MATERIALS FORUM VOLUME 32 - 2008 Edited by J.M. Cairney, S.P. Ringer and R. Wuhrer © Institute of Materials Engineering Australasia Ltd

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THE EFFECTS OF INITIAL MICROSTRUCTURE AND HEAT TREATMENT ON THE CORE MECHANICAL PROPERTIES OF CARBURIZED

AUTOMOTIVE STEELS

E. Boyle1, D.O. Northwood1*, R. Bowers1, X. Sun2, P. Bauerle2

1 Department of Mechanical, Automotive, and Materials Engineering, University of Windsor, 401 Sunset Ave., Windsor, Ontario, Canada N9B 3P4

2 Chassis & Powertrain Materials Engineering,

Chrysler LLC, 800 Chrysler Drive, Auburn Hills, MI, U.S.A. 48326-275

ABSTRACT SAE 8620 and other steels are typically used in the carburized condition for powertrain applications in the automotive industry. Steel parts are generally normalized before carburizing to produce a homogeneous austenite with a uniform grain size. Elimination of the intermediate normalizing treatment could lead to significant cost reductions. The intent of this work was to investigate the effects of composition and initial microstructure (as-received or normalized) on the hardenability as well as the core microstructure and mechanical properties of two case carburized (SAE 8620 and PS-18) steels. Both steels were received in the form of hot-rolled bars, but the PS-18 had received a special controlled-cool processing. This study complements our previous work on the effects of initial microstructure and heat treatment on the residual stress, retained austenite, and distortion of the same carburized automotive steels. The feasibility of using either a new steel or heat treatment process cycle requires acceptable properties in both the case and the core; both directly affect part quality and performance. Jominy curves revealed that the as-received hardenability of PS-18 was higher than that of the 8620 steel. The normalizing heat treatment was characterized for both steels by Jominy testing and by the effect on prior austenite grain size. In both the as-received and normalized heat treatment conditions, the PS-18 steel was noted to have higher ultimate tensile strength and lower Charpy impact toughness than the 8620 steel. The carburizing thermal cycle heat treatment could lead to unacceptable impact toughness values for the core of PS-18 steel components. 1. INTRODUCTION SAE 8620 and other steels are typically used in the carburized condition for powertrain applications in the automotive industry (i.e. differential ring gears, camshafts, and transmission gears). Gas carburizing is used most often for large-scale production, with the two-stage boost-diffuse technique being widely applied in order to produce the desired carbon profile1,2. Steel parts are generally normalized before carburizing to produce a homogeneous austenite with a uniform grain size3,4. Hardenability is essentially the ease of forming martensite and reflects the ability of a steel to be hardened to a specified depth5,6. The factors that increase hardenability are: 1) dissolved elements in austenite (except Co), 2) coarse grains of austenite, and 3) homogeneity of austenite3,6,7. The hardenability of low-alloy heat treatable steels is almost always tested and specified in terms of Jominy (end-quench) hardenability3,5-9. Methods for calculating hardenability have been proposed, from empirical compositional factors based on the work of Grossmann and others, to regression based equations or systems based on more fundamental metallurgical properties8,9.

The main objective of carburizing is to obtain a high-carbon martensitic case with good wear and fatigue resistance combined with a tough, low-carbon steel core1. In terms of core mechanical properties, it is generally the ultimate tensile strength property which is specified10. There is rarely any reference to the core microstructure, but strength properties are structure dependent1,10. The hardenability ultimately determines the microstructure present in the core material. Soft cores (< 750 N/mm2) suggest that the microstructure has a high ferrite content, while hard cores (>1240 N/mm2) suggest a predominantly martensitic structure10. The core material may also have a bainitic structure, or even a mixed microstructure containing ferrite, bainite, and martensite1,10. Because many applications require high core strength to support the case, alloy steels with good core hardenability that form martensite throughout a carburized part are often used1. The Charpy impact test is widely applied in industry to characterize the failure behaviour of metals and alloys, but is largely employed for determining the ductile-to-brittle transition temperature11. The general trend for many steels is that as the tensile strength increases, the impact strength falls10. Also, as the strength and hardenability of a steel increase, the elongation and reduction of area properties decrease10.

