A Low Temperature Brittle Fracture Investigation of Gray Iron and Nodular Iron on Machinery Casings

75-GT-80 Copyright 1975 by ASME $3.00 PER COPY author(s). $1.00 TO ASME MEMBERS The Society shall not be responsible for statements or opinions advanced in papers or in discussion at meetings of the Society or of its Divisions or Sections, or printed in its publications. Discussion is printeo only if the paper is published in an ASME journal or Proceedings. Released for general publication upon presentation. Full credit should be given to ASME, the Technical Division, and the A Low Temperature Brittle Fracture Investigation of Gray Iron and Nodular Iron on Machinery Casings R. EISENSTADT Professor of Mechanical Engineering, Union College, Schenectady, N.Y. Mom. ASME G. H. ROULSTON Design Engineer, KARL, General Electric Co., Schenectady, N.Y. The resistance to brittle fracture of two iron alloys commonly used in machinery components is investigated. The Stress Intensity Factor (KI) is calculated using the principles of Fracture Mechanics, and is compared to available fracture toughness data in the form of Charpy Energies, and Plane Strain Fracture Toughness (Kl c). Also presented are appropriate correlations to relate Charpy Energy and Kl c. The critical flaw size (i.e., the flaw size necessary for brittle fracture) is also calculated and is presented in a generalized form which is applicable for various materials. The results of the investigation are evaluated, and the value of examining both Charpy and Kl c data, particularly for irons, is discussed. Contributed by the Gas Turbine Division of The American Society of Mechanical Engineers for presentation at the Gas Turbine Conference & Products Show, Houston, Texas, March 2-6, 1975. Manuscript received at ASME Headquarters December 2, 1974. Copies will be available until December 1, 1975. THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS, UNITED ENGINEERING CENTER, 345 EAST 47th STREET, NEW YORK, N.Y. 10017

1 A Low Temperature Brittle Fracture Investigation of Gray Iron and Nodular Iron on Machinery Casings R. EISENSTADT G. H. ROULSTON INTRODUCTION Applications in Artic Environments have led to increased probability for brittle fracture failures. High operating stress in the presence of flaws, combined with temperatures as low as -50 F (-46 C) can create the conditions which lead to brittle fracture. It is the purpose of this investigation to examine this phenomenon for two materials commonly used in machinery components, nodular or ductile iron, and gray cast iron. One method of approach will use the principles of fracture mechanics to determine the Stress Intensity Factor (KI) for a given flaw size and compare this to available. fracture toughness data for the material. A second approach will use the available fracture toughness data to calculate a critical flaw size necessary for brittle fracture. This may then be compared to the maximum flaw size which is considered acceptable, CALCULATION OF THE STRESS INTENSITY FACTOR K I In order to calculate K I, it is necessary to have an estimate of the expected flaw size and shape, as well as an estimate of the nominal stresses encountered during operation. The type of flaw which will be considered here will be a buried or internal flaw, rather than a surface flaw. This is a common type of flaw for machinery applications, and of more concern than a surface flaw since it is more difficult to detect. Reference ( 1 ) 1 gives the following equation for a buried flaw under a stress normal to the flaw cross section. KI a Qa (1) where Q is a flaw shape parameter defined by the curves shown in Fig. 1, FLAW SIZE AND SHAPE Fig, 2 shows a cross section of a nodular iron casing containing flaws, The flaws shown in this figure are spongy shrink areas and are typical of both nodular and gray iron components, Flaws as large as 1/2 in. x 5 in. (12.7 x 127 mm) have been found in nodular iron casings, while in gray iron casings flaws 3/4 in. x 3 in, (19.0 x 76.0 mm) in cross section have been observed. The spongy shrink areas are groups of small flaws separated by ligaments of material, It is likely that early in the life of the component, these ligaments will fail, forming one large flaw, These shrink areas can best be modeled as elliptical flaws, (see Fig. 2) due 1 Underlined numbers in parentheses designate References at end of paper. NOMENCLATURE a = half minor diameter of flaw-(in,)-(mm) a c = critical minor diameter of flaw-(in.)- (mm) C = half major diameter of flaw-(in,)-(mm) KI = stress intensity factor-ksi (in,) 1/2 1/2 ) -(MPa(m) KI = plane strain fracture toughness, (slow cloading)-ksi (in.) 1/ 2 -(MPa(m) 1 / 2 ) = applied stress-(psi or ksi) - (MPa) Q = flaw shape parameter (dimensionless) f = ratio of applied stress to yield strength (dimensionless) KId = plane strain fracture toughness dynamic loading ksi (in.) 1 / 2 -(MPa(m) 1 / 2 ) a ys = yield stress - ksi - (MPa) 'max = maximum stress - ksi - (MPa) 2

