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Integrating physical and numerical simulation techniques to design the hot forging process of stainless steel turbine blades
International Journal of Machine Tools & Manufacture 44 (2004) 945–951 www.elsevier.com/locate/ijmactool
Integrating physical and numerical simulation techniques to design the hot forging process of stainless steel turbine blades P.F. Bariani, S. Bruschi , T. Dal Negro DIMEG, University of Padova, via Venezia 1, 35131 Padova, Italy Received 23 December 2003; accepted 21 January 2004
Abstract The paper presents a joint application of finite-element-based numerical simulation and real-material-based physical simulation techniques for design and optimisation of the hot forging operations to manufacture high strength stainless steel turbine blades. 2D simulations of the forging steps carried out using a suitably calibrated finite element model are combined with systematic analysis of microstructure evolution during forging experiments, with particular care to formation of the brittle d-ferrite phase at high temperatures. A correlation is established between microstructure and thermal and mechanical parameters characterising forging operations. On the bases of numerical and experimental results, the actual forging process is re-designed, reducing the total number of forging steps. Industrial trials, conducted with the optimised process parameters, demonstrate the effectiveness of the developed procedure. # 2004 Elsevier Ltd. All rights reserved. Keywords: Blade forging; Microstructure; FEM calibration
1. Introduction Due to complexity of the shape and low formability limits of the material, precision forging of turbine blades requires advanced tools for design of forging operations and close control of the process variables. In recent years, numerical simulation techniques, mainly based on finite element (FE) codes, have been extensively utilised to investigate blade forging, providing useful information in the prediction of several process parameters, as temperature distribution inside workpiece and dies, pressure on dies and forging loads [1–7]. For most of the alloys utilised in manufacturing turbine blades, such as stainless steels, titanium- and nickelbased alloys, process parameters have to be chosen and kept within narrow ranges due to material strict workability windows as well as close control in microstructure specifications of the final component. This represents the Corresponding author. Tel.: +39-049-8276822; fax: +39-0498276816. E-mail address: [email protected] (S. Bruschi).
0890-6955/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijmachtools.2004.01.020
main constraint in designing and optimising the forging process itself. One of the major goal in forging high strength stainless steels components is to obtain a welldefined microstructure at the end of the process characterised by a volume fraction of d-ferrite phase lower than a critical value: a higher percentage of d-ferrite phase weakens the mechanical strength and the ductility of the component during its service life. Consequently, temperature peaks during forging operations have not to exceed a critical value in order to keep the volume fraction of d-ferrite phase as lowest as possible [8–11]. Sound knowledge of the material rheological behaviour as well as accurate prediction of the microstructure evolution are therefore the prerequisites for the technical and economical success of the process design. Recent scientific literature [4–7] confirms that the research efforts are addressed to the development of models of the blade forging process where microstructure is strictly correlated to process parameters, i.e. forging temperature and strain rate. The paper presents a procedure for design and optimisation of the hot forging process of St70AH (DIN
946
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Table 1 Chemical composition of St70AH (DIN X5CrNiMoCuNb14-5) C
Si
Mn
Cr
Mo
Ni
Nb
0.07
0.7
1
15
2
5
0.5
X5CrNiMoCuNb14-5 whose chemical composition is reported in Table 1) turbine blades based on combined use of experimental and numerical techniques. The numerical analysis consists of 2D FE simulations of the deformation steps in the forging process, carried out using the commercial Forge 22 code suitably calibrated to the specific problem. The simulations provide accurate description of the thermal and mechanical events occurring during the process and that directly affect the microstructural evolution. The experimental analysis is addressed to properly calibrate the numerical model and combines (i) physical simulation experiments on real-material samples and (ii) systematic analysis of microstructure. Physical simulation experiments consist of hot compression tests, aimed at determining the material rheological behaviour as well as investigating the evolution of d-ferrite phase with process parameters, and laboratory forging experiments on a pilot plant to evaluate the boundary conditions during forging, i.e. heat transfer and friction coefficients. The developed procedure is applied to an industrial forging process where St70AH turbine blades are manufactured through three steps. The design and optimisation of the process is focused on the reduction of forging steps, thanks to a significant increase of the billet initial temperature.
