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Recent developments on high precision forging of aluminum and steel
Journal
of
71 (1997) Yl- Y9
Materials Processing Technology
_
--
Abstract Precision forming is a suitable means to meet the increasing demands for the cost-effective prediction of net-shape or near-net-shape parts in large quantities. This article demonstrates the different procedures applied to the precision forming of parts dependent on the specific critical aspects of the process. Examples are given for the geometrical and thermal layout of the forging process, the influence of material behavior on the forged part and a procedure for the compensation of the die deflection. A suitable layout of the forging sequences is the precondition for further optimization steps of the forging parameters. CAE techniques provide state-of-the-art simulation and reproducible manufacturing tools for precision forging. Computer based engineering development systems integrate design and simulation tools to optimize the forging process. The application of these tools increases the basic understanding for the net-shape production of complex shaped parts. Beneath the optimization of the forging process layout the attainable precision also depends on the accuracy of toolmaking and Drxess control. During the forging process the control of forging load and thermal conditions, including thermal treatment. k necessary to establish reproducible conditions for net-shape manufacturing. ‘_ 1997 Elsevier Science S.A. TJ Kqwwdv:
High precision forging; Aluminium: Steel
1. Introduction The increasing worldwide interest in the production of net-shape or near-net-shape products lead to the develornent of improved cold and warm forging tcchnologies. Warm forging is applied mainly to reduce the flow stress of the zzaterial, the forging loads and to increase the complexity of the part. The cold forging process avoids the problems connected with thermal distortion and surface degradation and made the production of ready-to-assembly parts possible. For automotive applications, cold forged steel parts are produced for power train and steering components in increasing quantities
PA. The search for weight reduction leads to an increasing use of light alloys like aluminum and magnesium, especially in the automotive industry. The application of new materials with improved properties is often restricted by increased material costs and production aspects. Hence, precision forging techniques is a promising means for material saving and the cost-effec0924-0136/97/$17.00 0 1997 Elsevier Science S.A. All rights reserved. Plt SO924-0136(97)00153-2
tive production of parts with improved properties at a competitive price. The reproducible production of nearnet-shape and net-shape forged parts in !arge quantities requires optimal process layout, precision tool manufacturing and the control of the forming parameters. Since material flow, tool life and forging loads are determined by the layout of the forging stages, the development of the forging process represents one of the most important steps [3,4]. The use of finite-element based computer aided engineering techniques is increasingly being applied in the development of forming processes [5]. With increasing computer performance, the three dimensional simulation of complex forming operations becomes feasible within reasonable time. Combined with state-of-the-art CAD-CAM techniques the precision and efficiency of tool design and tool manufacturing can be increased [6,7]. Expert systems have been developed for specific problems to support the designer during the layout ot the forming operations [S]. The future goal is the reduction to the minimum of the expensive ‘trial and error’ method using different tool designs. In the following, a
K. Siqyr t et (11./ Jowd
92
q/‘ Muterids
Processing
Tchrology
71 (1997)
91- 99
Fig. 3. Cold forging steps for drag links. Fig. 1. Slug and cold forged tripod CVJ inner race.
survey over recent developments of the precision forging of aluminum and steel components is presented.
2. Cold forging of a tripod constant velocity joint inner race ‘i’ke constant velocity joint (CVJ) is a commonly used part of the power train for modern front driven compact cars. The forming of a tripod CVJ inner race represents a nearly classical processing technique for a one-step cold forging operation with lateral material flow, Fig. 1. The lateral extrusion of the CVJ inner race is similar to the production of spiders for steering systems that represents an exceptional high developed and cost-effective precision forging for net-shape parts [2]. The part is designed mainly of two conical halves with three circumferential located radial trunnions, Fig. 2. The central body exhibits a conical recess at the top and a cylindrical recess at the bottom for further cutting operations. This part demonstrates how the design of the part affects the material flow, quality, tool wear and fatigue [3]. The tripod inner race was cold forged from cut Design“6”
,
10
2.9
y
rml
‘corm /
Shouldi
Fig. 2. Design variations of tripod CVJ inner race.
slugs of a medium-alloyed cold working steel 20MnCr4 (similar to AISI 5120) in an oil-gas spring supported horizontal split die. This tool provides a simultaneous axial mat xial flow during the forging operation. The accuracy of the part’s symmetry plane can be held to a reproducible tolerance of + 0.15 mm (0.0006 in.) to the top and bottomn surfaces. The design ‘B’ in Fig. 2 represents a design variation where a spherical plane is added at the transition between the trunnions and the double-conical central body of the CVJ inner race. For the parts with design ‘A’ of Fig. 2. the radial material flow into the cavity is: ua;iform and the surface of the trunnions is without any defects. While forming the inner race with design ‘B’ of Fig. 2, the lateral extruded material looses contact with the cavity shoulders which produces folding laps at the upper surface of the trunnions. These faults may cause some additional material allowance for grinding operations.
3. Cold forging of drag links The forging process of drag links demonstrates the necessity of an accurate layout of the forging stages for precision near-net-shape forging. For this part, cost savings of about 10% could be realized by precision cold forging instead of former hot forging. This part is produced by five cold forging operations in a multi-stage press from an upset wire slug of AISI 1035 steel, Fig. 3. The design of the forming stages is decisive for the productivity of the forging process of this part. For optimal material flow and complete filling of the cavity, the forming process had to be arranged as a five-step cold forging sequence. Fig. 4 illustrates the layout of the forming stages from the preformed slug to the final forging operation. For all five forming steps an oil-gas spring supported horizontal split die is required. The subsequent forging operations are strongly influenced by the preform design and material distribution.