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The intent of this work was to investigate the effects of composition and initial microstructure (as-received or normalized) on the hardenability as well as the core microstructure and mechanical properties of two case carburized (SAE 8620 and PS-18) steels. Elimination of the normalizing treatment prior to carburizing could lead to significant cost reductions. This study complements our previous work12,13 on the effects of initial microstructure and heat treatment on the residual stress, retained austenite, and distortion of the same carburized automotive steels. The feasibility of using either a new steel or heat treatment process cycle requires acceptable properties in both the case and the core; both directly affect part quality and performance. 2. EXPERIMENTAL DETAILS The steels were received in the form of hot-rolled bars; their compositions are given in Table 1. The hot-rolled bars are considered the as-received condition. The SAE 8620 steel was air cooled from the hot working temperature. The PS-18 material received an unspecified controlled-cooling process from the hot working temperature. The normalized condition involved heat treating the as-received materials at 924°C for 2 h, followed by air cooling. Table 1. Chemical composition of SAE 8620 and PS-18

steel (wt. %)

In a previous study12, Navy C-ring specimens were machined from SAE 8620 and PS-18 steels in both the as-received (A) and normalized (N) conditions. The C-ring specimens were then carburized using a boost-diffuse process. The specimens were held at 927°C for 9.75 hours with a carbon potential of 1.05 and at 843°C for 1.5 hours with a carbon potential of 0.88, followed by oil quenching at 149°C. In order to simulate the core material of the carburized parts, test specimens for the present study were subjected to a heat treatment which consisted of a thermal cycle similar to that of the carburizing process, but utilized a much lower carbon potential. This carbon potential (0.3) was sufficient to prevent decarburization of the specimens during heat treatment. The actual carburizing thermal cycle heat treatment consisted of 927°C for 1.5 hours at a carbon potential of 0.3, 843°C for 1.5 hours at a carbon potential of 0.3, followed by oil quenching at 149°C. Test specimens are identified using three characters relating to

the 1) material (8 for 8620, P for PS-18), 2) material condition (A or N), and 3) sample number (1, 2, 3). The end-quench test for hardenability of 8620 and PS-18 steel was performed in accordance with ASTM Standard A255-02: Standard Test Methods for Determining Hardenability of Steel14. End-quench (Jominy) Preferred Test Specimens were machined from the hot-rolled bars in both the as-received and normalized conditions. Two specimens were tested for each material condition. The only exception to the standard was that the as-received materials were not normalized prior to specimen preparation; the objective of this test was to compare the hardenability of the as-received materials to the normalized condition. For hardness measurements, two flats were ground 180° apart to a depth of 0.38 mm along the entire length of the bar. The calculation of hardenability for 8620 and PS-18 steels was also performed in accordance with ASTM Standard A255-02. The equations provided in the standard assume a No. 7 austenitic grain size. Metallographic analysis was conducted on four of the Jominy specimens, 8A1, 8N1, PA2, and PN2. The samples were etched using a picric acid solution to reveal the prior austenite grain boundaries. The etchant consisted of 1.5 g of picric acid dissolved in 150 ml of distilled water; to this solution 3 ml of soap detergent was added. The entire solution was then heated to a temperature range of 68-75°C. The sample was immersed in the solution in 15 s intervals, with light swabbing every 2-3 s to remove the dark film which formed on the surface. The sample was then rinsed with distilled water, followed by rinsing with ethanol and dried. This etching process continued until the prior austenite grain boundaries were sufficiently delineated. Prior austenite grain size measurements were performed using the Intercept Procedure in accordance with ASTM Standard E112-96(2004): Standard Test Methods for Determining Average Grain Size15. Tensile testing was conducted in accordance with ASTM Standard E8-04: Standard Test Methods for Tension Testing of Metallic Materials16. Tensile specimens were machined from 8620 and PS-18 steels in both the as-received and normalized conditions. The Standard 0.500-in. Round Tension Test Specimen with 2-in. Gage Length geometry was used for all test specimens. Some of these specimens were then subjected to the carburizing thermal cycle heat treatment. Two specimens were tested for each material condition. Charpy impact testing was conducted in accordance with ASTM Standard E23-05: Standard Test Methods for Notched Bar Impact Testing of Metallic Materials17. Charpy impact specimens were machined from 8620 and PS-18 steels in both the as-received and normalized conditions. The specimens were prepared along the longitudinal direction, and electron discharge machining (EDM) ensured precise machining of the notch. The Charpy (Simple-Beam) Impact Test Specimen – Type A geometry was used for all test specimens. Some of these specimens were then subjected to the carburizing thermal