2. The FEM model The current industrial process of St70AH blade forging consists of three forging steps carried out on a screw press. Forging temperature of the first step is v v 1100 C, the second one is at 1090 C and the third v one at 1070 C. After each blow, the blade is cooled in air until room temperature, the surface is grounded to remove surface cracks, blanks are then glazed and reheated up to the forging temperature with a soaking time at this temperature to uniform the thermal distribution inside the entire material volume. According to this procedure, the material conditions at the beginning of each step can be assumed to be the same for all the steps, then neglecting the effects of the previous thermal and mechanical history on the material flow strength. The combination of the process parameters in each step and the number of forging blows assure the desired microstructure at the end of the process.
Because of the thin section, deformation and thermal phenomena are much more significant in the airfoil region rather than in the root. In the airfoil, plane deformation conditions prevail and a 2D analysis of cross-sections is well appropriate. A 2D thermal-plastic coupled FE analysis was carried out to simulate the forging steps by using the commercial FE code Forge 22. A cross-section in the middle of the airfoil was taken as the representative one for the analysis. To obtain feasible results from the numerical simulations, an accurate calibration of the FE code is well known to be required [12,13]. In this work, particular attention was paid to calibration of those parameters that show to significantly affect the thermal and mechanical events occurring during forging operations. These parameters include the material flow strength under hot deformation conditions and the boundary conditions data, as heat transfer and friction coefficients at the die–workpiece interface. 2.1. Flow stress data The St70AH flow strength was evaluated through single-step hot compression experiments and its sensitivity to temperature and strain rate was analysed. All the compression tests were carried out on the thermomechanical simulator Gleeble 38002 at DIMEG lab, capable to assure a very accurate control of strain and strain rate along the whole deformation stroke. The experiments were performed on cylindrical samples of St70AH, 12 mm diameter and 14 mm long, machined form a drawn bar. During the test, the sample was v heated up to the testing temperature at 10 C/s, held at this temperature for 30 s and then compressed to the required amount of strain at constant strain rate. After deformation, the specimens were air-cooled and then analysed by optical microscope to evaluate their microstructure. The experimental plan is presented in Table 2. Fig. 1 presents the high material sensitivity to temperature at strain rate equal to 10 1/s. As the strain rate varies along the deformation stroke during the actual forging step, hot compression tests at continuously varying strain rate were also performed, aimed at evaluating the effects of strain rate histories on the material instantaneous flow strength [14]. In Fig. 2, the comparison between flow curves Table 2 Experimental plan to evaluate St70AH flow stress under single-step deformation conditions v
e
e_ (1/s)
T ( C)
1 1 1
1 10 50
1000 1000 1000
1100 1100 1100
1180 1180 1180
1250 1250 1250
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
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Table 3 Thermo-physical characteristics of St70AH
Fig. 1. St70AH sensitivity to temperature at strain rate equal to 10 1/s.
from constant and varying strain rate experiments is presented: the comparison is performed where the instantaneous values of the process parameters (amount of strain, strain rate and temperature) are the same. From the results presented in Fig. 2, the influence of the effects of varying strain rate on the material flow strength can be considered as negligible [15,16]. Rheological constants to be used in the constitutive equation implemented in the FE code were then calculated through non-linear regression analysis from flow stress data obtained during experiments at constant strain rate. Due to the very thin sections of the blade airfoil, temperature inside the blade airfoil during forging operations is greatly affected by thermal and mechanical interactions with the surrounding environment. Relevant thermo-physical characteristics of St70AH utilised in the FE model are reported in Table 3. 2.2. Boundary conditions The boundary conditions that mostly affect the accuracy of the numerical results are the heat transfer
Density (g/cm3)
Conductivity (W/m K)
Specific heat (J/g C)
7.85
16.3
1.06
v
coefficients (HTC) and the friction factor at the die– workpiece interface. The HTC value is identified through a combination of: (i) measurements of temperature histories performed at suitable locations inside the dies during the deformation phase; (ii) an FE model, that through fully coupled thermal and mechanical analysis relates the temperature readings to the unknown HTC distributions, and (iii) inverse analysis. This calculates the HTC values that minimise the difference between the temperature histories measured and those calculated by FEM [17]. Fig. 3 shows a schematic representation of the deformation phase of the forging operation. Deformation is carried out under plane-strain conditions so as to approximate the forging of turbine airfoil sections. This particular operation was chosen as an ideal case for applying this approach because of the thickness variations in the final section, the consequent distributions of contact pressure and sliding velocity at the interface and the high surface extension involved. The billet is a 30 mm diameter and 50 mm long cylinder. A laboratory forging plant based on a 3500 kN Vaccari2 screw press, a fast induction furnace and an automatic loading and unloading device, was used in the deformation experiments. Billet temperature at the v furnace exit was 1200 C and the deformation phase in the forging cycle lasted 0:6 101 s. In the described operation conditions, an average surface HTC of 30 kW/m2 K has been identified. On the same forging plant, hot ring compression tests have been performed to evaluate the friction factor. Die material was AISI H11 with a surface hardness, after grinding, in the range of 48–52 HRC. The St70AH ring geometry is 10 mm in height, with outside diameter 30 mm and inside diameter 15 mm. Testing v temperatures are in the range 1050–1200 C. The identified friction factor is 0.35. 2.3. Microstructural analysis
Fig. 2. St70AH sensitivity to strain rate history effects during deformation.