0.5 E i 2a INTERNAL FLAT! 0.4 0.0 a2c RATIO 0.2 r6- y 5=.8 6s O ^y 3.G 0.1 0 0.5 1.0 1.5 2.0 2.5 Fig, 1 Flaw shape parameter curves r to their shape and the numerous solutions available for a buried elliptical flaw. Fig, 2 Examples of spongy shrink areas OPERATING STRESSES Nominal operating stresses for nodular iron components appear to be higher thaii for gray iron components. In the case of the nodular iron casings, stresses, excluding thermal stresses, normal to the shrink areas are as large as 11,000 psi, (75,1 MPa). Stresses in the gray iron casings are limited to 3000 psi, (20.7 MPa), tension and 4500 psi, (31.1 MPa), in bending, While the bending stresses do not strictly apply to the model chosen (i.e., a uniform stress field normal to the flaw) the bending stress limit of 4500 psi, (31.1 MPa), will be used here as a conservative estimate of expected operating stresses, PHYSICAL AND MECHANICAL PROPERTIES The mechanical properties of both gray and nodular iron vary considerably depending upon the chemical composition of the particular alloy. The chemical composition of a nodular iron and a gray iron for which both mechanical properties and fracture toughness properties have been obtained is shown in Table 1 rreference (2)a STRESS INTENSITY FACTOR If it is assumed that the spongy shrink areas are buried elliptical cracks, then a value of Q in equation (1) may be obtained from Fig. 1. Flange and casing defects in nodular iron casings have been known to have elliptically spongy shaped defects as large as 1/2 in., (12.7 mm), minor diameter by 5 in,, (127 mm), major diameter. For this size flaw, a = 1/4 in., (6.4 mm), 2c = 5.0 in., (127 mm), and a/2c = 0.05, Assuming a stress level of 11 ksi, (75,1 MPa), and a yield strength for this material at the temperature of concern (-50 F), (-46 C) of 54.5 ksi, (375 MPa),a /vys = 0.20, then Fig. 1 shows that Q = 1.0. Substituting into equation (1) yields: K I = 11 ksi1 25 = 9.75 ksi in (ID i M Pa rm ') For the flaws in gray iron similar to that shown in Fig, 2 where the minor diameter = 0.75 in., (19.0 mm), and the major diameter = 2,5 in., (63.5 mm), a = 0.375 in., (8.5 mm), 2c = 2.5 in., (63.5 mm), and,,a/2c = 0.150, At a stress level of 4500 psi, (31.1 MPa), with ays = 37,300 psi, (256 MPa),c /Qys = 0.12. Fig. 1 shows that Q u 1.18, and equation (1) yields: K I = 4.5 ksi ^g75 = 4.9 ksi in MPaFr- TOUGHNESS DATA Most toughness data available for these materials is in the form of Charpy impact data, (3-7) with some limited Plane Strain Fracture 3

Table 1 Chemical Compositions Percent Material C Mn P S Si Ni Cr Mg Nodular Iron 3.1 0,23 0.02 0.019 2.39 0.97 0.04 Gray Iron 3.16 1,0.0.104 0.065 2.59 0.26 Table 2 shows the mechanical properties of interest for these two alloys at room temperature, and at -50 F (-46 C). Table 2 Mechanical Properties Material Room Temperature - 50 F (-46 C) Nodular Iron Yield = 46,300 psi, (319 MPa) Ultimate = 59,000 psi, (406 MPa) Yield = Ultimate = 54,500 psi, (375 MPa) Gray Iron Yield = Ultimate = 35,000 psi, Yield = Ultimate = 37,300 psi, (256 MPa) (2)41 MPa) Toughness data (KI c ) from [Reference (2)3. Since Charpy data is so commonly used as toughness criteria, it is worthwhile to examine the Charpy data and compare it with the available K I data. c Fig. 3, (2), shows Charpy energies obtained at various temperatures for the nodular and gray iron discussed in this paper. As is apparent from this figure Charpy energies at lower temperatures are extremely small, and in the case of gray iron, do not increase appreciable at higher temperatures. For this reason, a conversion of these Charpy values into fracture toughness values is of questionable value. However, in many cases this is the only method available to the design engineer and, therefore, will be investigated here. Fig. 4 shows two relationships between rrpy data and K Ic which were obtained from ferences (8) and (9)3. Both correlations are for steels and must be extrapolated to the lower Charpy values of irons. Fig, 3 shows that at -50 F (-46 C) nodular iron has a Charpy energy of 2 ft lb (2.7 J), while gray iron has approximately l.ft lb (1.4 J). Via Fig. 4 this yields values of KI of 10 ksi in., (11 MPa j '' ), and 5 ksi-in.,?5.5 MPa^ m ), respectively, or very nearly equal to the calculated Stress Intensity Factor for both materials. This would indicate that with the assumed flaw size and stresses, at -50 F (-46 C) either material could be susceptable to brittle fracture. However, fracture toughness data for both materials would indicate otherwise. Fracture toughness tests were performed on both nodular and gray iron using a three point bend bar (Reference (2)1. The tests were conducted at -320 F (-196 C) in order to obtain Plane Strain conditions for the size test bar used. The results show that nodular iron has a KI c of 21.6 ksi in., (23.8 MPa ), while gray iron has a KI c of 24.2 ksi in., (26.6 MPa[). Thus the fracture toughness tests indicate that nodular iron has approximately twice the toughness needed for these applications, and that gray iron has nearly five times the necessary toughness. Fracture toughness tests on 11 grey cast irons (10) resulted in values of K I varying from 11.3 to 18.5 ksi (in.) 1/2 (12.4-20.4 MPa m ), It is most important therefore to evaluate this quantity for the particular grey iron and heat treatment involved, Many materials are strain rate sensitive and toughness values should be obtained that simulate the most rapid loading rate expected, These impact fracture toughness values sometimes using the symbol KId are a more useful index that K I obtained under very slow loading conditions. c Reference (1) contains data of this kind for nodular irons at low temperatures and shows that KId (11) may be considerably lower than K I c CRITICAL FLAW SIZE Another way to examine this situation is 4