Upsetting
Cup Extwsion
Piercing/ Rod Extrusion
Cutting Load
Direction
Fig. 4. Forming
oprrations
tbr t!le cold forming
Production aspects and material flow considerations lead to the bail-like design of the preform head with differenc radii. This design prevents the complete tilling of the cavity already at the beginning of the first forging process. The uniform radial material flow provides a simultaneous die contact in the cavity that leads to reduced die loads compared with a spherical geometry. Before forming, the preform is annealed, zinc phosphated and coated with molybdenum-disuifide or soap. During the first forming stage the ball-like head of the ,eform is formed into an cylindrical shape and the angular transition between shaft and head is formed. The height of the final formed part, symbol ‘A’ in Fig. 4, is determined already at the first forming step. Further material Aow is prevented because of the perpendicular orientation of the shaft towards the forming direction. In order to avoid overfilling at the next forging stage, the upper side of the cylinder, symbol ‘B’ in Fig. 4, is inclined by 10” towards the horizontal axis. At the bottom of the head a bail shaped element is formed to improve the material distribution and to support the material flow. With the second forging operation a conical cup is formed at the upper side of the drag link head, symbol ‘D’ in Fig. 4. At the bottom an oval recess is formed, ‘C’ in Fig. 4, because the shape of the final part exhibits an oval hole in this area. In order to prevent folds. the punch cross-section has to be larger than the spherical cavity. This ensures lateral material flow and increase of the surface.
or dr;q
links.
The third forging stage combines a backward cup extrusion and a simultaneous forward rod extrusion. The backward cup extrusion transforms the conical cup into an oval shape. With the simultaneous rod extrusion process surplus material is displaced. A strong reduction of the forging loads is the second effect of the oval force relieving hole at the bottom. In order to Facilitate the piercing of the bottom at 111,~ last stage without a die. the shape of the cup-bottom at the Iast stage without a die, the shape of the cup-bottom transition exhibits a weakened cross section. In the fourth and fifth stage the drag link is cut at the transition between shaft and cup. symbol ‘E’ in Fig. 4 and pierced at the bottom.
4. Optimization of the tooling for the cold forging of helical gears Several forging methods like cold [9.10] and warm forging [ 1l] have been investigated to produce near-netshape helical gears. The goal is the fabrication of ready-to-install gears without any further grinding operation. The lateral cold extrusion process was developed to improve the precision of forged helical gears and to reduce the production costs compared with conventional cutting operations. Cold forged gears provide a high surface quality, load appropriate grain orientation and the effect of work hardening during extrusion that
94
K. Siegert et al. /Journal
of Materials Processing Tecltttolog~~71 (1997) 91-99
Fig. 5. Schematic of the lateral extrusion process for helical gears.
produces increased fatigue strength values [2]. Fig. 5 shows a schematic of the lateral extrusion process for helical gears. The die consists of an oil-gas spring supported split die. At the beginning of the extrusion process the slug is supplied into the die. After the traveling ram has closed the die, the upper and lower punch transform the ram load into the slug so that the material is displaced into the confined toothed cavity. In the last stage, a build-in ejector lifts the lower punch and ejects the forged gear while the die is free to rotate due to the helix angle. Previous studies [9,10] demonstrated the necessity to compensate the die deflection that occurs during the cold forging of gears. For common use like gearings, the gears should have a tolerance of IT 7 and better. For applications with reduced precision demands IT 9 is recommended. With uncorrected dies tolerances between IT 10 and IT 12 are attainable due to: ?? heat generation during forming ?? elastic spring-back of the workpiece ?? distortion due to work hardening of the teeth ?? die deflection With the die deflection as the dominant parameter [12] for cold forging, any method to improve the precision of cold forged helical gears has to emphasize the evaluation and the control of the die load and die deformation. In [13] a numerical analysis with 2D and 3D FEM- and BEM-codes was employed on the lateral extrusion of gears. Further investigations lead to the development of a self-optimizing program system with automated precorrection and optimization of the die geometry (Fig. 6)
v41. In combination with a finite-element simulation program this tool calculates a modified die geometry where profile mismatch and elastic die deflection will be compensated while the forging loads are applied. The calculation of a deflection-compensated die is based on a 3D finite-element simulation with the FORGE 3 code foi
the calculation of the forging loads and stresses. Because the periodical structure of the entire workpiece permits a meaningful subdivision, the slug was modeled as a segment from a torus for the simulation of one single tooth. The die is represented by the tooth’s outer surface. Fig. 7 displays the calculated pressure values perpendicular to the die surface at the end of the lateral extrusion. During the lateral extrusion process the material fills the tooth cavity from the bottom to the top. At the end of the forging process the upper and lower corner of the tooth are filled while the calculated pressure values can rise up to 4000 MPa and more. Other authors [15] found with theoretical investigations that the contact pressure values are mainly influenced by the pressure angle and the pitch displacement. Because FORGE3 is not suitable for the calculation of the die deflection, this simulation was performed in a separate step with the FEM-code-PERMAS. The calculated nodal displacements of the die are applied as input data for the developed optimization program that calculates the correction for the appropriate die dimensions. First, this program converts the nodal displacements into geometrical deviations as defined in DIN standard 3960 and calculates the radial deflection of the die. In the following iteration steps the program calculates the corrected die profile based on the modification of helix angle, pitch diameter and base circle. Generally, the modified tooth profile exhibits a reduced helix angle due to the deformation of the tooth profile. The iteration stops when the calculated data is below a defined limit. With the calculated toothing parameters the modified die geometry can be designed with a 3D CAD-system and produced via spark erosion. Fig. 8 shows the measured deviations of lateral extruded gears before and after correction of the die geometry. The maximum die deflection with x0.35 mm (0.014 in.) occurs in the radial direction. With the
95
EM - Net Generator
HelicalGear
Relational Data Selection to store the Results
Selected Hvdraulie Prew Cokd Forging of Helical Gears
II
CAD-Svstem
bptirnizetion Modul Data Selection
Fig. 6. Development system for simulation and optimization of the lateral forging of helical gears.
