TEMPERING EFFECT ON CYCLIC BEHAVIOUR OF A MARTENSITIC TOOL STEEL

Transcription

1 TEMPERING EFFECT ON CYCLIC BEHAVIOUR OF A MARTENSITIC TOOL STEEL Z. Zhang Institute of Metal and Technology Dalian Maritime University Dalian China D. Delagnes, G. Bernhart Research Center on Tools, Materials and Processes (CROMeP) Ecole des Mines d Albi-Carmaux Albi, CT Cedex 09 France Abstract A tempered martensitic steel is investigated in order to define a microstructural parameter which may be used in combination with a cyclic constitutive model for numerical simulation of forging dies. A tempering kinetic law was defined in the form of a Johnson-Mehl-Avrami law, by using a "tempering ratio" concept. The tempering ratio takes into account the actual, the as-quenched and the annealed hardness. Mechanical parameters are discussed with respect to tempering ratio and testing temperature and special attention was paid to the cyclic softening behaviour. INTRODUCTION Hot work tool steels are generally used at various tempering states, i.e. with different mechanical properties, depending on requirements of the industrial application (die dimension, workpiece temperature, forging equipment). Moreover, numerous investigations have shown that the die-workpiece interface may reach temperature levels higher than the tempering tempera- 687

2 688 6TH INTERNATIONAL TOOLING CONFERENCE ture. As a consequence, steel may be subjected to a continuous evolution of the microstructure and related properties during his life. Considering the tool lifetime increase, there is an interest in having a good understanding of the steel "ageing" effect on the cyclic fatigue behaviour, which corresponds to there classical loading condition. With the help of simulation, it may then be possible to try to optimise the tool design, if cyclic constitutive models are available. Such models have been investigated over the last years [1, 2, 3, 4] but assume microstructural stability of the steel. As a consequence, model parameters have to be identified for each tempering state and cannot take into account the over ageing during industrial life. In order to overcome these limitations, a new parameter has to be added in the model. This paper reports part of the work [5] performed to introduce a microstructural parameter in a cyclic constitutive model. It describes, in a first part, the tempering test program and the kinetic law defined to follow steel ageing. Then, it shows the effect of steel hardness and test temperature on the cyclic behaviour. Results are discussed in the last part in relation to a more industrial interest: i.e. relation between mechanical properties, hardness, ageing, testing temperature and strain rate. MICROSTRUCTURE AND TEMPERING KINETIC LAW TEMPERING TEST PROGRAM Material investigated in this work is a 55NiCrMoV7 (AISI L6/6F3) hot work tool steel widely used in forging industry for die manufacturing. The classical heat treatment consists in annealing, austenitising, quenching and one tempering. In order to establish a tempering kinetic law, a tempering test program including temperatures between 100 Cand 700 Cand times up to 660 hours was performed (Table 1). Initial state is the as vacuumquenched condition leading to a Rockwell Hardness of 60 HRC. Samples are introduced in a hot furnace and temperature are controlled by a thermocouple welded on the samples. At the end samples are taken out of the hot furnace and slowly cooled. As shown in Table 1, tempering test program consists in two parts: in the first part, time and temperature were chosen in order to permit the verification of the Hollomon and Jaffe [6] relation, as well as the Lifshitz and Wagner [7, 8] equation (rt 3 r3 0 = Kt) defining the increase of carbide size (diameter) during tempering. In the second part, times and temperatures are completed

3 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 689 Table 1. Tempering test program Tempering Tempering duration (hours) temperature( C) Part 1 Part , ,025 0,083 0, , ,63 0,025 0,083 0, , ,3 0,025 0,083 0, , ,87 0,025 0,083 0, , ,025 0,083 0, ,25 2 0,025 0,083 0, ,25 2 0,025 0,083 0, ,25 2 0,025 0,083 0, ,25 2 0,025 0,083 0, ,25 2 0,025 0,083 0,5 1 4 in order to get enough results for kinetic law definition. All samples were polished, and Vickers hardness (HV 0,2 ) was measured under a load of 200 g; thirty indentations were performed for each sample, general scatter was 10 HV. MICROSTRUCTURAL EVOLUTION DURING TEMPER- ING Part of the samples were subjected to further analysis in order to investigate evolution of microstructural features during tempering: this concerns grain sizes, martensitic laths width, length and aspect ratio, carbide volume fractions, compositions and sizes. Those observations were performed using, optical microscopy, Scanning Electron Microscopy, Transmission Electron Microscopy, Electron Dispersive Analysis, X-Ray Analysis and Image Analysis. More details can be found in reference [5]. Major results are summarised in the following: grain sizes and martensitic laths are not modified by tempering as shown in Fig. 1, 2 carbide volume fraction is constant and close to 8.2%, for every tempering temperature