J 3 0 KJ N ^ 2 I^ J 1 N ' cd IKI. IROJ 100 80 60 :iiiiiizit7 0.^ REF, 8 li. w 15 0 40 si REF, 9 r0 N 20 N ^ U 10 5-200 -100 0 100 200 NODULAR IRON TEMPERATURE -'C 9' 118 94 2631 300 400 500 600 &-8 13G C _. -tn LlI'arw ENMRGY JDi_ILES 2^4 2 72 _ - -firs CHARPY ENLRGY - FT, LB. Fig, 4 Charpy energy versus Klc 34 I 0 Mfr 4. B TEMPERATURE - `F Fig, 3 Charpy energy versus temperature for the irons in Tables 1 and 2 to determine the flaw size necessary to produce brittle fracture, given the fracture toughness and operating stress. When solved for a c, equation (1) reduces to the following: more than sufficient toughness for the intended applications. For either case, the flaw shapes encountered were relatively long and thin, with Q being close to 1.0. Fig. 1 shows that for a /uys less than 0.5, Q is approximately 1.0 for any long thin flaw (a/2c <0.1), Thus for small values of Q/Qys, it is conservative to assume that Q = 1.0, If Q max in equation (3) is expressed as a fraction of yield strength (i.e., u max = f (a ys ), then for f ' 0.5 equation (3) becomes: _ I ax 2 (3) acn m 1 /K I \2(EL) a c \f a ^) For nodular iron where Q = 1.0, K I = 21,6 ksi ;ft., (23.8 MPA J), and umax = 11 k si (75.1 MPa), equation (3) yields: a = n () 2 = 1.23(31.'3mm) This size flaw would be approximately equivalent to a 2.5 in. x 25 in., (63.5 mm x 635 mm), spongy shrink area, This is a large flaw which could be easily detected during an inspection. For gray iron where Q = 1.18, KI c = 24.2 ksi(, (26.6 MPa J), and max = 4.5 ksi, (31.1 MPa), equation (3) yields: 1.18 24.2 2 a c = n 4.5) = 11.6 11 (29q mm) This is an extremely large flaw and could not occur in a machinery component. The magnitude of these flaws sizes is another way of showing that both materials have Thus the critical flaw size can be expressed in terms of f and the ratio, KI c/sys. It should be remembered that equation (4) applies to an internal flaw and for values of 0/C 0.5, Fig. 5 is a plot of equation (4) for various values of KI c/ays, both of which are material properties. Expressed in this general form, Fig. 5 may be used for a variety of materials (depending upon the ratio of K Ic / ys ), and may be used conservatively for any elliptical flaw provided r/tys is less than approximately 0.5. This is often the case since the majority of brittle fracture problems occur at stress levels considerably below the yield strength. For nodular iron, KI /rys = 21,6/54.5 = 0.40, while for gray iron ki c / ys = 24.2/37.3 = 0. 65. These values are shown in Fig. 5 for flaws sizes up to 2.5 in., (63.5 mm). An excellent reference for the graphical solution of such problems can be found in the paper ( 12 ), H. D., Greenberg, and W. G. Clark "A Fracture Mechanics Approach to the Development of Realistic Standards for Heavy Walled Steel 5