developed optimization module, this radial die deflection can be compensated up to 95%. For an optimal die compensation the toothing deviations have to be considered additionally. Because the toothing of the die is produced by spark erosion, the tooth parameters cannot be varied along the vertical axis. With this limitatioi: a reduction of the toothing deviations up to 50% is attainable.
an
in N/mm2
5. Cold extrusion of P/M high Glicon con&es&s
al~jnu~
materials with
Hypereutectic aluminum-silicon alloys are commonly used for applications where high wear resistance and low thermal expansion coefficients are required. With conventional ingot (I,‘M) metallurgy techniques alloys with silicon contents up to a maximum of 25% can be produced. P/M processing routes, like the spray deposition process, allow the production of aluminum alloys with silicon contents up to 35% and more. These alloys exhibit a segregation free microstructure with a fine and uniform distribution of primary silicon crystals. For the production of thin-walled tubes with small tolerances from hot extruded high-silicon P/M grades, the cold extrusiou process was condidered to be the most effective forming technique. A technique to increase the formability of metal matrix composites (MMC) by means, of superimposed counter pressure is presented in [ 161.FOI manufacturing rea.sons, the cold extrusion of hypereutectic P/M tubes should Qe feasible with a single-action press. In this case, the material behavior during cold extrusion is the critical aspect that has to be considered. Like MMC, cracks are ini+$ted by the fracttrre of the embedded primary silicon p:u-titles. In order to optimize the forming process, the inflr;:ence of the following extrusion parameters on the materials microstructure was investigated: Equivalent strain. strain rate Max. straii, stat64 stress, strain diiributii &r.. &aiJl, sMe.of-Gress. straii distiibution .
Fig. 7. Cakulated
die contact pressure values from 3D FEM-simula-
tions at the end of the lateral extrusion (extended to three teeth).
sbain rate,fklw stress... Flaw stress. slrain hardening,famability
Tooth Profile Deviations -
“‘bmt
bmt
(b =
Radial Deviations
Angular Deflection
Profile Deformation
bmt
bmt
bmt
Root Circle
Crown Circle
Depth of Tooth
bmt
Bottom. m = Middle, t = Top of the Die Insert)
Fig. 8. h+zdsured profile deviations of cold forged helical gears.
Fig. 9 shows the workpieces from cut thick-walled tubes with outer diameters of 88 and 94 mm and the cold extruded thin-walled tubes. The investigated material Dispal S261X is based on an aluminum matrix with 25”/ silicon added to the melt before spray deposition. After spray deposition, the billet was extruded to thick walled tubes at 420°C with subsequent cooling at atmosphere (‘F’ condition). Fig. 10 shows the dimensions of the preform from cut P/M extrusions and the cold extruded tubes. The hot extruded thick-walled tubes were hollow forward extruded at room temperature in the ‘F’-condition and after soft-annealing. The parts extruded in the ‘F’-condition showed severe cracking of the primary silicon crystals whereas much less particle cracking occurred in the soft annealed extrudates. In order to optimize material flow and the state-ofstress during cold extrusion, simulations were performed with the finite element program DEFORM 2D. For a given temperature the material’s formability is mainly inf&enced by the state-of-stress, which is _I__-
represented by the quotient of mean stress G,,, and flow stress 6, [17]. Fig. I 1 illustrates the state-ofstress curve over the equivalent strain for several material points during the hollow forward extrusion of the tubes. For the most part of the extrusion process, the state-of-stress is characterized by large amounts of compressive stresses (negative values of g,,,/~,) which increase the material’s formability. For the materials points located at the bottom of the workpiece (symbol ‘e’ in Fig. 1I), the state-of-stress does not change significantly. For the other points (symbol ‘a’-‘d’ in Fig. 11) the state-of-stress drops significantly at the beginning of the extrusion process and increases with increasing strain. At the end of the deformation the material’s points are subjected to small amounts of tensile stresses (positive values of a,,,/~~). The state-of-stress is influenced by the die-openingangle 2a and the die fillet radius. With decreasing die-opening-angle, the state-of-stress becomes more fa-
1
Dlspal S281X (Al + 25%SI),
ColdExtid
Tube
2a
Dimension
88 doo=88mm
doo=94mm
(cpg= 1.22)
(qg = 1.52)
Fig. 9. Cut thick-walled tubes.
mm
preforms and cold extruded thin-walled Fig. IO. Dimensions of the preforms and the cold extruded tubes.
Al ; 25%S1 2ca=90” dao=94mm dal=74mm di=66mm
T=RT
2.5 Fig. 13. Production stages of ;I connecting rod forging. Fig. 1I. State-of-stress
during hollow forward extrusion of tubes.
vorable for the deformation to high strains. The quotient GJCJ, reaches higher negative values at the bcginning of the forming process and lower positive values at the end of the extrusion process. As the simulation results does not consider the characterisrics of the lubrecation system, the results may drive to the conclusion that a small die-opening-angle would be favorable. Extrusion tests with a die-opening-angle of 2r = 60” showed that the lubricant (zinc stearate) failed and severe welding of the workpiece material occurred on the die surface. Fig. 12 shows the microstructure of extruded tubes for different die fillet radii. Small fillet radii lead to a strong bending and shearing of the material while passing the die [18]. This produces increasing amounts df strain induced into the workpiece compared with a more rounded die and leads to an increased fracture of the silicon particles.
