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journa l homepage: www.e lsev ier .com/ locate / jmatprotec

D FEM simulation of the multi-stage forging processf a gas turbine compressor blade

heng Lva, Liwen Zhanga,∗, Zhengjun Mua, Qingan Taib, Quying Zhengb

School of Materials Science and Engineering, Dalian University of Technology, Dalian 116024, PR ChinaShenyang Liming Aero-Engine Group Corporation, Shenyang 110043, PR China

r t i c l e i n f o

rticle history:

eceived 2 December 2005

eceived in revised form

2 April 2007

ccepted 25 July 2007

eywords:

as turbine compressor blade

a b s t r a c t

Due to the complicated three-dimensional geometry and the non-steady state contact

between the workpiece and the die surface, the simulation of blade forging process

performed so far has been restricted to two-dimensional plane-strain problems or sim-

plified three-dimensional deformational cases throughout which some simplifications and

assumptions are employed. This study attempts to simulate an entire forging process of a

gas turbine compressor blade from a cylindrical billet to a complicated product, using 3D

rigid-viscoplastic FEM. Simulation successfully predicts a complete load/time diagram and

deformed configurations on the preforming stages and the following forging stage. Mean-

while, the distribution of different field-variables, such as strain and temperature, were

ulti-stage forging

D rigid-viscoplastic FEM

obtained. On the basis of these results, a change of the original forging stage is recom-

mended. The validity of simulation results was verified through comparisons with industrial

trials, which were conducted on the same process parameters as those in the simulation. The

simulation results may be effectively applied to other types of three-dimensional turbine

blade forging processes.

of its twist shape from the root to the end and uneven body at

. Introduction

he compressor blade is one of the most important mechan-cal components in gas turbine engine, which plays anmportant role in energy transformation, and it requires high-eometrical precision and mechanical properties due to itsevere working conditions. Forging of compressor blade is aomplex operation to describe quantitatively due to the sen-itivity of the properties of the material to process conditionsnd the complicated shapes of the products, which have awist shape from the root to the end of a blade. To enable the

anufacture of compressor blade with suitable mechanical

roperties and correct shape to be undertaken in a scientificanner, it is necessary to understand material flow, strain,

train rate, forging load and temperature histories during the

∗ Corresponding author. Tel.: +86 411 84706087; fax: +86 411 84708116.E-mail address: commat@student.dlut.edu.cn (L. Zhang).

924-0136/$ – see front matter © 2007 Elsevier B.V. All rights reserved.oi:10.1016/j.jmatprotec.2007.07.032

© 2007 Elsevier B.V. All rights reserved.

deformation process. To achieve the above goals, the finite ele-ment method can be used for its ability to account for thecomplex tool/workpiece interactions and boundary conditionsthat would occur during the manufacturing processes.

Until now, in applying FEM to gain an understanding ofthe thermomechanical characteristics of blade forging, mostof the simulation work performed has been treated as two-dimensional plane-strain problems (Morita et al., 1991; Dungand Mahrenholtz, 1982; Kang et al., 1990; Soltsni et al., 1994),which inevitably suffer from a lack of practical significance.However, blade forging is a three-dimensional process because

different positions. Therefore, a full three-dimensional sim-ulation is required in order to study the material flow in allthe regions and obtain more realistic information to improve

n g t

464 j o u r n a l o f m a t e r i a l s p r o c e s s i

the precision of die design. In an early exercise with a sim-ple geometry, Argyris et al. (1985) analyzed the blade forgingprocess using the three-dimensional finite element methodby considering the thermal effect, but interface friction wasomitted. Isothermal forging of a turbine blade was analyzedby Yang et al. (1993) using 3D rigid-viscoplastic FEM. In thesimulation, the interface friction was included and remeshingwas carried out using a modular remeshing scheme. However,the arc transition between the tenon and the body on the diecavity was simplified with a right angle, which is far from theactual situation in the blade forging process. Zhan et al. (Zhanet al., 1999, 2001; Yang et al., 2002; Liu et al., 2002) simulatedprecision forging process of a compressor blade by 3D FEM, butthe preforming stages were not considered in the analysis.

In the present work, a three-dimensional finite elementanalysis of non-isothermal forging process of a gas turbinecompressor blade is carried out, using 3D rigid-viscoplasticFEM. This simulation includes all the deforming stages fromthe initial cylindrical billet to the final blade and considers thefrictional condition and practical arc transition between thetenon and the body. The deformational characteristics of theblade forging process are revealed and an optimized forgingscheme is recommended. As an experimental validation ofthe simulation results, forging trials of the compressor bladehave been carried out under the same forging conditions asthose in the simulation.

