COMPARISON OF THE TRIBOLOGICAL PROPERTIES OF DIFFERENT COLD WORK STEELS AT TEMPERATURES UP TO 250 C

Transcription

1 COMPARISON OF THE TRIBOLOGICAL PROPERTIES OF DIFFERENT COLD WORK STEELS AT TEMPERATURES UP TO 250 C G. A. Fontalvo Materials Center Leoben Franz-Josef Strasse 13 A-8700 Leoben Austria C. Mitterer Department of Physical Metallurgy and Materials Testing University of Leoben Franz-Josef Strasse 18 A-8700 Leoben Austria Abstract Tools for cold forming are subjected to mechanical and thermal loads, resulting in abrasion, adhesion, tribooxidation, and surface fatigue, i.e. wear mechanisms which determine tool life. The manufacture of austenitic stainless steel, for example, is especially difficult due to the strong tendency of that steel to work harden and to adhere to the tool material. In the present work the tribological behavior of five different cold work steels at temperatures ranging up to 250 Cwas investigated. Cold work steel discs were worn against austenitic stainless steel balls in ambient air using a high-temperature tribometer. Wear tracks on the discs and balls were characterized using scanning electron microscopy (SEM), electron probe microanalysis (EPMA) and an optical profiler. This combination allowed for the determination of the wear mechanisms at low and high temperatures. At room temperature, the adhesion of austenitic slider material to the cold work steels dominates the wear behavior, plowing the austenitic slider and protecting the discs from wear. The increasing plowing action of loose wear debris and strength loss 223

2 224 6TH INTERNATIONAL TOOLING CONFERENCE cause an increase in the wear of the cold work steels at elevated temperatures. Keywords: cold work steel, friction, wear, high temperature, austenitic stainless steel, carbide content INTRODUCTION Tools for cold forming are subjected to mechanical and thermal loads, resulting in abrasion, adhesion, tribooxidation and surface fatigue. The progress of these wear mechanisms determines tool life. The interactions that take place at the interface between two or more bodies in relative motion control the tribological behavior of the materials involved [1]. As interface interactions determine reliability as well as tool lifetime in any machining process, it is very important to maintain adequate control of friction and wear during sliding contact. A fundamental knowledge base of the mechanisms of tribological interactions is necessary to establish that control. Friction is determined by the contributions of asperity deformation, adhesion and plowing by hard asperities and wear debris, whereas the main wear mechanisms are adhesion, abrasion, surface fatigue and tribo-oxidation. For a particular application, these mechanisms contribute to a certain extent to the tribological behavior, giving rise to the fact that friction and wear are system-dependent properties. Although many theories have been developed through the years to explain the phenomenology of friction and wear [1, 2, 3, 4], none allows the prediction of the tribological behavior of materials in any given situation. The manufacturing of austenitic stainless steel is especially difficult due to the strong tendency of these steels to work harden and to adhere to tool material [5]. In many cases, even a minor change in the tool shape due to wear debases the surface quality of the workpiece. Because of the complexity of tribological systems, simple laboratory tests are commonly used to study tribology. They often allow information suitable for the selection and optimization of material combinations to be obtained. The present work was carried out to compare the dry sliding behavior of different cold work steels at their working temperature and to relate their tribological properties to their microstructure. For this purpose, different cold work steel discs were worn against austenitic stainless steel balls using a high-temperature tribometer.

3 Comparison of the Tribological Properties of Different Cold Work Steels at EXPERIMENTAL Steel discs of 40 mm diameter and 10 mm thickness were made of five different cold work steels which are commercially available. The samples were first quenched and tempered to a hardness of 60 HRC, then mirror-polished to an average roughness of 5 nm and ultrasonically cleaned in ethanol. Dry sliding tests were carried out in ambient air at three different temperatures (25 C, 150 Cand 250 C) using a CSEM high-temperature tribometer. The sliding speed and the normal load were kept constant at 0.1 m/s and 2 N, respectively. Austenitic stainless steel balls (DIN ) of 6 mm diameter and with a hardness of 274 HV1 were used as counterbodies. The number of laps performed and the wear track radius was 50 and 10 mm, respectively, resulting in a sliding distance of 3.14 m. The wear tracks of the samples were characterized using scanning electron microscopy (SEM; Cambridge Instruments Stereoscan 360) and white light interferometry (Wyko NT1000). The latter was also used to determine the wear volume of the discs and the amount of material transferred to the discs after ball-on-disc testing. Wear debris were stuck to a copper-strip and characterized using energy-dispersive electron probe microanalysis (EPMA; Link exl). RESULTS AND DISCUSSION MICROSTRUCTURAL CHARACTERIZATION In this paper, we will refer to the five cold work steels investigated as steel 1 to steel 5, as indicated by the order of their optical micrographs shown in Fig. 1. One sample of each steel was etched using a 5% HNO 3 solution, so that the carbides appear white and the matrix dark under the microscope. The main differences among the steels are the carbide contents and sizes. The first three steels Figs. 1a, 1b and 1c, contain the same alloying elements but they have different carbide contents. Unlike the other steels, steel 3 is prepared by powder metallurgy. It shows the highest carbide content and the lowest carbide size of the steels tested. Steels 4 and 5, Fig. 1d and 1e, respectively, are similar to steel 2 from the point of view of microstructure, but they have been developed to be more ductile than steel 2. Thus, their alloy content is slightly different and they show a lower carbide content.