4 690 6TH INTERNATIONAL TOOLING CONFERENCE inter-laths carbides are M 3 C cementites (Fig. 3) whereas intra-laths carbides (Fig. 4) are M 3 C or V 8 C 7 vanadium carbides mean size of intra-laths carbides (which strongly contribute to the strengthening) does not follow the Lifshizt and Wagner equation, but increases rapidly with time and linearly with temperature as shown in Fig. 5a, 5b A linear relation between carbide mean size and Vickers hardness (Fig. 6) was established. This is an indication that among the two mechanisms which may explain the softening during tempering, i.e. dislocation structure evolution and carbide coalescence, the second one has been clearly demonstrated. Dislocation evolution analysis was not performed during this work, but is in progress in an other work [9] TEMPERING RATIO AND KINETIC LAW To establish the tempering kinetic law, it was assumed that tempering corresponds to a phase transformation promoted by a diffusion process between a martensite state towards a ferrite + globular carbide. Such solid phase transformations may be described by the general Johnson, Mehl, Avrami (JMA) [10, 11] relationship as follow: f v = 1 exp( (bt) m ) (1) where f v corresponds to the volume fraction of the new solid phase, m is a material constant and, if we assume that tempering is a thermally activated process, b may be expressed with an Arrhenius equation, ( ) Q b = b 0 exp (2) RT with b 0 constant, Q activation energy, R perfect gaz constant and T temperature in Kelvin. As it was found a direct relation between carbide coalescence and hardness, Vickers hardness was chosen to define the tempering ratio, considering that the tempered state, with a hardness HV is an intermediate state between the as quenched (hardness HV 0 ) and the annealed state (hardness HV ). As a consequence, tempering ratio τ v is defined by the equation τ v = H v H 0 H H 0 (3)

5 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 691 Figure 1. Tempering effect and mean grain size. Figure 2. Tempering effect and martensitic laths width.

6 692 6TH INTERNATIONAL TOOLING CONFERENCE Figure 3. Interlath carbides (SEM). Figure 4. Intralaths carbides (TEM).

7 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 693 Figure 5a. Relation between mean carbide size and tempering time at 600 C. Figure 5b. Relation between mean carbide size and tempering temperature after 2 hours. Figure 6. Relation between mean carbide size and Vickers hardness.

8 694 6TH INTERNATIONAL TOOLING CONFERENCE With such a definition, tempering ratio is between 0 (as quenched state) and 1 (annealed state). For the steel investigated, HV 0 =776 and HV =210. Combining equations (1) and (3), the current hardness can be written as H v = H 0 (H H 0 )(1 exp( (bt) m )) (4) The parameters identifications b 0, Q and m are performed using the SidoLo software [12]. The values obtained are b 0 = (s 1 ), m = and Q = (J K 1 mol 1 ). The latter value is close to the values reported in literature for low or medium alloyed steels and also close to the activation energy of the diffusion of classical alloying elements (Cr, Mn, Ni, V) in ferrite [5]. Figures 7 and 8 show respectively hardness evolution and tempering ratio evolution for short times (less than 4 hours). On Fig. 8 are also drawn the curves coming from the model, showing the validation of the choice of a JMA kinetic law to describe tempering. When complex Figure 7. Vickers hardness and tempering temperature and time. time-temperature routes are followed, differential equation of kinetic law has to be used ( ) m 1 1 m τ v = (1 τ v )mb ln (5) 1 τ v This equation was validated by multiple level tempering experiments [5]. If we consider equation (4), various routes can be followed to reach a given