The maximum amount of d-ferrite phase considered acceptable in an St70AH blade corresponds to 5%; for higher percentage of d-ferrite phase, mechanical strength and ductility together with fatigue resistance of the alloy drastically decrease [18]. To determine the actual evolution of d-ferrite phase within the range of testing temperatures and strain rates, the microstructure of the specimens deformed
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Fig. 3.
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Geometry of the simulative test for heat transfer evaluation and locations of thermocouples embedded into the upper and lower die.
according to Table 2 was observed and the amount of d-ferrite phase evaluated. The specimens were cut into transverse sections, polished to 1 lm and then etched by electrolysis to reveal the microstructure. For each specimen, the microstructure was observed in three different zones that are represented in Fig. 4. The influence of both testing temperature and strain rate on the amount of d-ferrite phase was evaluated: for all the testing temperatures, it is proved that strain rate does not influence the microstructure in a significative way. As example, in Fig. 5, St70AH microv structures at T ¼ 1100 C at different strain rate are shown. Temperature, instead, strongly affects the d-ferrite phase formation (Fig. 6 on the left), leading to an v unacceptable microstructure at T ¼ 1250 C, characterised by more than 5% of d-ferrite. Fig. 6 on the right shows also the influence of soaking time just before deformation: if the specimen is held at the testing temperature for 20 min (instead of 30 s), its microstructure results almost unchanged for temperatures v v less than T ¼ 1180 C, while at T ¼ 1250 C the amount of d-ferrite phase drastically increases. On the basis of results of microstructural observations, the control of temperature inside the workpiece allows to control the amount of d-ferrite phase. If temperature inside the workpiece during deformation is v kept lower than 1180 C, then the blade microstructure can be considered acceptable.
Fig. 4. Zones of microstructural analysis on the deformed specimen.
3. Results of FE simulations By using the calibrated FE model, the temperature distribution inside the airfoil region was calculated at the end of each forging blow. The maximum value of temperature was compared with the limit value that corresponds to the critical amount of d-ferrite phase. Particular attention was paid to temperature increase in the central zone and along the middle line of the airfoil section. Temperature of the material close to the interface with the dies and in the flash was not taken into account since these portions of material are removed by machining after forging. Fig. 7 shows the temperature distribution inside the cross-section at the end of the second (left) and the third steps (right). The highest values of temperature in the central zone v of the airfoil region are always lower than 1100 C at the end of the relavant forging steps, particularly at the end of the second step, which is the most significant as temperature increase in the central zone of the airfoil. From the microstructural stand-point, the amount of d-ferrite phase will be within the acceptable limits.
4. The optimised forging process The accurate analysis of the temperatures in the airfoil at the end of the second and the third forging steps revealed that the temperature and the corresponding amount of d-ferrite phase are rather lower than the respective critical values. The lack of a quantitative knowledge of thermal and microstructural events during the forging process induces the process planner to be too conservative. Consequently, the design of the current industrial process proves to be quite inefficient. The quite high number of forging blows causes long lead times to process the batch and multiplication of the costly operations to carry out at the end of each blow. It becomes than clear that the total number of forging blows could be reduced, increasing the initial billet temperature and, at the same time, maintaining the maximum amount of d-ferrite phase lower than 5%.