6. Forging of connecting rods Connecting rods were forged from extruded rods of a high silicon P/M aluminum alloy 1191. The alloy is
R = 0.5 mm
Fig. 12. Microstructure
R=2mm
of extruded tubes for different die lillet radii.
composed of a conventional AA-2024 matrix with additional 17% Si, 5% Fe and 0.6% Zr. A two-step w;~r~l~ forging process ~/as applied to overcome the limited formability of the P/M material at room temperature. Fig. 13 shows the different forging stages from the extrusions to the completely machined part. The forging process was arranged as a two stage sequence of closed die forging with flash and closed die forging without flash. Closed die flashless forging was chosen to improve the surface and the material properties of the rods. For the complete filling of the confined &e cavity an exact mass distribution of the preformed part is required. On the other hand, deticiencies of material lead to unfilled corners of the cavity. The preform shape was derived from the shape of the finally forged workpiecc by volume constant geometry variations to optimize the material flow. Closed die precision forging without Rash rcquirc~ enhanced accuracy and reduced tolerances in tool design. Compared to DIN and international standards. the side wall thickness (3 mm) and the base thickness (3 mm) of the connecting rods are designed smaller than for conventional forgings [20,2I]. General tolerances for precision forging are reduced to typically half of those for conventional forgings. Thickness tolerances are reduced to about two thirds of their former values. For precision forging of aluminum alloys in steel cavities the shrinkage of the forged and cooled down workpiece has to be considered. The different thermal expansion coefficients of workpiece and die cause different thermal extensions at elevated temperatures. Fig. 14 reveals the curves for a constant thermal expansion difference between workpiece and die in the thermal working area for the forging of the connecting rods. The cavity dimension, c were extended by 03’41 to reduce the shrinkage of the forged workpiece in the thermal working range. The forging tool, Fig. 15, was designed as an electric heated die with changeable die inserts for the first and the final forging stage. Closed loop temperature cOiltrOl systems and temperature measurement systems were used to monitor the temperatures of the dies.
K. Siegert et al. / Jotrrttal of’Materials Promsittg
98
Tedtttology 71 (1997) 91-99
wrM@aa~Dm~s2.m (Al l,si Sk3.Scu C.Ch!+cl.~ oii mlww 233s fX3ZCrMOV33.wmpnd) II
t*)r
I
=a30
r3.340%
.IT.i04lK
l-hWd~dikN=+ALl Am=m_.A&--.A&,, ” _._.._
+,$.
-0 -me
.I&
-
-560 ._._._
workpkce
Temperature
lncmsse
d&rp in
-sdm
K
Fig. 14. Temperature influence on thermal tween workpiece and die.
expansiondifferencebe-
CAD designed die inserts were CNC-milled of a prehardened tool steel, whereas high temperature steel X32 CrMoV 33 was used for the die inserts. For the first forging stage, the die inserts were designed with a symmetrical parting line and conventional [20] draft angles and transition radii. The tool for the fin,al forging stage consists of punch and die witil a la.teral tolerance of + 0.01 mm and draft angles of 0”. Two ejectors were placed in the die near the bearing areas of the connecting rod to release the forged workpiece. The preforms were closed die forged with flash from turned extrusions in a temperature controlled upper and lower die. After separating the flash the workpiece was precision forged in a closed die without flash. For preforging of the connecting rods, the workpiece temperature was varied between 300 and 500°C, die temperature was adjusted to 300°C. Graphite-water emulsion was used to reduce friction between workpiece and the cavity surface. The
Fig. 16. Height measurements ent forging temperatures.
for connecting rod preforms for differ-
For a combined forging and heat treatment process, the workpiece preheating temperature is determined from the annealing temperature of the material. Thus, the preheating temperature of the final forging stage was varied between 460 and 5OO”C,the die temperature was 450°C. Conventional furnace heating and also induction heating were used for preheating the initial and preforged workpieces. The volume of the finished part is directly determined by the volume of the preforged workpiece. Hence, the complete filling of the confined cavity requires a minimum volume of the workpiece. As a measure for the volume of the preforged workpiece, the height of the bearings was examined for different forging temperatures. Because of the elastic tool deformation, the height of the preforged parts is also affected by the flow stress of the workpiece. Fig. 16 shows the influence of tlhe workpiece temperature on the height measurements. for preforged connecting rod and the reference lines for the minimum heights required for complete filling of the cavity. Both curves decrease with increasing temperature. Between 380 and 460°C the influence of the workpiece temperature on the height measurements is reduced to 0.05 mm. For the finished forged connecting rods height measurements showed a scatter of about &-0.1 mm. At the end face of the rods a small vertical flash with a thickness of x 0.2 mm occurred. With a subsequent cold sizing operation a further reduction of the tolerances may be achieved. Quenching and arfificial aging of the forged connecting rods produces a significant increase in hardness of z 50% compared with the extruded material. Compared with conventional heat treatment the distortion of the workpiece produced by annealing and quenching can be reduced. References
Fig. IS. Connecting rod forging tool. Right side: final forging stage (without flash); and left side: first forging stage (with flash).
[I] M. Hirschvogel, Transmission shaft forgings-technical and economical aspects of new developments, Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May, 1995.