9 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 695 Figure 8. Tempering ratio with respect to time for various temperature (experimental and simulation). hardness H v ; those routes are defined by the relation ( t exp Q ) = const (6) RT which gives the well known Hollomon and Jaffe relation T(k + log t) = const (7) TEMPERING EFFECT ON CYCLIC FATIGUE BEHAVIOUR FATIGUE TEST PROGRAM In order to investigate the effect of tempering on the cyclic fatigue behaviour, samples have been manufactured with four different hardneses. Tempering conditions, Rockwell and Vickers hardness and corresponding tempering ratio s are reported in Table 2. For each hardness, samples have been subjected to cyclic loads at different temperatures (20 C, 300 C, 400 C, 500 C, 600 C). Tests were carried out with a closed-loop 810 MTS servo-hydraulic testing machine. The round specimen was mounted in water-cooled grips and heating was achieved with a 6 kw induction generator. Temperature was controlled by thermocouples mechanically applied on the middle of the sample. Strain is recorded with a 12 mm gauge length contact extensometer with alumina rods. The test itself consists in a total

10 696 6TH INTERNATIONAL TOOLING CONFERENCE strain amplitude reversed (+0.8/-0.8 %) fatigue test under triangular wave form divided in two stages: during the first stage, strain rate is fixed on a level of s 1, and number of cycles were selected in order to reach a near constant cumulated plastic strain close to 4 mm/mm whatever the test temperature and hardness. As an example, Fig. 9 shows the evolution of the semi-stress amplitude with the cumulated plastic strain for the 50 HRc samples at all temperatures. during the second stage, strain rate is varied form s 1 to s 1 and s 1, but only three cycles were performed at each strain rate. This allows to get information on the steel strain rate sensitivity that may modify considerably the mechanical properties, especially at high temperatures. Fig. 10 shows the change of stressstrain loops for a 42 HRc specimen at 600 C. The three levels of strain rates have been selected to cover the industrial strain rate conditions, between mechanical and hydraulic forging. Figure 9. hardness. Semi-stress amplitude with respect to cumulated plastic strain for the 50 HRC

11 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 697 Table 2. Calculated and experimental hardness of fatigue equivalent sample Initial tempering temperature ( C) temperature ( C) duration (s) Fatigue test Hardness before test (HV 0,2) Hardness after test (HV 0,2) Initial tempering temperature ( C) Temperature of equivalent ageing ( C) Equivalent temperature ageing duration (s) Hardness after ageing (HV 0,2) Hardness calculated with kinetic law (HV 0,2)

12 698 6TH INTERNATIONAL TOOLING CONFERENCE Figure 10. Stress-total strain for 3 strain rates at 600 Cfor a 42 HRC sample. CYCLIC BEHAVIOUR IN RELATION WITH HARDNESS As shown in Fig. 9, tempered martensitic steels undergo cyclic softening during fatigue testing. This softening is divided in two parts which are generally explained by the rapid (exponential) modification of the dislocation density and structure for the first one, and the modification of dislocation sub-structure and carbide morphology for the second linear one. Figures 11 and 12 show the semi-stress amplitude evolution at 20 Cand 600 Cfor the four different hardneses. At 20 C, curves are clearly separated: stress amplitudes are increasing with the hardness level. Softening is very important and may reach level as high as 250 MPa. At 600 C, three of the curves are very close, and only the 35 HRC sample is lower. This surprising result can easily be explained considering the tempering ratio after testing which are respectively 0.79, 0.67, 0.68 and 0.72 for the initial 35 HRC, 42 HRC, 45.5 HRC and 50 HRC samples. It becomes obvious that samples with similar tempering ratio show same cyclic softening behaviour. The origin is related to the fatigue testing procedure which requires a heating time and a temperature stabilisation time of at least 75 seconds. When using the tempering kinetic law, tempering ratio may be calculated after heating and after fatigue testing. Results show that when testing temperature is greater than the tempering temperature, samples undergo a rapid

13 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 699 Figure 11. Semi-stress amplitude evolution at room temperature for the 50, 45.5, 42 and 35 HRC samples. Figure 12. samples. Semi-stress amplitude evolution at 600 Cfor the 50, 45.5, 42 and 35 HRC microstructural evolution to reach a similar state. Moreover, this modification continues during fatigue testing whereas this is not the case for samples tested at temperatures lower than the tempering temperature.