4 226 6TH INTERNATIONAL TOOLING CONFERENCE (a) steel 1 (b) steel 2 (c) steel 3 (d) steel 4 (e) steel 5 Figure 1. Optical micrographs of the five cold work steels investigated.

5 Comparison of the Tribological Properties of Different Cold Work Steels at FRICTION BEHAVIOR The measured mean friction coefficients are shown in Fig. 2. It can be seen that the steels 2, 3, 4 and 5 have friction coefficients of about 0.65 at room temperature and about 0.9 at 150 and 250 C, respectively, although slight differences in the friction coefficient between these steels can be observed, particularly at the highest temperature. In contrast, steel 1 shows a high friction coefficient at room temperature which decreases as temperature rises. During ball-on-disc testing, the friction coefficient is expected to have three dominating components, each of them caused by one of the primary friction mechanisms adhesion, asperity plowing and debris plowing. The microstructure of the cold work steels and its influence on the components of the friction coefficient should provide an explanation of their frictional behavior. Although all steels have a martensitic matrix, there is a significant difference in the volume fraction of carbides between them. Poggie and Wert observed a decrease of friction coefficient and wear rate of tool steels with increasing carbide volume fraction [6]. At room temperature, a reduction in the adhesive interaction between the austenitic steel slider and the cold work steel discs could be expected as a consequence of the increasing carbide volume fraction. Since steel 1 has the lowest carbide volume fraction of all samples investigated, it can be assumed that the adhesive component of the friction coefficient of steel 1 cannot be reduced in the same extent as for the others steels, compare Fig. 2. At higher temperatures, the plowing components presumably play a major role. Here, the carbide volume fraction should again affect the friction behavior as it determines the resistance against plowing. Thus, a higher carbide volume fraction results in higher friction coefficients. This premise is supported by the results of the ballon-disc tests at 250 C, see Fig. 2, where steel 3 shows the highest friction coefficient and steel 1 the lowest. WEAR BEHAVIOR Fig. 3a shows the surface profile of the wear track of a steel 1 disc after ball-on-disc testing. The way in which the surface roughness dramatically changes within the wear track can be clearly seen; there is transferred material and evidence of plowed grooves. As temperature increases, these plowed grooves become wider and there is almost no evidence of transferred material, as Fig. 3c clearly shows.

6 228 6TH INTERNATIONAL TOOLING CONFERENCE Figure 2. Mean friction coefficient for the five cold work steels investigated against austenitic steel for a sliding distance of 50 laps at different temperatures. The software of the optical profiler can be used to estimate the volume occupied by the space between the actual surface profile and a plane representing the ideal surface. A zero level is established and interpreted as cross-horizontal section through the roughness profile at height zero, which intercepts both sample material and air, see Figs. 3a and 3c. The negative volume is defined as the volume above the actual sample surface and below the zero level, or approximately the amount of material been worn away i.e., the wear volume, when the amount of material transferred to those areas below zero level is neglected. The positive volume is defined as the volume below the actual sample surface and above the zero level, or the amount of material transferred. Although the surface peaks and valleys outside the wear track are also account for in the volume calculation, their influence on the results is small as previous measurements indicated. Furthermore, the effect of surface features outside of the wear track was almost constant for all samples investigated since surface roughness does not vary significantly among the steels. Fig. 4a shows the results of calculating the amount of material transferred onto the discs following the method explained above, Fig. 4b shows the