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Fig. 5.
Fig. 6.
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v
Microstructure sensitivity to strain rate at T ¼ 1100 C.
Microstructure sensitivity to temperature at strain rate equal to 50 1/s and with different soaking times before deformation.
Thanks to St70AH high temperature sensitivity (Fig. 1), temperature increase leads to a substantial reduction of the total forging load, thus reducing the overall production costs.
The actual forging process was re-designed reducing the number of forging steps from three to two: the initial v v billet temperature was increased to 1150 C (50 C more than the actual forging process).
Fig. 7. Temperature distribution at the end of the second (left) and the third (right) steps of the actual process.
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P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Fig. 8.
Temperature distribution at the end of the second (left) and the third (right) steps of the optimised process.
transformation. In the case of St70AH, accurate prediction of temperature and d-ferrite phase distributions has allowed the reduction of the number of forging steps of the actual process from three to two, increasing the overall initial process temperature and thus leading to a more effective forging process. Industrial trials have demonstrated the effectiveness of this forging process optimisation.
Acknowledgements Fig. 9. Example of blade microstructure in the central zone of the airfoil after heat treatment.
As shown in Fig. 8, the values of temperature calculated in the cross-section at the end of the second blow are almost everywhere lower than the critical one corresponding to unacceptable amount of d-ferrite phase. v Only in the flash area temperature reaches 1180 C, corresponding to an amount of d-ferrite phase greater than 5%. 4.1. Industrial validation Fig. 9 shows the airfoil microstructure of an St70AH blade, forged with the optimised value of initial temperature. The microstructure is relevant to the blade after heat treatment: however, it does not present any evident tracks of d-ferrite phase (note that d-ferrite phase after formation at high temperatures remains in the component even if heat treated).
5. Concluding remarks Accurate knowledge of the thermal and mechanical events determining the evolution of microstructure during deformation phases is a primary prerequisite for design and optimisation of forging St70AH turbine blades. The combination of an FE model calibrated to the specific forging problem with a systematic microstructure analysis on samples prepared through physical simulation experiments proved to be an effective and successful tool in predicting temperatures and phase
The work on which this paper is based is a part of a scientific co-operation between PIETRO ROSA TBM and DIMEG- University of Padova. The Authors wish to thank PIETRO ROSA TBM for collaboration in analysing the microstructural results and conducting the industrial trials and validation. References [1] T. Altan, N. Akgerman, Application of CAD/CAM in forging turbine and compressor blades, J. Eng. Power 98 (1978) 290– 298. [2] C. Boer, N. Rebelo, G. Schroeder, Process Modelling of Metal Forming and Thermomechanical Treatment, Springer Verlag, Berlin, 1986. [3] Z.M. Hu, J.W. Brooks, T.A. Dean, Experimental and theoretical analysis of deformation and microstructure evolution in the hot die forging of titanium alloy aerofoil sections, J. Mater. Process. Technol. 88 (1999) 251–265. [4] L. Yuli, D. Kun, Z. Mei, Z. Fuwey, Physical modelling of blade forging, J. Mater. Process. Technol. 99 (2000) 141–144. [5] P.F. Bariani, M. Dal Negro, M. Fioretti, Hot workability studies of Nimonic 80A applied to the nett-shape forging of aerofoil blades, Proceedings of the AMST, Udine, Italy, 1999. [6] P.F. Bariani, T. Dal Negro, M. Fioretti, Joint use of physical and numerical simulation techniques in predicting process parameters evolution and final microstructure in forging Nimonic 80A turbine blades, Proceedings of the 2nd Esaform, Guimares, Portugal, 1999. [7] H. Ou, R. Balendra, Modelling techniques for nett-shape forging of turbine blades, Proceedings of the ICFT, 1998. [8] P.F. Bariani, G. Berti, T. Dal Negro, S. Masiero, Experimental evaluation and FE simulation of thermal cycle at tool surface during cooling and deformation phases in hot and warm forging operations, Ann. CIRP 51 (1) (2002) 219–222. [9] P.F. Bariani, T. Dal Negro, S. Masiero, Influence of oxide scale on the heat transfer coefficient during hot forging of turbine
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[10] [11]
[12]
[13]
blades, Proceedings of the 5th ESAFORM, Kracow, Poland, 2002. Y.V.R.K. Prasad, T. Seshacharyulu, Processing maps for hot working of titanium alloys, Mater. Sci. Eng. A243 (1998) 82–88. G. Shen, S.L. Semiatin, T. Altan, Investigation of flow stress and microstructure development in non-isothermal forging of Ti-6242, J. Mater. Process. Technol. 36 (1993) 303–319. P.H.S. Cardoso, C. Kwietniewski, J.P. Porto, A. Reguly, T.R. Strohacher, The influence of delta ferrite in the AISI 416 stainless steel hot workability, Mater. Sci. Eng. A351 (2003) 1–8. J. Glownia, B. Kalandyk, K. Hubner, Delta ferrite predictions for cast duplex steels with high nitrogen content, Mater. Charact. 47 (2001) 149–165.