2. Simulation details

In the present work, the DEFORM 3D software package basedon an updated Lagrangian description was employed to simu-late the blade forging process. To allow the focus to be placedon the thermomechanical effects on the workpiece, a rigid-viscoplastic material formulation coupled with a heat transferformulation was used for the workpiece. The governing equa-tions that have to be satisfied during the forging process areas follows:

• Equilibrium condition:

�ij,i = 0 (1)

• Compatibility condition:

εij =(

12

)(vi,j + vj,i) (2)

• Constitutive relation:

εij =(

3˙ε2�

)�′

ij (3)

• Incompressibility condition:

˙εkk = 0 (4)

• Boundary condition:

�ijnij = Fi on Sf, vi = vi on Sv (5)

e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470

The equations given above can be solved by a variationalprinciple expressed as

ı˘ =∫

V

�ı ˙ε dV + k

∫V

εVıεV dV −∫

Sf

Fiıvi ds = 0 (6)

where

� =√

3/2(�′ij�

′ij)

1/2, ˙ε =

√2/3(εijεij)

1/2 (7)

In the above equations, �, ˙ε, εV and �′ij

are the effective stress,effective strain rate, volumetric strain rate and deviatorystress components, respectively. V the volume of the billet,Sf the force surface, Sv the velocity surface, Fi the tractionstress and k is the large positive constant to penalize volumechange.

Eq. (6) can be converted into non-linear algebraic equationsby utilizing the standard FEM discretization procedure. Due tothe non-linearity involved in the material properties and fric-tional contact conditions, the solution is obtained iteratively.

This rigid-viscoplastic material model is coupled with aheat transfer model, expressed by the energy-balance equa-tion:

(kT,i),i + r − (�cPT) = 0 (8)

where k denotes thermal conductivity, T the temperature, r theheat generation rate, � the specific density and cP is the specificheat. The first-term (kT,i),i and the third-term �cPT representthe heat transfer rate and the internal energy rate, respec-tively. The rate of the heat generation in the deforming billetdue to plastic deformation is given below:

r = ˛� ˙ε (9)

where ˛ represents the fraction of mechanical energy con-verted to heat, usually assumed to be 0.9. The temperaturedistribution of the workpiece and/or dies can be obtained read-ily by solving the energy balance equation rewritten, by usingthe weighted residual method, as

∫V

kT,iıT,i dV +∫

V

�cPTıT dV −∫

V

˛� ˙εıT dV =∫

S

qnıT ds (10)

where qn is the heat flux normal to the boundary surface,including heat loss to the environment and friction heatbetween two contacting objects. By applying the FEM dis-cretization procedure, Eq. (10) can also be converted intoa system of algebraic equations and solved by a standardmethod. In practice, the solutions of mechanical and thermalproblems are coupled in a staggered manner.

In the blade forging process, there are four stages dur-ing deforming, namely upsetting, heading, busting, and finalforging stage, illustrated in Fig. 1. Through the first three pre-forming operations, a cylindrical billet is formed to a preformwith a head. After preforming, the workpiece is cooled in air

to room temperature, the surface is descaled, blanks are thensand blasted and reheated up to the forging temperature witha soaking time to uniform the thermal distribution inside theentire material volume. Then, the forging operation is con-

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470 465

Fig. 1 – Forming stages of the blade manufacturing process: (a) initial billet; (b) upsetting preform; (c) heading preform; (d)busting preform; (e) final forging.

ture

dasp

wbrbt

�

Iicu

tion in the process, elastic deformation is negligible and allthe dies are considered as rigid bodies. The friction at theworkpiece–tooling interfaces was assumed to be of shear type,

Table 1 – Process conditions for FE-simulation

Preforming stages Forging stage

Initial workpiecetemperature (◦C)

1160 1160

Fig. 2 – Flow stress curves under different tempera

ucted to forming the final product. In order to understandnd analyze the actual blade forging process, all stages areimulated, including the heat transfer from furnace to diesrior to forging.

The material for the billet was a new stainless steel, whichas developed for the production of gas turbine compressorlade. During hot deformation of the workpiece, strain, strainate and temperature have a great influence on the flow andehavior of the material, which can be expressed as the equa-ion:

¯ = �(ε, ˙ε, T) (11)

n this paper, the material flow behavior can be realized bynputting the flow stress data, gained in the thermomechani-al simulation experiments. Fig. 2 shows the flow stress curvesnder different temperature.