PI R. Geiger, M.
tinsel, From near net-shape to net-shape cold forging--state of the art. Proc. 9th Pm. Cold Forging Congr.. Solihull, UK, 22-26 May, 1995. [31 M. Kammerer. Th. Werle. K. Piihlandt, Entwicklung von Stadienpllnen zur Fertigong schwieriger Fliessressteile, in: K. Siegert (Ed.), Ncuere Entwicklungen in dcr Massivumformung 1995 DGM Informationsges. mbh, Oberursel. 1995. [41 M. Kammerer, Werkzeuge zum Fliesspressen von Aluminium, in: K. Siegert (Ed.), Fliesspressen von Aluminium, Vol. 1, DGM lnformationsges. mbh, Qberursel. 1995. 151 T. Alum, K. Sweeney, V. Vazquez, H. Kim. M. Knoerr, Cold forging of complex shaped parts to close tolerance-application of metal flow simulation to process and tool design, Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May, 1995. Konstruktion und [61 K. Siegert, R. Neher, Computerunterstiitzte Fetigung sowie massliche Kontrolle mit Datenriickfiihrung bei der Herstellung von Schmiedegesenken und ..teilen, in: K. Siegert (Ed.), Neuerc Entwicklungen in der Massivumformung 1991 DGM Informationsges. mbk, Oberursel. 1991. Hochgeschwindigkeitsfrasen von Neher, .I. Boxier. [71 R. Graphitelektroden. in: K. Siegert (Ed.), Rechneranwendung in der Umfonntechnik. DGM Informationsges. mbh, Oberursel, 1992. [El G. Du. Ein wissernbasiertes System zur Stadienplanermittlung beim Kaltmassivumformen. Berichte aus dem Institut fur Umformtechnik Nr. I 11. Springer-Verlag. Berlin, 1991. Stirn[91 F. Schmieder, Beitrag zur Fertigung von schragvenahnten radern durch Querpressen. Berichte aus dem lnstitut fiir Umformtechnik Nr. 118. Springer-Verlag. Berlin. I99 I. HOI V. Szentmihalyi, Beitrag der Prozesssimulation zur Entwicklung komplexer Umforrnteile. Berichte aus dem lnstitut fur Umformtechnik Nr. I2 I. Springer-Verlag, Berlin. 1994.
[I I] E. Doege. H. Nagele. FEM-gcstlutzte Auslrgung van Prazisionsschmiedepro/e+icn ~.rlnlornrtechnik 29 ( I ) ( I9951 ?? 10. [I’] M. Lamer. Untersuchungen ubrr d&s KaltHicssp~e~~rn geradund schragverzahnter Stirnrider. VDI Fortschrittsberichtc Rcihe 3 Nr. 221, VDI-Verlag. Dusseldorf, 1991. p, Ihl. [l3] K. Lange. V. Szentmihalyi. Optimized cold fot,ginp of helical gears by FEM-simulation. Proc. 9ih Inl. Cold Forging Congr.. Solihull, UK, 22-26 May. 1995. 1141 T. Kepplcr-Ott. Rechnerunterstiitzte Optimierung der Werkzeuggeometrie fur die Herstellung von gerad- und schragverzdhnlen Stirnradern durch Verfahren der Massivumformung, cnpublished report, 1996. [l5] F. Dohmann, N. Lintel, Konlaktspannungen beim Verzahnungspressen, Umformtechnik 29 (I) (19S5) 46-51. [I61 H.W. Wagner, J. Wolf. Kaltmassivumforrinmg von MMCs in Form der Matrixlegierungen AlMgSit und AlSil2 mit dem Ver-stiirkungswerkstoff SIC. Verbundwerkstoffe und Werkstoffverbunde, Symp. Deutschen Gesellschaft fiir Materialkunde, Chemnitz. 17-19 Juni. 1993. DGM Informtionsges. mbH. Oberursel. 1993, S.43350. [I 71 K. Siegert, D. Ringhand, Formability of dispersion strengthened P/M aluminium alloys. Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May. 1995. [I 81 K. Sicgert. K.-J. Fann. D. Engelmann. Simulation des Yoll-Vorwarts-Fliesspressens von Aluminium, in: K. Siegert (Ed.), Rechneranwendung in der Umformtechnik. DGM Informationsges. mbh. Oberursrl, 1992. [l9] K. Siegert. D. Ringhand. Flashless and precision forging of connecting rods from P,:M aluminum alloys. J. Mater. Proc. Technol. 36 (1994) 157-167. [ZO] DIN l?49, Gesenkschmiedestiicke aus Aluminium. 1211 H. Meyer-Nolkemper. Gesenkschmieden von Aluminiumwerkstoffen I-IX. Aluminium 55 (1979).
of
71 (1997) Yl- Y9
Materials Processing Technology
_
--
Abstract Precision forming is a suitable means to meet the increasing demands for the cost-effective prediction of net-shape or near-net-shape parts in large quantities. This article demonstrates the different procedures applied to the precision forming of parts dependent on the specific critical aspects of the process. Examples are given for the geometrical and thermal layout of the forging process, the influence of material behavior on the forged part and a procedure for the compensation of the die deflection. A suitable layout of the forging sequences is the precondition for further optimization steps of the forging parameters. CAE techniques provide state-of-the-art simulation and reproducible manufacturing tools for precision forging. Computer based engineering development systems integrate design and simulation tools to optimize the forging process. The application of these tools increases the basic understanding for the net-shape production of complex shaped parts. Beneath the optimization of the forging process layout the attainable precision also depends on the accuracy of toolmaking and Drxess control. During the forging process the control of forging load and thermal conditions, including thermal treatment. k necessary to establish reproducible conditions for net-shape manufacturing. ‘_ 1997 Elsevier Science S.A. TJ Kqwwdv:
High precision forging; Aluminium: Steel
1. Introduction The increasing worldwide interest in the production of net-shape or near-net-shape products lead to the develornent of improved cold and warm forging tcchnologies. Warm forging is applied mainly to reduce the flow stress of the zzaterial, the forging loads and to increase the complexity of the part. The cold forging process avoids the problems connected with thermal distortion and surface degradation and made the production of ready-to-assembly parts possible. For automotive applications, cold forged steel parts are produced for power train and steering components in increasing quantities
PA. The search for weight reduction leads to an increasing use of light alloys like aluminum and magnesium, especially in the automotive industry. The application of new materials with improved properties is often restricted by increased material costs and production aspects. Hence, precision forging techniques is a promising means for material saving and the cost-effec0924-0136/97/$17.00 0 1997 Elsevier Science S.A. All rights reserved. Plt SO924-0136(97)00153-2
tive production of parts with improved properties at a competitive price. The reproducible production of nearnet-shape and net-shape forged parts in !arge quantities requires optimal process layout, precision tool manufacturing and the control of the forming parameters. Since material flow, tool life and forging loads are determined by the layout of the forging stages, the development of the forging process represents one of the most important steps [3,4]. The use of finite-element based computer aided engineering techniques is increasingly being applied in the development of forming processes [5]. With increasing computer performance, the three dimensional simulation of complex forming operations becomes feasible within reasonable time. Combined with state-of-the-art CAD-CAM techniques the precision and efficiency of tool design and tool manufacturing can be increased [6,7]. Expert systems have been developed for specific problems to support the designer during the layout ot the forming operations [S]. The future goal is the reduction to the minimum of the expensive ‘trial and error’ method using different tool designs. In the following, a
K. Siqyr t et (11./ Jowd
92
q/‘ Muterids
Processing
Tchrology
71 (1997)
91- 99
Fig. 3. Cold forging steps for drag links. Fig. 1. Slug and cold forged tripod CVJ inner race.