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[14] E. Doege, N. Werner, Algorithms for calibration of finite element models, Prod. Eng. 5/1 (1998) 127–130. [15] P.F. Bariani, S. Bruschi, T. Dal Negro, Reliability of flow stress evaluation in numerical simulation of hot forging operations, Proceedings of the 6th ICTP, Nuremberg, Germany, 1999. [16] P.F. Bariani, S. Bruschi, T. Dal Negro, Analysis of varying strain rate conditions on the material flow stress in hot forging operations, Proceedings of the 7th ICTP, Yokohama, Japan, 2002. [17] Z.M. Hu, J.W. Brooks, T.A. Dean, The interfacial heat transfer coefficient in hot die forging of titanium alloy, J. Mech. Eng. 212 (1998) 485–496. [18] Materials Properties Handbook, ASM International.
Integrating physical and numerical simulation techniques to design the hot forging process of stainless steel turbine blades P.F. Bariani, S. Bruschi , T. Dal Negro DIMEG, University of Padova, via Venezia 1, 35131 Padova, Italy Received 23 December 2003; accepted 21 January 2004
Abstract The paper presents a joint application of finite-element-based numerical simulation and real-material-based physical simulation techniques for design and optimisation of the hot forging operations to manufacture high strength stainless steel turbine blades. 2D simulations of the forging steps carried out using a suitably calibrated finite element model are combined with systematic analysis of microstructure evolution during forging experiments, with particular care to formation of the brittle d-ferrite phase at high temperatures. A correlation is established between microstructure and thermal and mechanical parameters characterising forging operations. On the bases of numerical and experimental results, the actual forging process is re-designed, reducing the total number of forging steps. Industrial trials, conducted with the optimised process parameters, demonstrate the effectiveness of the developed procedure. # 2004 Elsevier Ltd. All rights reserved. Keywords: Blade forging; Microstructure; FEM calibration
1. Introduction Due to complexity of the shape and low formability limits of the material, precision forging of turbine blades requires advanced tools for design of forging operations and close control of the process variables. In recent years, numerical simulation techniques, mainly based on finite element (FE) codes, have been extensively utilised to investigate blade forging, providing useful information in the prediction of several process parameters, as temperature distribution inside workpiece and dies, pressure on dies and forging loads [1–7]. For most of the alloys utilised in manufacturing turbine blades, such as stainless steels, titanium- and nickelbased alloys, process parameters have to be chosen and kept within narrow ranges due to material strict workability windows as well as close control in microstructure specifications of the final component. This represents the Corresponding author. Tel.: +39-049-8276822; fax: +39-0498276816. E-mail address: [email protected] (S. Bruschi).