7 Comparison of the Tribological Properties of Different Cold Work Steels at (a) two-dimensional profile after testing at room temperature (b) three-dimensional image of the same wear track as in a) (c) two-dimensional profile after testing at 250 C Figure 3. Surface of the wear tracks on steel 1 after ball-on-disc testing against austenitic stainless steel balls. wear volumes. It can generally be concluded that, as temperature rises, the amount of transferred material decreases and wear increases. Once again, the behavior of steel 1 is different from that of the other steels. At 150 C, the amount of material transferred onto the discs of steels 2 to 5 is nearly as high as at room temperature. In contrast, at elevated temperatures the material transferred onto steel 1 is low and almost the same at 150 and 250 C. Steel 3 shows the lowest amount of transferred material at room temperature and the lowest wear volume at any temperature compared to the other steels. The differences in adhesive wear behavior at room temperature and at 250 Ccan clearly be seen in the SEM micrographs shown in Fig. 5. At room temperature, large numbers of lumps of austenitic slider material can be found on the wear track, see Figs. 5a and 5c, as indicated by SEM and EPMA. However, there is also evidence of plowed microgrooves. At 250 C, almost no material transfer takes place, but the plowed grooves are more evident, see Fig. 5b. It can also be seen from Fig. 4c that the disc wear at 150 Cis about the same as at room temperature. Furthermore, no loose

8 230 6TH INTERNATIONAL TOOLING CONFERENCE (a) Sample A (b) Sample B Figure 4. a) Amount of material transferred from the austenitic stainless steel to the cold work steel discs and b) disc wear for 50 laps sliding distance.

9 Comparison of the Tribological Properties of Different Cold Work Steels at (a) room temperature (b) detail of Fig 5a (c) at 250 C Figure 5. SEM micrographs of wear tracks obtained by ball-on-disc testing of austenitic stainless steel balls on steel 1 discs. wear debris was visible to the naked eye at room temperature. At 150 C, loose wear particles could be found only for steels 1 and 4. Unfortunately, the material recovered was not sufficient for proper analysis. Fig. 6a, where a SEM micrograph of the wear scar of the austenitic steel slider after testing at room temperature is shown, confirms this observation. Virtually no wear debris or lumps can be seen but plowed grooves are visible. On the other hand, at 250 Cwear debris accumulate on the side of the wear scar which comes at first in contact with the steel disc during rotation, see Fig. 6c. Figure 7 shows a SEM micrograph of the accumulated wear debris on the ball after it has been stuck to a copper-strip. Some big wear particles can be seen but the vast majority of wear particles are small with a size of about 2 to 5 'µm. The results of EPMA of the largest wear particle seen in Fig. 7 are shown in Fig. 8a. The Ni peak demonstrates that the large particle is mainly composed of ball material as nickel is not an alloy element from any

10 232 6TH INTERNATIONAL TOOLING CONFERENCE (a) room temperature (b) at 250 C Figure SEM micrographs of the austenitic stainless steel balls after testing against steel of the steels investigated. The fine wear particles, however, are primarily composed of disc material as shown by EPMA, see Fig. 8b, where the Ni peak is not as pronounced as in Fig. 8a. In the tribological situation being investigated two different wear mechanisms are present pre- dominantly, i.e. adhesion and plowing. Adhesion of material from the austenitic steel slider to the rotating disc primarily occurs at room temperature. As temperature rises, ball and disc wear due to the plowing action by wear debris increasingly determines wear. The microstructure of the steels should play a very important role in both mechanisms. As previously mentioned, it is commonly accepted that a high carbide content should reduce the adhesion of counterbody material when sliding against a steel surface provided that the counterbody material is softer than the steel [6]. The reason for that behavior is that a metal-metal contact is necessary for material transfer (where the counterbody is a metal). This explains why steel 3 shows the lowest amount of transferred material at room temperature. The fact that both steels 2 and 5 show a relatively high amount of transferred material in comparison to steels 1 and 4 which have a lower carbide content, leads to the conclusion, that the carbide content, but probably also the carbide size and distribution, could affect the resistance against adhesive wear. As suggested by many previous investigations [7, 8], the formation of discrete

11 Comparison of the Tribological Properties of Different Cold Work Steels at Figure 7. SEM micrograph of the wear debris collected after ball-on-disc testing of austenitic stainless steel balls against steel 5 at 250 C. (a) (b) Figure 8. (a) EDX-analyses of the big flake-like wear particle in Fig.7, (b) EDX- Analysis of the little wear particles in Fig. 7.