Element 8620 PS-18C 0.22 0.27

Mn 0.84 1.00Cr 0.51 0.57Mo 0.19 0.15Ni 0.57 0.12P 0.009 0.018Si 0.23 0.26S 0.021 0.026

Cu 0.20 0.22Al 0.026 0.030

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cycle heat treatment. Impact testing was performed on three specimens for every material condition at -40, -20, 0, 23, and 40°C. For specimens with test temperatures below room temperature (23°C), liquid bath thermal conditioning was employed using a dry ice and alcohol solution. For specimens with test temperatures above room temperature, furnace thermal conditioning was employed. Specimens were placed in the bath or furnace for 20 min. and removed immediately before testing. After testing, the fracture surface of each specimen was examined, and the proportion of shear fracture was determined in accordance with method A6.1.217 of the standard. This method involved comparing the appearance of the fracture surface with a fracture appearance chart. Two independent measurements were made for each specimen, and these values were averaged for analysis. 3. RESULTS AND DISCUSSION 3.1 Microstructures For both 8620 and PS-18, the effectiveness of the normalizing process in refining and homogenizing the grain structure is shown in Figures 1 to 4. Figures 1 and 2 show the as-received and normalized microstructure of 8620, respectively. Figures 3 and 4 show the as-received and normalized microstructure of PS-18, respectively. Both the 8620 and PS-18 as-received microstructures exhibit proeutectoid and Widmanstätten ferrite. Optical metallography could not resolve the pearlite layers. For the 8620 material, the Widmanstätten ferrite constituent persists in the normalized condition; it is not noted in the normalized PS-18 microstructure. Scanning electron microscopy (SEM) was required to resolve the pearlite layers. Figures 5 and 6 show the as-received and normalized microstructure of 8620, respectively. Figures 7 and 8 show the as-received and normalized microstructure of PS-18, respectively. Fine pearlite layers are observed in all of the microstructures. In some samples, the orientation of the plates appears somewhat irregular in some grains.

Figure 1. Optical micrograph of as-received SAE 8620, showing proeutectoid and Widmanstätten ferrite and

pearlite. Nital etch.

Figure 2. Optical micrograph of normalized SAE 8620,

showing proeutectoid and Widmanstätten ferrite and pearlite. Nital etch.

Figure 3. Optical micrograph of as-received PS-18, showing proeutectoid and Widmanstätten ferrite and

pearlite. Nital etch.

Figure 4. Optical micrograph of normalized PS-18, showing proeutectoid ferrite and pearlite. Nital etch.

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Figure 5. SEM micrograph of as-received SAE 8620, showing proeutectoid and Widmanstätten ferrite and

pearlite. Nital etch.