: (a) 950 ◦C, (b) 1000 ◦C, (b) 1100 ◦C and (d) 1150 ◦C.

The process conditions in the FE-simulation are givenin Table 1. Due to high temperature and large deforma-

Tool temperature (◦C) 20 300Environment

temperature (◦C)20 20

Friction factor 0.7 0.3

466 j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470

Fig. 3 – FEM simulation for preforming operations: (a) initial billet; (b) upsetting preform; (c) heading preform; (d) bustingperform.

Table 2 – Thermal physical properties

Thermal conductivity (N/(s ◦C)) 36.5Heat capacity (N/(mm2 ◦C)) 7.74Heat transfer coefficient between

workpiece and die (N/(s mm ◦C))11

Convection coefficient toenvironment (N/(s mm ◦C))

0.02

Emissivity 0.8

Fig. 4 – Flow net pattern at various stages: (a) initial billet;(b) upsetting preform; (c) heading preform; (d) busting

Fig. 3 illustrates the stages of deformation during the pre-forming process. Due to the large deformation occurring inthe vicinity of head, the FEM mesh degenerate severely dur-

perform.

expressed as

fS = mk (12)

where fS is the frictional stress, k the shear yield stress and mis the friction factor. In the preforming operations, a frictionfactor of 0.7 is applied to model the dry forging conditions.While, the friction factor is assumed to be 0.3 in the followingforging stage because of the lubricated conditions between theforging die and workpiece. During forging, the upper die speedchanges with the ram movement of mechanical press, whichcan significantly affect the average strain rates and thereforestiffness within the workpiece. To model this change in com-pression rate, a ram-dependent die speed function is enteredinto the program for the preforming and forging simulations.A mechanical press with an effective load capacity of 25 MN isused in the forging operation. And the press used to preform

the billet is an upsetting machine with the load capacity of5 MN. The thermal physical properties of workpiece are givenin Table 2.

Fig. 5 – Temperature distribution inside the section B–B atthe end of preforming operations.

3. Results and discussion

3.1. Analysis of preforming operations

Fig. 6 – Load–time curves (preforming operations).

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470 467

Fig. 7 – FEM simulation for forging operation at a height reduction of: (a) 0.0 mm; (b) 29.7 mm; (c) 37.7 mm; (d) 40.1 mm.

F and

iftppflhticsi

bdcpbaWlrmsrdomte11

ppu

range of 1.33–1.61 in most parts of blade body. Relatively lowvalues of equivalent strain exist in the tenon, with the range of0.578–0.756 in most part of this region. In blade forging process,the critical deformation is an important factor, which deter-

ig. 8 – Equivalent strain distribution inside the section C–C

ng simulation. Therefore, remesh procedure is carried outrequently to complete the simulation. One of the most impor-ant information in metal forming analysis is the metal flowattern. However, FEM meshs are not adequate for this pur-ose when remeshing is involved. In order to visualize metalow pattern through remeshing, a procedure called flow netas been developed. Fig. 4 shows the flow net pattern of sec-ion A–A at different stages during the preforming process. Its clear that high strain caused by large deformation is con-entrated in the head. This concentrated distribution of hightrain leads the head material to be hardened and temperaturen this region is higher than others.

For the stainless steel utilized in manufacturing turbinelades, the temperature has to be kept within narrow rangesue to material strict workability windows as well as closeontrol in microstructure specifications of the final com-onent. High temperature in workpiece may induce somerittle phases, which can weaken the mechanical strengthnd the ductility of the component during its service life.hen the temperature in workpiece is lower, the forging

oad may increase abruptly due to the sensitivity of mate-ial flow strength to temperature, and internal damages of

aterials like as micro-crack may occur. For the stainlessteel reported in this paper, the temperature after forging isequired to be larger than 950 ◦C. Fig. 5 shows the temperatureistribution inside the section B–B at the end of preformingperations. As shown in the figure, the temperature in theajority region of the perform is 1140 ◦C, which is higher

han 950 ◦C. And due to the heat generation from deformationnergy and friction, there is a slight temperature rise of about0 ◦C in the head region compared to the initial temperature of160 ◦C.

Fig. 6 shows the load–time curves on each operation for thereforming process. It can be seen that the maximum load inreforming process is about 700 kN, which takes place in thepsetting operation because of high resistance of metal flow.

D–D at the end of forging operation: (a) C–C and (b) D–D.

The computered forming load does not exceed the effectiveload capacity of the upsetting machine, 5 MN.