14 700 6TH INTERNATIONAL TOOLING CONFERENCE As a consequence, mechanical properties may change drastically when the steel is subjected to temperatures higher than tempering temperatures, even for short times, and the tempering kinetic law seems to be a good tool to get a reliable estimation of the hardness variation associated to the microstructural evolution. DISCUSSION TEMPERING AND TEST TEMPERATURE EFFECT ON MECHANICAL PROPERTIES Previous results have been obtained in order to define a cyclic constitutive model able to describe the fatigue behaviour and the effect of timetemperature ageing on this cyclic behaviour. More information can be found in references 1 to 5. It allows to reproduce all experimental features of the fatigue behaviour of tempered martensitic steels: fatigue loops using two kinematic variables (back stress), two phase cyclic softening using two isotropic variables (drag stress), strain memory accounting for the increase of softening when strain range is increased, and at least time-temperature ageing with the tempering ratio variable (equation (5)). After parameter identification, it is possible to predict with simulation all mechanical properties, whatever tempering and testing history. Figure 13. Yield stress with respect to temperature for different initial hardness.

15 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 701 Figure 14. Strain rate effect on yield stress for the 50 and 35 HRC material. In the first example, model results are presented to discuss the evolution of the classical yield stress (R 0.2 ). In Figures 13 and 14, yield stress is drawn with respect to testing temperature for the different tempering conditions and for three different strain rates (10 2 s 1, 10 3 s 1, 10 4 s 1 ) that correspond to typical strain rates seen by the various forging equipments. The more the testing temperature is high, the closer are the yield stresses, and the more the yield stress differs between low and high strain rates (this difference reaches values as high as 150 MPa at 600 C). As a consequence, the useful yield stress for die design is equipment dependent and not a material constant. Moreover, this yield stress decreases when subjected to cyclic fatigue due to the material softening, as reported in Fig. 12 at 600 C. Combined influences of temperature, strain rate and cyclic softening may decrease the conventional yield stress by a factor of two in the worst case (at highest temperatures). The relation between yield stress and tempering ratio is plotted in Fig. 15 for different temperatures. It can be seen that there is a linear relationship, but the higher the testing temperature is, the lower is the slope. That means that the important effect of the tempering ratio level at low temperatures vanishes at high temperature, where the heating up to testing temperature has a major effect.

16 702 6TH INTERNATIONAL TOOLING CONFERENCE Figure 15. Yield stress in relation with tempering ratio for different test temperatures. Figure 16. Simulated forging cycle. In the second example the evolution of an initial 45.5 HRC material exposed to a forging cycle of 20s between 400 Cand 600 C(Fig. 16) is simulated. It can be seen in Fig. 17, that tempering ratio increases with the number of cycles, as a result of the in-service temperature which is higher than the tempering temperature (460 C). Room temperature and 600 Cyield

17 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 703 Figure 17. Material evolution with respect to the number of forging cycles. stresses predicted by the model are also reported in Fig. 17, in order to show the capacity of the model to take into account the material ageing. HARDNESS EVOLUTION AFTER CYCLIC FATIGUE TEST After fatigue testing, up to a cumulated plastic strain close to 4mm/mm (i.e. without rupture), Vickers hardness was measured in the middle of each fatigue specimen. The amount of hardness decrease is reported in Fig. 18 with respect to the difference between the testing and tempering temperature. If testing temperature is lower than tempering temperature, hardness decrease is low and no evolution occurs at room temperature. A maximum decrease of 50 HV at tempering temperature is measured. Conversely, when fatigue testing is performed at temperatures more than 50 Chigher than tempering temperature, hardness decrease is much higher and can reach 50% of the initial hardness as for example for the 50 HRC sample tested at 600 C. To distinguish the temperature and load effect on hardness evolution, some special ageing were performed reproducing exactly the time-temperature history of the fatigue test specimens. It can be seen in Table 2, that the calculated and measured hardness of those fatigue equivalent time-temperature ageing are very close (that validates again the reliability of the kinetic law).