0890-6955/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijmachtools.2004.01.020
main constraint in designing and optimising the forging process itself. One of the major goal in forging high strength stainless steels components is to obtain a welldefined microstructure at the end of the process characterised by a volume fraction of d-ferrite phase lower than a critical value: a higher percentage of d-ferrite phase weakens the mechanical strength and the ductility of the component during its service life. Consequently, temperature peaks during forging operations have not to exceed a critical value in order to keep the volume fraction of d-ferrite phase as lowest as possible [8–11]. Sound knowledge of the material rheological behaviour as well as accurate prediction of the microstructure evolution are therefore the prerequisites for the technical and economical success of the process design. Recent scientific literature [4–7] confirms that the research efforts are addressed to the development of models of the blade forging process where microstructure is strictly correlated to process parameters, i.e. forging temperature and strain rate. The paper presents a procedure for design and optimisation of the hot forging process of St70AH (DIN
946
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Table 1 Chemical composition of St70AH (DIN X5CrNiMoCuNb14-5) C
Si
Mn
Cr
Mo
Ni
Nb
0.07
0.7
1
15
2
5
0.5
X5CrNiMoCuNb14-5 whose chemical composition is reported in Table 1) turbine blades based on combined use of experimental and numerical techniques. The numerical analysis consists of 2D FE simulations of the deformation steps in the forging process, carried out using the commercial Forge 22 code suitably calibrated to the specific problem. The simulations provide accurate description of the thermal and mechanical events occurring during the process and that directly affect the microstructural evolution. The experimental analysis is addressed to properly calibrate the numerical model and combines (i) physical simulation experiments on real-material samples and (ii) systematic analysis of microstructure. Physical simulation experiments consist of hot compression tests, aimed at determining the material rheological behaviour as well as investigating the evolution of d-ferrite phase with process parameters, and laboratory forging experiments on a pilot plant to evaluate the boundary conditions during forging, i.e. heat transfer and friction coefficients. The developed procedure is applied to an industrial forging process where St70AH turbine blades are manufactured through three steps. The design and optimisation of the process is focused on the reduction of forging steps, thanks to a significant increase of the billet initial temperature.
2. The FEM model The current industrial process of St70AH blade forging consists of three forging steps carried out on a screw press. Forging temperature of the first step is v v 1100 C, the second one is at 1090 C and the third v one at 1070 C. After each blow, the blade is cooled in air until room temperature, the surface is grounded to remove surface cracks, blanks are then glazed and reheated up to the forging temperature with a soaking time at this temperature to uniform the thermal distribution inside the entire material volume. According to this procedure, the material conditions at the beginning of each step can be assumed to be the same for all the steps, then neglecting the effects of the previous thermal and mechanical history on the material flow strength. The combination of the process parameters in each step and the number of forging blows assure the desired microstructure at the end of the process.
Because of the thin section, deformation and thermal phenomena are much more significant in the airfoil region rather than in the root. In the airfoil, plane deformation conditions prevail and a 2D analysis of cross-sections is well appropriate. A 2D thermal-plastic coupled FE analysis was carried out to simulate the forging steps by using the commercial FE code Forge 22. A cross-section in the middle of the airfoil was taken as the representative one for the analysis. To obtain feasible results from the numerical simulations, an accurate calibration of the FE code is well known to be required [12,13]. In this work, particular attention was paid to calibration of those parameters that show to significantly affect the thermal and mechanical events occurring during forging operations. These parameters include the material flow strength under hot deformation conditions and the boundary conditions data, as heat transfer and friction coefficients at the die–workpiece interface. 2.1. Flow stress data The St70AH flow strength was evaluated through single-step hot compression experiments and its sensitivity to temperature and strain rate was analysed. All the compression tests were carried out on the thermomechanical simulator Gleeble 38002 at DIMEG lab, capable to assure a very accurate control of strain and strain rate along the whole deformation stroke. The experiments were performed on cylindrical samples of St70AH, 12 mm diameter and 14 mm long, machined form a drawn bar. During the test, the sample was v heated up to the testing temperature at 10 C/s, held at this temperature for 30 s and then compressed to the required amount of strain at constant strain rate. After deformation, the specimens were air-cooled and then analysed by optical microscope to evaluate their microstructure. The experimental plan is presented in Table 2. Fig. 1 presents the high material sensitivity to temperature at strain rate equal to 10 1/s. As the strain rate varies along the deformation stroke during the actual forging step, hot compression tests at continuously varying strain rate were also performed, aimed at evaluating the effects of strain rate histories on the material instantaneous flow strength [14]. In Fig. 2, the comparison between flow curves Table 2 Experimental plan to evaluate St70AH flow stress under single-step deformation conditions v
e
e_ (1/s)
T ( C)
1 1 1
1 10 50
1000 1000 1000
1100 1100 1100
1180 1180 1180
1250 1250 1250
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
947
Table 3 Thermo-physical characteristics of St70AH
Fig. 1. St70AH sensitivity to temperature at strain rate equal to 10 1/s.