survey over recent developments of the precision forging of aluminum and steel components is presented.
2. Cold forging of a tripod constant velocity joint inner race ‘i’ke constant velocity joint (CVJ) is a commonly used part of the power train for modern front driven compact cars. The forming of a tripod CVJ inner race represents a nearly classical processing technique for a one-step cold forging operation with lateral material flow, Fig. 1. The lateral extrusion of the CVJ inner race is similar to the production of spiders for steering systems that represents an exceptional high developed and cost-effective precision forging for net-shape parts [2]. The part is designed mainly of two conical halves with three circumferential located radial trunnions, Fig. 2. The central body exhibits a conical recess at the top and a cylindrical recess at the bottom for further cutting operations. This part demonstrates how the design of the part affects the material flow, quality, tool wear and fatigue [3]. The tripod inner race was cold forged from cut Design“6”
,
10
2.9
y
rml
‘corm /
Shouldi
Fig. 2. Design variations of tripod CVJ inner race.
slugs of a medium-alloyed cold working steel 20MnCr4 (similar to AISI 5120) in an oil-gas spring supported horizontal split die. This tool provides a simultaneous axial mat xial flow during the forging operation. The accuracy of the part’s symmetry plane can be held to a reproducible tolerance of + 0.15 mm (0.0006 in.) to the top and bottomn surfaces. The design ‘B’ in Fig. 2 represents a design variation where a spherical plane is added at the transition between the trunnions and the double-conical central body of the CVJ inner race. For the parts with design ‘A’ of Fig. 2. the radial material flow into the cavity is: ua;iform and the surface of the trunnions is without any defects. While forming the inner race with design ‘B’ of Fig. 2, the lateral extruded material looses contact with the cavity shoulders which produces folding laps at the upper surface of the trunnions. These faults may cause some additional material allowance for grinding operations.
3. Cold forging of drag links The forging process of drag links demonstrates the necessity of an accurate layout of the forging stages for precision near-net-shape forging. For this part, cost savings of about 10% could be realized by precision cold forging instead of former hot forging. This part is produced by five cold forging operations in a multi-stage press from an upset wire slug of AISI 1035 steel, Fig. 3. The design of the forming stages is decisive for the productivity of the forging process of this part. For optimal material flow and complete filling of the cavity, the forming process had to be arranged as a five-step cold forging sequence. Fig. 4 illustrates the layout of the forming stages from the preformed slug to the final forging operation. For all five forming steps an oil-gas spring supported horizontal split die is required. The subsequent forging operations are strongly influenced by the preform design and material distribution.
Upsetting
Cup Extwsion
Piercing/ Rod Extrusion
Cutting Load
Direction
Fig. 4. Forming
oprrations
tbr t!le cold forming
Production aspects and material flow considerations lead to the bail-like design of the preform head with differenc radii. This design prevents the complete tilling of the cavity already at the beginning of the first forging process. The uniform radial material flow provides a simultaneous die contact in the cavity that leads to reduced die loads compared with a spherical geometry. Before forming, the preform is annealed, zinc phosphated and coated with molybdenum-disuifide or soap. During the first forming stage the ball-like head of the ,eform is formed into an cylindrical shape and the angular transition between shaft and head is formed. The height of the final formed part, symbol ‘A’ in Fig. 4, is determined already at the first forming step. Further material Aow is prevented because of the perpendicular orientation of the shaft towards the forming direction. In order to avoid overfilling at the next forging stage, the upper side of the cylinder, symbol ‘B’ in Fig. 4, is inclined by 10” towards the horizontal axis. At the bottom of the head a bail shaped element is formed to improve the material distribution and to support the material flow. With the second forging operation a conical cup is formed at the upper side of the drag link head, symbol ‘D’ in Fig. 4. At the bottom an oval recess is formed, ‘C’ in Fig. 4, because the shape of the final part exhibits an oval hole in this area. In order to prevent folds. the punch cross-section has to be larger than the spherical cavity. This ensures lateral material flow and increase of the surface.
or dr;q
links.
The third forging stage combines a backward cup extrusion and a simultaneous forward rod extrusion. The backward cup extrusion transforms the conical cup into an oval shape. With the simultaneous rod extrusion process surplus material is displaced. A strong reduction of the forging loads is the second effect of the oval force relieving hole at the bottom. In order to Facilitate the piercing of the bottom at 111,~ last stage without a die. the shape of the cup-bottom at the Iast stage without a die, the shape of the cup-bottom transition exhibits a weakened cross section. In the fourth and fifth stage the drag link is cut at the transition between shaft and cup. symbol ‘E’ in Fig. 4 and pierced at the bottom.