Figure 6. SEM micrograph of normalized SAE 8620, showing proeutectoid and Widmanstätten ferrite and

pearlite. Nital etch. Figure 7. SEM micrograph of as-received PS-18, showing proeutectoid and Widmanstätten ferrite and pearlite. Nital etch. Figures 9 to 12 show the microstructure of the steels after the carburizing thermal cycle heat treatment. Figures 9 and 10 show the carburizing thermal cycle heat treated

as-received and normalized microstructure of 8620, respectively. Figures 11 and 12 show the carburizing thermal cycle heat treated as-received and normalized microstructure of PS-18, respectively.

Figure 8. SEM micrograph of normalized PS-18, showing proeutectoid ferrite and pearlite. Nital etch.

Figure 9. Optical micrograph of as-received 8620 after the carburizing thermal cycle heat treatment, showing

martensitic structure. Nital etch.

Figure 10. Optical micrograph of normalized 8620 after

the carburizing thermal cycle heat treatment, showing martensitic structure. Nital etch.

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Figure 11. Optical micrograph of as-received PS-18 after

the carburizing thermal cycle heat treatment, showing martensitic structure. Nital etch.

Figure 12. Optical micrograph of normalized PS-18 after

the carburizing thermal cycle heat treatment, showing martensitic structure. Nital etch.

A martensitic structure is observed for the as-received and normalized condition of both steels. While there remains a difference between the as-received and normalized microstructures, it appears that the magnitude of this difference is minimized by the carburizing thermal cycle heat treatment. 3.2 Hardenability Table 2 summarizes the hardness data for the end-quench test specimens. Figure 13 shows the end-quench hardenability curves for the 8620 steel. It plots the average value of the two as-received specimens, the average value of the two normalized specimens, and the calculated hardenability based on the ASTM standard. Figure 14 shows the equivalent information for the PS-18 steel. The hardenability data was fairly reproducible for each material condition, as evidenced by hardness values for equivalent specimens which differed by 2 HRC or less. The PS-18 specimens had a higher hardenability than the 8620 specimens for both the as-received and normalized conditions. For PS-18, the normalized

specimens were noted to have a slightly higher hardenability than the as-received specimens, whereas the opposite was observed for the 8620 specimens. However, the hardness values between the conditions varied by 5 HRC or less for PS-18 and by 3 HRC or less for 8620. There was generally a good agreement between the end-quench test data and the calculated data for hardenability, with the possible exception of the as-received PS-18 specimens between 4.8 mm and 14.3 mm from the quenched end. Optical micrographs of specimens 8A1, 8N1, PA2, and PN2 etched to reveal the prior austenite grain boundaries are shown in Figures 15 to 18, respectively. The prior austenite grain size data for these four specimens is given in Table 3. It should be noted that the hardenability equations from the standard are based on the assumption of a No. 7 grain size, while the measured grain sizes of the specimens ranged from No. 8 to 9.5; correlation of the calculated and experimental values showed little effect of this difference. Normalizing refined the grain size for both steels, but this effect was more evident for PS-18. 3.3 Tensile Properties Figure 19 shows the ultimate tensile strength (UTS), yield strength (YS), percent elongation (%EL), and percent reduction in area (%RA) for 8620 and PS-18 steel in the as-received and normalized conditions. Figure 20 shows the same data for the equivalent specimens after the carburizing thermal cycle heat treatment. PS-18 had a higher ultimate tensile strength than 8620. For 8620 and PS-18 specimens not subjected to the carburizing thermal cycle heat treatment, normalizing consistently decreased the ultimate tensile strength and increased the percent elongation, or ductility; there was relatively little change in the yield strength. Normalizing had little effect on the reduction in area values for 8620, but slightly increased this property for PS-18. For 8620 and PS-18 carburizing thermal cycle heat treated specimens, normalizing slightly increased the ultimate tensile and yield strength; the ductility and reduction in area decreased for 8620, but increased for PS-18. The carburizing thermal cycle heat treatment led to an increase in ultimate tensile and yield strength and a decrease in elongation and reduction in area for both steels in either the as-received or normalized condition. These trends agree with those documented by Parrish10. 3.4 Impact Properties Figures 21 and 22 show the Charpy impact test data for as-received and normalized 8620 and PS-18 specimens before and after the carburizing thermal cycle heat treatment, respectively. The impact energy for each specimen, as well as an average impact energy for each material-temperature combination, is shown. The error bars shown reflect the standard deviation of these energy readings. PS-18 had a lower impact toughness than 8620. For the specimens not subjected to the carburizing thermal cycle heat treatment, normalizing consistently increased the impact toughness of both steels.