12 234 6TH INTERNATIONAL TOOLING CONFERENCE particles during wear occurs by a surface fatigue process. It is conceivable that the contact of the soft slider to the hard carbides could lead to higher slider fatigue on a microscopic scale, and thus relieve the adhesion of the fatigued volume as soon as it comes into metal-metal contact. It can easily be imagined that the area of contact affected by this mechanism depends on the carbide content and carbide distribution as both determine the probability of the prior metal-carbide contact and the posterior metal-metal contact, and on the carbide size. Although steel 3 has the highest carbide content, and thus the highest probability of carbide-metal contact compared to the other steels, it also has the lowest probability of metal-metal contact necessary for adhesion of wear particles. In other words, the resulting adhesive behavior is a superposition of two mechanisms: On the one hand, there is a reduction of metal-metal contact due to a higher carbide content and a corresponding reduction in the transfer of material, while on the other hand, an increase in the micro-scale-fatigue of the austenitic steel slider results in an increase of the production of wear particles which adhere to the discs. Heating the samples prior to ball-on-disc testing takes about two hours, i.e. a sufficient time for a thicker than the natural oxide layer to be formed on the samples, particularly at 250 C. The wear reducing effects of oxidation have been recognized for years [9], and it is thus plausible that the oxide layer reduces the adhesion of ball material on the discs as Fig. 4a shows. Many investigators have pointed out that there is a minimum wear particle size necessary for material to leave a surface in the form of a loose wear particle [?, 11]. The wear particles generated during sliding are retained within the wear scars and grow by agglomeration until they reach a critical size. The agglomeration of particles increases as temperature rises. Nevertheless, the wear particles can remain adhered to a surface due to adhesion forces between solid surfaces and, by further agglomeration, develop compact layers which can become load-bearing [9]. If the wear particles are removed from the surface, there is more wear damage due to the plowing action of the hard particles. In our experiments at room temperature, the transferred particles mainly stuck to the steel disc and plowed the ball, but evidently some of the wear particles were removed and plowed the disc as well. The reason for the high hardness of the wear particles from the austenitic ball, which has a lower hardness than the cold work steels, lies in the fact that modest amounts of plastic strain can cause a rapid martensitic transformation, making the particles hard enough to abrade the tool steels. Grain refinement and

13 Comparison of the Tribological Properties of Different Cold Work Steels at work hardening could also play a role in increasing the hardness of wear particles [12]. At higher temperatures it appears that the removal of wear particles becomes easier compared to room temperature. First, wear particles are removed from the surface, then they agglomerate on one side of the ball wear scar and then they plow the disc. As a consequence of wear particle removal and strength loss at higher temperatures, wear at 250 Cis 3 to 6 times higher than at room temperature, compare Fig. 4b, where a higher carbide content reduces the plowing action of the wear particles. The ball is protected from wear by the agglomeration of wear particles on its surface which decreases contact with the disc, resulting in a decrease of the diameter of the wear scar of the ball at 250 C, compare Figs. 8a and??b. CONCLUSIONS Within this work, the tribological behavior of several cold work steels in ball- on-disc sliding experiments against austenitic stainless steel balls has been determined in the temperature range between 25 and 250 C. It has been found that the carbide content determines friction behavior of the cold work steels by reducing the adhesive component of the friction coefficient and by increasing the resistance against plowing. At room temperature, the adhesion of austenitic slider material to the cold work steels dominates the wear behavior, plowing the austenitic slider and protecting the discs from wear. The increasing plowing action of loose wear debris and strength loss cause an increase in the wear of the cold work steels at elevated temperatures. Summing up, the carbide content, carbide distribution and carbide size determine the friction and wear behavior of the cold work steels against austenitic stainless steel. ACKNOWLEDGMENTS Financial support of this work by the Technologie Impulse G.m.b.H. in the frame of the K-plus competence center program and by Boehler Edelstahl GmbH and Boehler Uddeholm AC is highly acknowledged. The authors are also grateful to Gerhard Hawranek for performing the SEM and EPMA investigations. REFERENCES [1] N. P. SUH, in "Tribophysics" (Prentice Hall, Boston, 1986).

14 236 6TH INTERNATIONAL TOOLING CONFERENCE [2] F. P. BOWDEN and D. TABOR, in "The Friction and Lubrication of Solids" (Clarendon Press, Oxford, 1954). [3] E. RABINOWICZ, in "Friction and wear of materials" (John Wiley and Sons Inc., New York, 1965). [4] Y. BERTHIER, M. GODET adn M. BRENDLE, Tribology Trans. 32(4) (1989) 490. [5] E. SCHEDIN, in Applications on Stainless Steel 92, Proceedings of the Conference on Applications of Stainless Steels, (Jernkontoret, Stockholm, 1992) p [6] R. A. POGGIE and J. J. WERT, Wear 149 (1991) 209. [7] I. KRAGELSKY, in "Friction and Wear" (Butterworths, London, 1965). [8] N. P. SUH, Wear 25 (1973) 111. [9] F. H. STOTT, Tribology International 31 (1998) 61. [10] S. TUERKER OKTAY, N. P. SUH, Journal of Tribology 114 (1992) 379. [11] E. FINKIN, Materials in Engineering Applications 1 (1979) 154. [12] D. A. RIGNEY, Tribology International 30 (1997) 361.