4. Optimization of the tooling for the cold forging of helical gears Several forging methods like cold [9.10] and warm forging [ 1l] have been investigated to produce near-netshape helical gears. The goal is the fabrication of ready-to-install gears without any further grinding operation. The lateral cold extrusion process was developed to improve the precision of forged helical gears and to reduce the production costs compared with conventional cutting operations. Cold forged gears provide a high surface quality, load appropriate grain orientation and the effect of work hardening during extrusion that
94
K. Siegert et al. /Journal
of Materials Processing Tecltttolog~~71 (1997) 91-99
Fig. 5. Schematic of the lateral extrusion process for helical gears.
produces increased fatigue strength values [2]. Fig. 5 shows a schematic of the lateral extrusion process for helical gears. The die consists of an oil-gas spring supported split die. At the beginning of the extrusion process the slug is supplied into the die. After the traveling ram has closed the die, the upper and lower punch transform the ram load into the slug so that the material is displaced into the confined toothed cavity. In the last stage, a build-in ejector lifts the lower punch and ejects the forged gear while the die is free to rotate due to the helix angle. Previous studies [9,10] demonstrated the necessity to compensate the die deflection that occurs during the cold forging of gears. For common use like gearings, the gears should have a tolerance of IT 7 and better. For applications with reduced precision demands IT 9 is recommended. With uncorrected dies tolerances between IT 10 and IT 12 are attainable due to: ?? heat generation during forming ?? elastic spring-back of the workpiece ?? distortion due to work hardening of the teeth ?? die deflection With the die deflection as the dominant parameter [12] for cold forging, any method to improve the precision of cold forged helical gears has to emphasize the evaluation and the control of the die load and die deformation. In [13] a numerical analysis with 2D and 3D FEM- and BEM-codes was employed on the lateral extrusion of gears. Further investigations lead to the development of a self-optimizing program system with automated precorrection and optimization of the die geometry (Fig. 6)
v41. In combination with a finite-element simulation program this tool calculates a modified die geometry where profile mismatch and elastic die deflection will be compensated while the forging loads are applied. The calculation of a deflection-compensated die is based on a 3D finite-element simulation with the FORGE 3 code foi
the calculation of the forging loads and stresses. Because the periodical structure of the entire workpiece permits a meaningful subdivision, the slug was modeled as a segment from a torus for the simulation of one single tooth. The die is represented by the tooth’s outer surface. Fig. 7 displays the calculated pressure values perpendicular to the die surface at the end of the lateral extrusion. During the lateral extrusion process the material fills the tooth cavity from the bottom to the top. At the end of the forging process the upper and lower corner of the tooth are filled while the calculated pressure values can rise up to 4000 MPa and more. Other authors [15] found with theoretical investigations that the contact pressure values are mainly influenced by the pressure angle and the pitch displacement. Because FORGE3 is not suitable for the calculation of the die deflection, this simulation was performed in a separate step with the FEM-code-PERMAS. The calculated nodal displacements of the die are applied as input data for the developed optimization program that calculates the correction for the appropriate die dimensions. First, this program converts the nodal displacements into geometrical deviations as defined in DIN standard 3960 and calculates the radial deflection of the die. In the following iteration steps the program calculates the corrected die profile based on the modification of helix angle, pitch diameter and base circle. Generally, the modified tooth profile exhibits a reduced helix angle due to the deformation of the tooth profile. The iteration stops when the calculated data is below a defined limit. With the calculated toothing parameters the modified die geometry can be designed with a 3D CAD-system and produced via spark erosion. Fig. 8 shows the measured deviations of lateral extruded gears before and after correction of the die geometry. The maximum die deflection with x0.35 mm (0.014 in.) occurs in the radial direction. With the
95
EM - Net Generator
HelicalGear
Relational Data Selection to store the Results
Selected Hvdraulie Prew Cokd Forging of Helical Gears
II
CAD-Svstem
bptirnizetion Modul Data Selection
Fig. 6. Development system for simulation and optimization of the lateral forging of helical gears.
developed optimization module, this radial die deflection can be compensated up to 95%. For an optimal die compensation the toothing deviations have to be considered additionally. Because the toothing of the die is produced by spark erosion, the tooth parameters cannot be varied along the vertical axis. With this limitatioi: a reduction of the toothing deviations up to 50% is attainable.
an
in N/mm2
5. Cold extrusion of P/M high Glicon con&es&s
al~jnu~
materials with
Hypereutectic aluminum-silicon alloys are commonly used for applications where high wear resistance and low thermal expansion coefficients are required. With conventional ingot (I,‘M) metallurgy techniques alloys with silicon contents up to a maximum of 25% can be produced. P/M processing routes, like the spray deposition process, allow the production of aluminum alloys with silicon contents up to 35% and more. These alloys exhibit a segregation free microstructure with a fine and uniform distribution of primary silicon crystals. For the production of thin-walled tubes with small tolerances from hot extruded high-silicon P/M grades, the cold extrusiou process was condidered to be the most effective forming technique. A technique to increase the formability of metal matrix composites (MMC) by means, of superimposed counter pressure is presented in [ 161.FOI manufacturing rea.sons, the cold extrusion of hypereutectic P/M tubes should Qe feasible with a single-action press. In this case, the material behavior during cold extrusion is the critical aspect that has to be considered. Like MMC, cracks are ini+$ted by the fracttrre of the embedded primary silicon p:u-titles. In order to optimize the forming process, the inflr;:ence of the following extrusion parameters on the materials microstructure was investigated: Equivalent strain. strain rate Max. straii, stat64 stress, strain diiributii &r.. &aiJl, sMe.of-Gress. straii distiibution .
