 3/4 inch round 1100~F 20, 000 [3 1950~F - Norm. -3/4 inch round 1950~F - Norm. -3 inch round ( 1950~F - Norm. -6 inch round A 2100~F - Norm. -3/4 inch round 10,000 I I I I 1 111 I I I - 60,4000 0- 1 - - - _____ ~~1200~F __ 40, 000 30, -oo - 0 U 0 20,000 10,000 1 1 0 0 1000 Time - hours Figure 38.- Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Rupture Properties of H-40 Steel at 1100~ and Z1200~F. Curves for H-40 Turbine Wheels Included for Comparison.

60,000 __ ~v -Oil-Quenched Wheel o0 17501F - Norm. -3/4 inch roun 50,000 __ - * 1950*F-0. Q. - 3/4 inch roun o1 1950OF - Norm. -3/4 inch roun 11000F i 1950iF - Norm. - 3 inch roun 1950*F - Norm. -6 inch roun 40,000 — 3 0 0A~2100OF - Norm. -3/4 inch roun co ~~~~~12000FH 30,000 7 7I1 Figre 9. or1 0 j1~hrtI00 II __ ]Fiure39.- Effect of Various Cooling Rates and Austenitizing Temperatures on the Time to 0. 5 Percent Total DeformainDt frH-40 Steel at 1 100' and 1200'F. Curves for H1-40 Turbine Wheels Included for Conmpe-rrson.

50,000, I,i,,i,, -- - Oil-Quenched Wheel I 1 i T 0 1750F - Norm.-3/4 inch round 1100F 40,000 A 1950F O.Q. - 3/4inchround _ 71'13 1950~F - orm. -3/4 inch round |.~. |s 1950*F - Norm.- 3 inch round I X( 1950'F - Norm. - 6 inch round 1200~F ~, 30, oo 00 4 3 ~ A 2100'F - Norm. -3/4 inch round 1. 30, 000 d I I -I 20,00_ 10 100 1000 Time - hours Figure 40. - Effect of Various Cooling Rates and Austenitising Temperatures on the Time to 1.0 Percent Total Deformation Data for H-40 Steel at 1100- and 1200F. Curves for H-40 Turbine Wheels Included for Comparison. 80,000 ||9I |- 00 | IIo F z~~zzz~~~zlz~~~zL~~bif4Kzzz II~~~~~~~arnj-~ 60,000.... 0.011 40,000 - Oil-Quenched Wheel -0~ | |O 1750-F - Norm. -3/4 inch round |0 30,000 1950F - O.Q. - 3/4 inch round - - 0 1950F - Norm. -3/4 inch round.| b 200o00 3 1 1950-F - Norm.- 3 inch round - | 1950-F - Norm.- 6 inch round 1 | Z100'F - Norm.-3/4 inch round 01,0.000!! Ii ll 0. 00001 0. 0001 0.001 0.01 Minimum Creep Rate - Percent per hour Figure 41. - Effect of Various Cooling Rates and Austenitizing Temperatures on the Stress-Creep Rate Data for H-40 Steel at 900' and 1100-F. Curves for H-40 Turbine Wheels Included for Comparison.

3, 00O~0. 0003 0.4700"FFI'4 Time for 1 Percent Total Deformation O, at 90, 000 psi.02,OO0~0 O. 002 14 4) O),000 go. 0001 o0 Minimum Creep Rate -Percent per hour at 90, 000 psi 0 0 3, 000 ~0. 006 900"F'4I 14 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~~~~STime for 1 Percent Total Deformation o 2,000 ~0. 004 - ___-at 65,OO00psi 14 4) S 1,00 pO002 0 --— __O__Minimum Creep Rapur trenath6,00 60, 000 ~'~b~reuufor 1 Percent Total Deformation I~~~~. a ~~'osrRupture Strength 0 ~~~~~~~ ~~Stress for Minimum Creep Rate of U) ~~~~~~~~~~~~~ ~~0. 001 Percent per hour __________ ~Stress for 1 Percent Total Deformna20, 000 tion in 1000 hours 3000 12001F T~...........1o00.hour Rupture Strength *20, 000 -_ _ _ _ 10,000 Oil Q. Norm. Norm. Norm.

3,000. oo0003 _7001F O. w *,*Time for 1 Percent Total Deformation 2,000.-0002 at 90,000 psi, 0 0. 2~ 00!,00 0~~~ 1, 000 ~ 0001, " k 0 h ~O\ U O Minimum Creep Rate - Percent per hour at 90,000 psi 0 0 3,000 O0.006 I., ~ 0 900F (U zs~~~~~~~~~ Time for 1 Percent Total Deformation.o c~ ~ ~~~~~~ \ / ~at 65, 000psi 2, Z000'0. 004 1 1, U 1,000 g0.002 o_ U Minimum Creep Rate - Percent per 0 0 -, * " hour at 65,000 psi 60, 000 1 100F 100-hour Rupture Strength Stress for 1 Percent Total Deformaa'~~~ 40,000_..~..~ 8 tion in 100 hours'40,000' -0.~~~ * 1000-hour Rupture Strength * X ~ ~ ~ / -- __ Stress for 1 Percent Total Deforma-'X~0 i i _- -=.tion in 1000 hours 0 [5-' I\Stress for Creep Rate of 0.001 Per20,000 cent per hour 30,000 Normal g0 100-hour' Rupture strength 0.~~~~~~~ ~20, 000 _ _ _ _ _ _ _ _ ). -~2o o0 ~ 0~0 1700 1800 1900 2000 2,100 Normalizing Temperature -' Figure 43. - Influence of Normalizing Temperature on the Elevated Temperature Properties of H-40 Steel at 700 to 1200~F.