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Table 2. Hardness Data for End-Quench Specimens

Figure 13. Hardenability curves for SAE 8620

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8620 A Avg. 8620 N Avg. 8620 Calc.

8A1 8A2 8N1 8N2 PA1 PA2 PN1 PN21.6 46 46 45 44 49 49 49 493.2 44 45 43 42 48 48 48 484.8 40 40 41 39 45 45 47 476.4 35 35 35 34 39 39 43 447.9 31 32 31 29 35 35 39 399.5 29 29 28 27 32 32 35 36

11.1 27 27 26 26 30 30 33 3312.7 26 26 25 24 29 28 31 3114.3 25 25 24 23 27 28 29 3015.9 23 24 23 23 27 27 28 2817.5 23 24 23 22 26 26 27 2719.1 22 23 22 22 25 25 27 2820.6 22 22 22 21 25 25 26 2722.2 21 22 21 20 24 24 26 2623.8 - 21 - - 24 24 25 2625.4 - 21 - - 23 24 25 2528.6 - - - - 23 22 24 2531.8 - - - - 23 22 23 2334.9 - - - - 22 21 23 2338.1 - - - - 21 21 22 2244.5 - - - - - - 21 2150.8 - - - - - - 21 -

Distance from Quenched End

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Figure 14. Hardenability curves for PS-18

Figure 15. Optical micrograph of prior austenite grain

boundaries in specimen 8A1. ASTM grain size no. 8.5-9.0. 1% Picric acid + surfactant.

Figure 16. Optical micrograph of prior austenite grain

boundaries in specimen 8N1. ASTM grain size no. 9.0-9.5. 1% Picric acid + surfactant.

Figure 17. Optical micrograph of prior austenite grain

boundaries in specimen PA2. ASTM grain size no. 8.0-8.5. 1% Picric acid + surfactant.

Figure 18. Optical micrograph of prior austenite grain

boundaries in specimen PN2. ASTM grain size no. 9.0-9.5. 1% Picric acid + surfactant.

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Table 3. Prior Austenite Grain Size Data

Figure 19. Tensile test data for non-carburizing thermal

cycle heat treated specimens

Figure 20. Tensile test data for carburizing thermal cycle

heat treated specimens

For carburizing thermal cycle heat treated specimens, normalizing had less of an effect on the impact toughness; in both materials, the impact toughness increased at certain test temperatures and decreased at others. The carburizing thermal cycle heat treatment led to a decrease in impact toughness for both steels in either the as-received or normalized condition. This decrease in impact toughness corresponds with an increase in tensile strength, as mentioned by Parrish10. Also, the difference between the minimum energy (i.e. at -40°C) and the maximum energy (i.e. at 40°C) for each material condition is reduced by the carburizing thermal cycle heat treatment. The carburizing thermal cycle heat treatment could lead to unacceptable impact toughness values for the core of PS-18 steel components. In Figures 21 and 22, there is no distinct upper or lower shelf. From this data, the ductile-to-brittle transition temperature (DBTT) cannot be accurately determined. It