from constant and varying strain rate experiments is presented: the comparison is performed where the instantaneous values of the process parameters (amount of strain, strain rate and temperature) are the same. From the results presented in Fig. 2, the influence of the effects of varying strain rate on the material flow strength can be considered as negligible [15,16]. Rheological constants to be used in the constitutive equation implemented in the FE code were then calculated through non-linear regression analysis from flow stress data obtained during experiments at constant strain rate. Due to the very thin sections of the blade airfoil, temperature inside the blade airfoil during forging operations is greatly affected by thermal and mechanical interactions with the surrounding environment. Relevant thermo-physical characteristics of St70AH utilised in the FE model are reported in Table 3. 2.2. Boundary conditions The boundary conditions that mostly affect the accuracy of the numerical results are the heat transfer
Density (g/cm3)
Conductivity (W/m K)
Specific heat (J/g C)
7.85
16.3
1.06
v
coefficients (HTC) and the friction factor at the die– workpiece interface. The HTC value is identified through a combination of: (i) measurements of temperature histories performed at suitable locations inside the dies during the deformation phase; (ii) an FE model, that through fully coupled thermal and mechanical analysis relates the temperature readings to the unknown HTC distributions, and (iii) inverse analysis. This calculates the HTC values that minimise the difference between the temperature histories measured and those calculated by FEM [17]. Fig. 3 shows a schematic representation of the deformation phase of the forging operation. Deformation is carried out under plane-strain conditions so as to approximate the forging of turbine airfoil sections. This particular operation was chosen as an ideal case for applying this approach because of the thickness variations in the final section, the consequent distributions of contact pressure and sliding velocity at the interface and the high surface extension involved. The billet is a 30 mm diameter and 50 mm long cylinder. A laboratory forging plant based on a 3500 kN Vaccari2 screw press, a fast induction furnace and an automatic loading and unloading device, was used in the deformation experiments. Billet temperature at the v furnace exit was 1200 C and the deformation phase in the forging cycle lasted 0:6 101 s. In the described operation conditions, an average surface HTC of 30 kW/m2 K has been identified. On the same forging plant, hot ring compression tests have been performed to evaluate the friction factor. Die material was AISI H11 with a surface hardness, after grinding, in the range of 48–52 HRC. The St70AH ring geometry is 10 mm in height, with outside diameter 30 mm and inside diameter 15 mm. Testing v temperatures are in the range 1050–1200 C. The identified friction factor is 0.35. 2.3. Microstructural analysis
Fig. 2. St70AH sensitivity to strain rate history effects during deformation.
The maximum amount of d-ferrite phase considered acceptable in an St70AH blade corresponds to 5%; for higher percentage of d-ferrite phase, mechanical strength and ductility together with fatigue resistance of the alloy drastically decrease [18]. To determine the actual evolution of d-ferrite phase within the range of testing temperatures and strain rates, the microstructure of the specimens deformed
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Fig. 3.
P.F. Bariani et al. / International Journal of Machine Tools & Manufacture 44 (2004) 945–951
Geometry of the simulative test for heat transfer evaluation and locations of thermocouples embedded into the upper and lower die.
according to Table 2 was observed and the amount of d-ferrite phase evaluated. The specimens were cut into transverse sections, polished to 1 lm and then etched by electrolysis to reveal the microstructure. For each specimen, the microstructure was observed in three different zones that are represented in Fig. 4. The influence of both testing temperature and strain rate on the amount of d-ferrite phase was evaluated: for all the testing temperatures, it is proved that strain rate does not influence the microstructure in a significative way. As example, in Fig. 5, St70AH microv structures at T ¼ 1100 C at different strain rate are shown. Temperature, instead, strongly affects the d-ferrite phase formation (Fig. 6 on the left), leading to an v unacceptable microstructure at T ¼ 1250 C, characterised by more than 5% of d-ferrite. Fig. 6 on the right shows also the influence of soaking time just before deformation: if the specimen is held at the testing temperature for 20 min (instead of 30 s), its microstructure results almost unchanged for temperatures v v less than T ¼ 1180 C, while at T ¼ 1250 C the amount of d-ferrite phase drastically increases. On the basis of results of microstructural observations, the control of temperature inside the workpiece allows to control the amount of d-ferrite phase. If temperature inside the workpiece during deformation is v kept lower than 1180 C, then the blade microstructure can be considered acceptable.