Castings." These solve the previously mentioned analytical expressions graphically and yield rapid results for scoping the brittle fracture susceptibility of a structure. I a DISCUSSION It is interesting to note that the Charpy impact data and other physical properties for both gray and nodular iron do yield fracture toughness values less than the KI c determined in the fracture toughness tests ( 13-15 ). Thus, from interpretation of the Charpy data alone, it would be concluded that brittle fracture was a distinct possibility for both materials for machinery applications. The actual fracture toughness is, however, more than adequate for the applications. One explanation of this discrepency between toughness predictions could be the scatter of the Charpy data, particularly in the lower ranges. An absolute error of 2 ft lb (2.7 J) for a Charpy test is not uncommon. In addition a certain amount of the Charpy energy (approximately 0.5 ft lb, (1,4 J) is transformed into kinetic energy as the specimen fails. When these sources of error are combined for lower range Charpy values a percentage error of 200 to 300 percent can result. Fig. 4 shows that such an uncertainity in Charpy values will produce similar changes in KI c. In fact, the uncertainity is enough to account for the discrepancy in fracture toughness agreement noted here, Another possible explanation of the discrepancy could be the strain rate effect caused by the difference in loading rates between the Charpy tests and the KI tests. The loading rate for the KI c tests was such that all specimens failed in from 1 to 1.5 min, while the loading rate for the Charpy tests is such that failure is nearly instantaneous. Reference ( 1 ) confirms that some nodular irons will give good ductility in relatively slow strain rate tension tests, yet will fail in a brittle manner on impact. The value of examining both K Ic data and Charpy data is apparent. For machinery applications, or other applications where no impact type of loading is anticipated, the use of Charpy data alone as a toughness criteria could lead to an over conservative design, or use of material with unnecessarily excessive toughness. Likewise for applications where impact loading is anticipated, interpretation of only KI c data as toughness criteria could lead to the use of a material with insufficient toughness. 1, 0. DL- 0.1 0,2 2.3 0,1 t = ^ax/jys Fig,5 Generalized critical flaw curves ACKNOWLEDGMENT The authors wish to acknowledge the support of the Gas Turbine Department of the General Electric Co., Schenectady, New York for this work and in particular Mr. D. E. Brandt of the same department for his advice and help on this paper. REFERENCES 1 "Progress in Measuring Fracture Toughness and Using Fracture Mechanics," Materials and Research Standards, March 1964, p. 115. 2 Brandt, D. E., "Crack Growth and Fracture Characterization of Ductile Iron, Gray Iron, and 0.25 Percent Carbon Cast Steel," General Electric Report 71-GTD-17, 1971. 3 Gilbert, G. N. J., "Ductile and Brittle Failure in Ferritic Nodular Irons," Symposium of Metallic Materials at Low Temperatures, ASTM STP-158, 1953, pp. 415-431. 4 Sandoz, G. A., Bishop, H. F., and Pellini, W. S., "Notch Ductile High Strength Nodular Irons," Transactions of the ASM, Vol. 48, 1956, p. 974. 5 Gilbert, G. N. J., "Low Temperature Properties of Cast Irons," Symposium of Metallic Materials at Low Temperatures, ASTM-STP-158, 1953, p. 432-435. 6 Pellini, W. S., Sandoz, G., and Bishop, H. F., "Notch Ductility of Nodular Iron," Transactions of ASM, Vol. 46, 1954, p. 418-450. 7 Mogford, I. L., Brown, J. L., and Hull, D., "Fracture of Nodular Cast Iron," 6

Journal of the Iron and Steel Institute, July 1967, 8 Pellini, W. S., "Advances in Fracture Toughness Characterization Procedures and in Quantitative Interpretations to Fracture Safe Design of Structural Steels," NRL Report 6713, 9 Barsom, J. M., and Rolfe, S. T., "Correlation Between Klc and Charpy V Notch Test Results in Transition Temperature Range," Presented at the ASTM Meeting, June 1969, U. S. Steel Paper, Fig, 14, 10 Grover, A. G., and Pollard, G., "Brittle Failure of Grey Cast Irons," International Conference on Fracture, 2nd,, Proceedings Brighton England, April 1969. 11 Bishop, H. F,, Sandoz, G., Howells and Pellini, "Notch Ductility of Commercial Malleable Iron," NRL Report 4713, March 1956. 12 Greenberg, H, D,, and Clark, W. G., "A Fracture Mechanics Approach to the Development of Realistic Standards for Heavy Walled Steel Casting," Metals Engineering Quat., August. 1969. 13 Pellini, W. S., Sandoz, G. A,, and Bishop, H. F., "Drop Weight Test Measures Notch Ductility," Iron Age, July 8, 1954, 14 Pellini and Loss, "Integration of Metallurgical and Fracture Mechanics Concepts of Transition Temperature Factors Relating to Fracture Safe Design for Structural Steels," NRL Report 6900, April 1969, 15 Vanick, J. S., "Low Temperature Toughness of Flake and Spherodal Graphite Cast Iron Symposium on Metallic Materials at Low Temperature," ASTM-STP-158, 1958, pp. 405-414. 7