3.2. Analysis of forging operation

The deformed shapes of the blade at different stroke of forgingoperation are shown in Fig. 7. It can be seen that the spread ofthe blade sections resulting from height reduction takes placeduring the process. The elongation of the blade is small andthe spread of the blade sections is almost straight so that theuse of transverse sections to study the deformed characteris-tics would be reasonable. The equivalent strain distributionsinside the cross-section C–C and D–D at the end of forgingoperation are shown in Fig. 8. It can be seen that large defor-mation takes place in the blade body, especially in the flashlands. The values of equivalent strain are found to be in the

Fig. 9 – Thermal plasticity curve of blade steel.

468 j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470

and D–D at the end of forging operation: (a) C–C and (b) D–D.

Fig. 10 – Temperature distribution inside the section C–C

mines the degree of deformation in single step. Fig. 9 showsthe thermal plasticity curve of blade steel, which was testedby thermal compression experiments. As shown in this figure,the deformation limit in every heat number is about 0.7. Thus,the equivalent strain in blade body is much higher, exceedingthe material plastic forming limit, which may induce someforming defects.

As mentioned above, high temperature in workpiece mayinduce some brittle phases, which can weaken the mechanicalstrength and the ductility of the product during its service life.Therefore, the temperature distribution in workpiece is mostsignificant during forging operation. Fig. 10 shows the temper-ature distribution inside the cross-section C–C and D–D at theend of forging operation. It is clear that there is little declineof temperature in blade body due to the heat generation fromdeformation and friction. The highest values of temperaturein the central zone of cross-sections are about 1170 ◦C at theend of forging stage. Based on physical modeling experiments,we find that this temperature value is much higher and themechanical strength and the ductility of the final forging part,which is forged under this temperature, are much lower for itsapplication.

Fig. 11 shows the load–time curve for the forging operation.Here, the forging load increases abruptly due to flash forma-tion. Maximum load for this finish operation is lower than20 MN, less than the effective load capacity of the mechanicalpress, 25 MN.

3.3. Optimization of forging operation

The accurate analysis of above forging operation revealsthat the temperature and equivalent strain values after final

Fig. 12 – Distribution of field-variables inside section C–C at the e

Fig. 13 – Equivalent strain distribution inside section C–C an

Fig. 11 – Load–time curve of forging operation.

forging stage are rather higher than the respective values. Con-sequently, the forging process should be amended. It is clearthat the large deformation could be reduced through adding apreforging blow prior to finish forging and, at the same time,reducing the finish forging temperature to decrease the pos-sibility of brittle phase occurrence. Thus, current industrialforging process consists of two forging steps, namely preforg-ing and finish forging. And the finish forging temperature isreduced to 1120 ◦C.

Fig. 12 shows the distribution of field-variables inside sec-tion C–C at the end of preforging operation. The values ofequivalent strain are found to be in the range of 0.422–1.07 inthe cross-section. And the highest temperature values in the

central zone of cross-section are about 1170 ◦C. These temper-ature and strain values are a little higher, but do not affect thefinal forging part due to the reheating and finish forging afterpreforging stage.

nd of preforging: (a) equivalent strain and (b) temperature.

d D–D at the end of finish forging: (a) C–C and (b) D–D.

j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470 469

and D–D at the end of finish forging: (a) C–C and (b) D–D.

ttgzwFattiaatm

ppdfls

Fp

Fig. 14 – Temperature distribution inside section C–C

Fig. 13 shows the equivalent strain distribution inside sec-ion C–C and D–D at the end of finish forging step. As shown inhis figure, the equivalent strain after finish forging is reducedreatly. The values are in the range of 0.3–0.8 in the centralone of section C–C and less values exit in the section D–D,hich are rather lower than the material plastic forming limit.

ig. 14 shows the temperature distribution inside section C–Cnd D–D at the end of finish forging. It can be seen that lit-le temperature change has happened compared to the initialemperature of 1120 ◦C, which could ensure the required qual-ties of the final products. There is a little temperature rise ofbout 20 ◦C in the flash lands due to the severe deformationnd friction in that region. However, temperature change inhe flash was not taken into account since these portions of

aterials are removed by machining after forging.Fig. 15 shows the load–time curves of optimized forging

rocess. It can be seen that the forging load is very low in

reforging stage and increases abruptly due to flash formationuring the finish forging stage. Thanks to high sensitivity ofow strength to temperature, temperature reduction leads to aubstantial increase of finish forging load. The maximum load

ig. 16 – Products through the optimized process: (a) preformingroduct.