18 704 6TH INTERNATIONAL TOOLING CONFERENCE Figure 18. Relation between test and tempering temperatures and hardness decrease Figure 19. Temperature and mechanical hardness decrease. This allows now to discuss the effect of temperature and cyclic fatigue on the hardness evolution. Two examples are reported in Fig. 19 for the 45.5 and 50 HRC samples. Three main conclusions can be drawn from these results:

19 Tempering Effect on Cyclic Behaviour of a Martensitic Tool Steel 705 if testing temperature is more than 50 Clower than tempering temperature, no temperature ageing happens and hardness is not modified by fatigue testing, if testing temperature is close to the tempering one (-50 Cto +50 C), temperature ageing is limited, but hardness evolution due to cyclic loading is observed, if testing temperature is more than 50 Chigher than tempering temperature, hardness variation due to the temperature becomes predominant, but a second evolution related to cyclic fatigue is noticed. It seems that the amount of hardness decrease due to fatigue is nearly constant to a value of 50 HV. In conclusion, it seems that there is a synergetic effect between temperature ageing and cyclic fatigue only when testing temperature is higher than tempering temperature. It induces a complementary decrease of hardness of 50 HV, which seems to be a constant whatever the initial tempering state of the material. CONCLUSION The results reported are part of more important research activity that aims to work out a cyclic constitutive model for tempered martensitic tool steels [5]. As numerous tempering conditions are used in industry for the same steel grade, and as in service temperatures may overshoot the tempering one, at least for short times, a parameter that takes into account the microstructural evolution has to be added into the model. Microstructural investigation on various tempered samples have shown that only the intra-laths carbides undergo changes during tempering (dislocations were not investigated in this work); a linear relation between hardness and carbides mean size was established. As a consequence, hardness was chosen to define a tempering ratio parameter. The evolution of this parameter was described with a Johnson-Mehl-Avrami kinetic law. Cyclic response was discussed with respect to initial hardness. It was shown that if the testing temperature is higher than the tempering one, samples undergo evolutions, even for short times of sample heating before testing which may modify the microstructural state and the resulting mechanical properties. Hardness measurements on fatigue tested samples have been performed and were analysed in combination with

20 706 6TH INTERNATIONAL TOOLING CONFERENCE the kinetic law. It was shown that if testing temperature is lower than tempering one, no significant hardness evolution was measured. In other cases, temperature ageing becomes predominant, and superposed fatigue induces a synergetic effect which leads to a supplementary hardness decrease of 50 HV, which seems to be a material constant, whatever the hardness and testing temperature. ACKNOWLEDGMENT The authors gratefully acknowledge THYSSEN France for supplying the steel rods. REFERENCES [1] G. BERNHART, G. MOULINIER, O. BRUCELLE and D. DELAGNES, Int. Journal of Fatigue, 21 (1999) [2] Z. ZHANG, D. DELAGNES and G. BERNHART. Proceedings of 5th Int. Tooling Conference, , Sept 1999, Leoben, Austria. [3] Z. ZHANG, D. DELAGNES and G. BERNHART, Int. Journal of Fatigue, 24 (2002) [4] V. VELAY, G. BERNHART, Z. ZHANG and L. PENAZZI, Proceedings of High- Temperature Fatigue Conference, CAMP 2002, April, paderborn, germany, pp [5] Z. ZHANG. PhD thesis, ENSMP (2002) [6] J. H. HOLLOMON and L. D. JAFFE, Trans. AIME, 162 (1945) p.727. [7] I. M. LIFSHITZ and V. V. SLYOZOV, J. Phys. Chem. Solids, 19, (1961) p. 35. [8] C. WAGNER. Zeitschrift f'ür Electrochemie, Bd 65, Nr 7/8 (1961) [9] N. MEBARKI, P. LASMESLE, F. DELMAS, D. DELAGNES, C. LEVAILLANT, 6th tooling conference, Karlstad, [10] W. A. JOHNSON and R. F. MEHL. Transactions of the american Institute of mining, Metallurgical and petroleum engineers, 135 (1939) [11] M. AVRAMI, Journal of Chemical Physics, 7 (1939) [12] P. PILVIN, SoDoLo, User Manual (1985).