appears that the DBTT may, indeed, occur over a range of temperatures. For the materials not subjected to the carburizing thermal cycle heat treatment, this transition region is most likely within the -20°C to 20°C range. Testing of additional specimens at a larger number of temperatures would likely help to determine the DBTT. A separate set of Charpy impact specimens was machined and carburizing thermal cycle heat treated apart from the other heat treated (carburizing thermal cycle) samples. The impact energy of these specimens, which were tested at 0°C, varied from 4.1 to 7.1 J. These data are not included in the plot of carburizing thermal cycle heat treated impact properties. Further investigation is required to understand the discrepant nature of these data. Figures 23 and 24 show the percentage of shear fracture measurements for as-received and normalized 8620 and PS-18 specimens before and after the carburizing thermal cycle heat treatment, respectively. The fracture-appearance methods utilized are based on the concept that 100% shear (ductile) fracture occurs above the transition-temperature range and cleavage (brittle) fracture occurs below the range17. In both figures, the measurement for each sample, as well as an average value with standard deviation error bars, is shown. As previously noted, the data from the carburizing thermal cycle heat treated specimens tested at 0°C is not included. For specimens not subjected to the carburizing thermal cycle heat treatment, normalizing increased the percentage of shear fracture for both steels. For 8620 and PS-18 steels, there was not much variation in the percentage of shear fracture observed in either the as-received or normalized specimens at test temperatures below 0°C for the normalized specimens and 20°C for the as-received specimens. At the higher test temperatures, 8620 had a higher percentage of shear fracture than PS-18. For the carburizing thermal cycle heat treated specimens, normalizing had little effect on the percentage of shear fracture for both steels. 8620 consistently had a higher percentage of shear fracture than PS-18 at all test temperatures, but, similar to the non-carburizing thermal cycle heat treated specimens, this difference was greater at higher test temperatures. With respect to the percentage of shear fracture, the carburizing thermal cycle heat treatment led to a slight increase in the values for the as-received specimens and a significant decrease in the values for the normalized specimens; this decrease was more pronounced at the higher test temperatures. From the appearance of the curves in Figures 21 to 24, there appears to be some correspondence between the impact energy and percentage of shear fracture. In general, higher impact energies corresponded to a higher percentage of shear fracture. 3.5 Case and Core Properties In a previous study12, noticeable differences in the as-received and normalized microstructures for both steels did not result in significant variation in size

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8A1 0.0148 0.0167 8.5 - 9.08N1 0.0139 0.0156 9.0 - 9.5PA2 0.0179 0.0201 8.0 - 8.5PN2 0.0126 0.0141 9.0 - 9.5

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distortion after carburizing. The higher scatter in shape distortion data made it difficult to determine the effect, if any, of normalizing. The small sample size of the study limited the ability to quantify the effect of composition and initial microstructure on the amount of surface retained austenite and residual stress. From analysis conducted on the core mechanical properties, the effect of normalizing is more easily quantified. Normalizing of the non-carburizing thermal cycle heat treated materials lead to a slight decrease in tensile strength, a slight increase in ductility, and a

significant increase in impact properties. Following the carburizing thermal cycle heat treatment, the normalized specimens had a slightly higher tensile strength than the as-received specimens for both steels. The normalized 8620 specimens had a slightly lower percent elongation and reduction in area, while the opposite was observed for PS-18. Normalizing had less of an effect on the impact toughness and percentage of shear fracture. It should be noted that, with respect to the magnitude of these properties, the difference between the normalized and as-received specimens was relatively small.

Figure 21. Charpy impact test data for non-carburizing thermal cycle heat treated specimens

Figure 22. Charpy impact test data for carburizing thermal cycle heat treated specimens

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Figure 23. Percentage of shear fracture data for non-carburizing thermal cycle heat treated specimens

Figure 24. Percentage of shear fracture data for carburizing thermal cycle heat treated specimens