Fig. 15 – Load–time curves of optimized forging process.

in finish forging stage is 23 MN, which is much higher than themaximum load during the original forging operation, but it islower than the effective load capacity of the mechanical pressall the time.

billet; (b) preforging part; (c) finish forging part; (d) final

n g t

r

470 j o u r n a l o f m a t e r i a l s p r o c e s s i

4. Experimental verification

After convinced by numerical simulation that the optimizedforging process can give a realistic possibility for the pro-duction of gas turbine compressor blade, forging tests wereperformed to validate the results of FE simulation. The forg-ing process parameters are the same as those in the optimizedprocess. The lubricant used in the experiments was the mix-ture of commercial grease and MoS2. The forging tests wereconducted on a mechanical press with 25 MN load capacityfor preforging and finish forging operations. And the pressused to preform the billet was an upsetting machine with theload capacity of 5 MN. Fig. 16 shows the geometries of a forgedblade and its preform produced with the optimized process.Compared with the simulated results shown in Figs. 3 and 7,the similarity of the geometry between the forging trials andthe simulation results is clearly apparent. Meanwhile, themechanical properties such as strength and ductility satisfythe application requirements well, according to the testingexperiments. Thus, industrial trials show good agreementswith the FEM simulation of the blade forming process.

5. Conclusions

Using the rigid-viscoplastic finite element method, a three-dimensional finite element analysis of non-isothermal forgingof a gas turbine compressor blade is carried out, includingall the deforming stages from the initial cylindrical billet tothe final blade. FE analysis provides detailed information onthe material flow, load, strain and temperature, which can beincorporated into process design. The conclusions with thefinite element analysis can be summarized as follows:

(1) At the end of preforming operations, the temperature inmost parts of the workpiece is 1140 ◦C, which is higherthan the temperature corresponding to unacceptable lowtemperature. At the same time, the loads during the pre-forming operations are much lower than the effective loadcapacity of upsetting machine.

(2) At the end of original forging operation, the values ofequivalent strain in most parts of blade are rather higher,in the range of 1.33–1.61. This highly accumulated strainexceeds the material plastic forming limit and may causesome forming deficiencies. Due to heat generation fromdeformation energy and friction, the temperature in work-

piece after forging operation is much higher than thereceivable value and may induce some brittle phaseswhich can weaken the mechanical strength of the com-ponent during its service life.

e c h n o l o g y 1 9 8 ( 2 0 0 8 ) 463–470

(3) On the basis of the analysis results of above forging oper-ation, an optimized forging scheme is recommended toeliminate the strain accumulation and reduce the tem-perature rise in the workpiece after finish forging. In theoptimized forging process, the simulation results showthat the equivalent strain and temperature are reducedgreatly and can ensure the required quality of the finalproducts. Industrial trials have demonstrated the validityof simulation results.

Acknowledgement

The authors would like to express their appreciation for thefinancial support of the National High Technology Devel-opment Program of China (2004AA503010) for the presentresearch work.

e f e r e n c e s

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Liu, Y.L., Yang, H., Zhan, M., 2002. A study of the influence of thefriction conditions on the forging process of a blade with atenon. J. Mater. Process Technol. 123, 42–46.

Morita, A., Hattori, S., Tani, K., Takemura, A., Ashida, Y., 1991.Near net shape forging of titanium alloy turbine blade. ISIJ Int.31 (8), 824–833.

Soltsni, B., Mattiasson, K., Samuelsson, A., 1994. Implicit anddynamic explicit solutions of blade forging using the finiteelement method. J. Mater. Process Technol. 45, 69–74.

Yang, D.Y., Lee, N.K., Yoon, J.H., 1993. A three-dimensionalsimulation of isothermal compressor blade forging by therigid-viscoplastic finite element method. J. Mater. Eng.Perform. 2 (1), 119–124.

Yang, H., Zhan, M., Liu, Y.L., 2002. A 3D rigid-viscoplastic FEMsimulation of the isothermal precision forging of a blade witha damper platform. J. Mater. Process Technol. 122, 45–50.

Zhan, M., Liu, Y.L., Yang, H., 1999. Research on a new remeshing

for the 3D FEM simulation of blade forging. J. Mater. ProcessTechnol. 94, 231–234.

Zhan, M., Liu, Y.L., Yang, H., 2001. A 3D rigid-viscoplastic FEMsimulation of compressor blade isothermal forging. J. Mater.Process Technol. 117, 56–61.