Fig. 4. Zones of microstructural analysis on the deformed specimen.
3. Results of FE simulations By using the calibrated FE model, the temperature distribution inside the airfoil region was calculated at the end of each forging blow. The maximum value of temperature was compared with the limit value that corresponds to the critical amount of d-ferrite phase. Particular attention was paid to temperature increase in the central zone and along the middle line of the airfoil section. Temperature of the material close to the interface with the dies and in the flash was not taken into account since these portions of material are removed by machining after forging. Fig. 7 shows the temperature distribution inside the cross-section at the end of the second (left) and the third steps (right). The highest values of temperature in the central zone v of the airfoil region are always lower than 1100 C at the end of the relavant forging steps, particularly at the end of the second step, which is the most significant as temperature increase in the central zone of the airfoil. From the microstructural stand-point, the amount of d-ferrite phase will be within the acceptable limits.
4. The optimised forging process The accurate analysis of the temperatures in the airfoil at the end of the second and the third forging steps revealed that the temperature and the corresponding amount of d-ferrite phase are rather lower than the respective critical values. The lack of a quantitative knowledge of thermal and microstructural events during the forging process induces the process planner to be too conservative. Consequently, the design of the current industrial process proves to be quite inefficient. The quite high number of forging blows causes long lead times to process the batch and multiplication of the costly operations to carry out at the end of each blow. It becomes than clear that the total number of forging blows could be reduced, increasing the initial billet temperature and, at the same time, maintaining the maximum amount of d-ferrite phase lower than 5%.
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Fig. 5.
Fig. 6.
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v
Microstructure sensitivity to strain rate at T ¼ 1100 C.
Microstructure sensitivity to temperature at strain rate equal to 50 1/s and with different soaking times before deformation.
Thanks to St70AH high temperature sensitivity (Fig. 1), temperature increase leads to a substantial reduction of the total forging load, thus reducing the overall production costs.
The actual forging process was re-designed reducing the number of forging steps from three to two: the initial v v billet temperature was increased to 1150 C (50 C more than the actual forging process).
Fig. 7. Temperature distribution at the end of the second (left) and the third (right) steps of the actual process.
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Fig. 8.
Temperature distribution at the end of the second (left) and the third (right) steps of the optimised process.
transformation. In the case of St70AH, accurate prediction of temperature and d-ferrite phase distributions has allowed the reduction of the number of forging steps of the actual process from three to two, increasing the overall initial process temperature and thus leading to a more effective forging process. Industrial trials have demonstrated the effectiveness of this forging process optimisation.
Acknowledgements Fig. 9. Example of blade microstructure in the central zone of the airfoil after heat treatment.
As shown in Fig. 8, the values of temperature calculated in the cross-section at the end of the second blow are almost everywhere lower than the critical one corresponding to unacceptable amount of d-ferrite phase. v Only in the flash area temperature reaches 1180 C, corresponding to an amount of d-ferrite phase greater than 5%. 4.1. Industrial validation Fig. 9 shows the airfoil microstructure of an St70AH blade, forged with the optimised value of initial temperature. The microstructure is relevant to the blade after heat treatment: however, it does not present any evident tracks of d-ferrite phase (note that d-ferrite phase after formation at high temperatures remains in the component even if heat treated).
5. Concluding remarks Accurate knowledge of the thermal and mechanical events determining the evolution of microstructure during deformation phases is a primary prerequisite for design and optimisation of forging St70AH turbine blades. The combination of an FE model calibrated to the specific forging problem with a systematic microstructure analysis on samples prepared through physical simulation experiments proved to be an effective and successful tool in predicting temperatures and phase
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