From data obtained on the carburized parts, as well as case and carburizing thermal cycle heat treated core properties, normalizing did not have a significant impact on any of the parameters investigated. Since the required material properties are largely dependant on the application, it is recommended that additional testing of actual parts be conducted under conditions that are more representative of those experienced in service. As such, the performance of the part as a whole could be evaluated. The feasibility of using either a new steel or heat treatment process cycle requires acceptable properties in

both the case and the core; both directly affect part quality and performance. 4. CONCLUSIONS The normalizing process was effective in refining and homogenizing the grain structure for both 8620 and PS-18; this refining effect was more evident for PS-18. For 8620, the ASTM grain size number increased from 8.5-9.0 to 9.0-9.5, while for PS-18, it increased from 8.0-8.5 to 9.0-9.5. Normalizing reduced the amount of

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Widmanstätten ferrite in the microstructure of both steels. Normalizing increased the hardenability of PS-18, as noted by a 1-5 HRC increase across the Jominy specimen length. Normalizing resulted in a 1-3 HRC decrease across the Jominy specimen length for 8620. PS-18 had a higher hardenability than 8620. There was generally a good agreement between the end-quench test data and the calculated data for hardenability. For the non-carburizing thermal cycle heat treated materials, normalizing consistently lowered the tensile strength and increased the ductility, impact toughness, and percentage of shear fracture of both steels. Normalizing had little effect on the reduction in area property for 8620, and only slightly increased this property for PS-18. For the carburizing thermal cycle heat treated materials, normalizing slightly increased the ultimate tensile and yield strength; the ductility and reduction in area decreased for 8620, but increased for PS-18. Normalizing had a minimal effect on the impact toughness and percentage of shear fracture in both materials. PS-18 exhibited a higher tensile strength and lower impact toughness than 8620. The carburizing thermal cycle heat treatment led to an increase in tensile strength and a decrease in elongation, reduction in area, and impact toughness for both steels in both the as-received or normalized condition. The carburizing thermal cycle heat treatment could lead to unacceptable impact toughness values for the core of PS-18 steel components. With respect to the percentage of shear fracture, this heat treatment slightly increased the values for the as-received materials and significantly decreased the values for the normalized materials. The effect of normalizing that was noted for the non-carburizing thermal cycle heat treated samples was negated by the carburizing thermal cycle heat treatment process. Since the required material properties are largely dependent on the application, it is recommended that additional testing of actual parts be conducted under conditions that are more representative of those experienced in service. Acknowledgments This work was supported by the Chassis & Powertrain Materials Engineering Department of Chrysler LLC and by the Natural Science and Engineering Research Council of Canada through Discovery Grants awarded to Professors Derek O. Northwood and Randy Bowers. Erin Boyle acknowledges financial support through the award of Ontario Graduate Scholarships. References 1. G. Krauss: Steels: Processing, Structure, and

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13. D.O. Northwood, L. He, E. Boyle, and R. Bowers: Materials Science Forum, 2007, vol. 539-543, pp. 4464-4469.

14. ASTM International, Annual Book of ASTM Standards - 2006, ASTM Standard A255-02, vol. 01.05, V.A. Mayer et al., ed., ASTM International, West Conshohocken, PA, 2006, pp. 30-53.

15. ASTM International, Annual Book of ASTM Standards - 2006, ASTM Standard E112-96(2004), vol. 03.01, V.A. Mayer et al., ed., ASTM International, West Conshohocken, PA, 2006, pp. 273-298.

16. ASTM International, Annual Book of ASTM Standards - 2006, ASTM Standard E8-04, vol. 03.01, V.A. Mayer et al., ed., ASTM International, West Conshohocken, PA, 2006, pp. 64-87.

17. ASTM International, Annual Book of ASTM Standards - 2006, ASTM Standard E23-05, vol. 03.01, V.A. Mayer et al., ed., ASTM International, West Conshohocken, PA, 2006, pp. 160-186.

Contact Prof. Derek O. Northwood; Tel: (519) 253-3000 ext.4785; Fax: (519) 973-7007; E-mail: dnorthwo@uwindsor.ca