Fig. 7. Cakulated
die contact pressure values from 3D FEM-simula-
tions at the end of the lateral extrusion (extended to three teeth).
sbain rate,fklw stress... Flaw stress. slrain hardening,famability
Tooth Profile Deviations -
“‘bmt
bmt
(b =
Radial Deviations
Angular Deflection
Profile Deformation
bmt
bmt
bmt
Root Circle
Crown Circle
Depth of Tooth
bmt
Bottom. m = Middle, t = Top of the Die Insert)
Fig. 8. h+zdsured profile deviations of cold forged helical gears.
Fig. 9 shows the workpieces from cut thick-walled tubes with outer diameters of 88 and 94 mm and the cold extruded thin-walled tubes. The investigated material Dispal S261X is based on an aluminum matrix with 25”/ silicon added to the melt before spray deposition. After spray deposition, the billet was extruded to thick walled tubes at 420°C with subsequent cooling at atmosphere (‘F’ condition). Fig. 10 shows the dimensions of the preform from cut P/M extrusions and the cold extruded tubes. The hot extruded thick-walled tubes were hollow forward extruded at room temperature in the ‘F’-condition and after soft-annealing. The parts extruded in the ‘F’-condition showed severe cracking of the primary silicon crystals whereas much less particle cracking occurred in the soft annealed extrudates. In order to optimize material flow and the state-ofstress during cold extrusion, simulations were performed with the finite element program DEFORM 2D. For a given temperature the material’s formability is mainly inf&enced by the state-of-stress, which is _I__-
represented by the quotient of mean stress G,,, and flow stress 6, [17]. Fig. I 1 illustrates the state-ofstress curve over the equivalent strain for several material points during the hollow forward extrusion of the tubes. For the most part of the extrusion process, the state-of-stress is characterized by large amounts of compressive stresses (negative values of g,,,/~,) which increase the material’s formability. For the materials points located at the bottom of the workpiece (symbol ‘e’ in Fig. 1I), the state-of-stress does not change significantly. For the other points (symbol ‘a’-‘d’ in Fig. 11) the state-of-stress drops significantly at the beginning of the extrusion process and increases with increasing strain. At the end of the deformation the material’s points are subjected to small amounts of tensile stresses (positive values of a,,,/~~). The state-of-stress is influenced by the die-openingangle 2a and the die fillet radius. With decreasing die-opening-angle, the state-of-stress becomes more fa-
1
Dlspal S281X (Al + 25%SI),
ColdExtid
Tube
2a
Dimension
88 doo=88mm
doo=94mm
(cpg= 1.22)
(qg = 1.52)
Fig. 9. Cut thick-walled tubes.
mm
preforms and cold extruded thin-walled Fig. IO. Dimensions of the preforms and the cold extruded tubes.
Al ; 25%S1 2ca=90” dao=94mm dal=74mm di=66mm
T=RT
2.5 Fig. 13. Production stages of ;I connecting rod forging. Fig. 1I. State-of-stress
during hollow forward extrusion of tubes.
vorable for the deformation to high strains. The quotient GJCJ, reaches higher negative values at the bcginning of the forming process and lower positive values at the end of the extrusion process. As the simulation results does not consider the characterisrics of the lubrecation system, the results may drive to the conclusion that a small die-opening-angle would be favorable. Extrusion tests with a die-opening-angle of 2r = 60” showed that the lubricant (zinc stearate) failed and severe welding of the workpiece material occurred on the die surface. Fig. 12 shows the microstructure of extruded tubes for different die fillet radii. Small fillet radii lead to a strong bending and shearing of the material while passing the die [18]. This produces increasing amounts df strain induced into the workpiece compared with a more rounded die and leads to an increased fracture of the silicon particles.
6. Forging of connecting rods Connecting rods were forged from extruded rods of a high silicon P/M aluminum alloy 1191. The alloy is
R = 0.5 mm
Fig. 12. Microstructure
R=2mm
of extruded tubes for different die lillet radii.
composed of a conventional AA-2024 matrix with additional 17% Si, 5% Fe and 0.6% Zr. A two-step w;~r~l~ forging process ~/as applied to overcome the limited formability of the P/M material at room temperature. Fig. 13 shows the different forging stages from the extrusions to the completely machined part. The forging process was arranged as a two stage sequence of closed die forging with flash and closed die forging without flash. Closed die flashless forging was chosen to improve the surface and the material properties of the rods. For the complete filling of the confined &e cavity an exact mass distribution of the preformed part is required. On the other hand, deticiencies of material lead to unfilled corners of the cavity. The preform shape was derived from the shape of the finally forged workpiecc by volume constant geometry variations to optimize the material flow. Closed die precision forging without Rash rcquirc~ enhanced accuracy and reduced tolerances in tool design. Compared to DIN and international standards. the side wall thickness (3 mm) and the base thickness (3 mm) of the connecting rods are designed smaller than for conventional forgings [20,2I]. General tolerances for precision forging are reduced to typically half of those for conventional forgings. Thickness tolerances are reduced to about two thirds of their former values. For precision forging of aluminum alloys in steel cavities the shrinkage of the forged and cooled down workpiece has to be considered. The different thermal expansion coefficients of workpiece and die cause different thermal extensions at elevated temperatures. Fig. 14 reveals the curves for a constant thermal expansion difference between workpiece and die in the thermal working area for the forging of the connecting rods. The cavity dimension, c were extended by 03’41 to reduce the shrinkage of the forged workpiece in the thermal working range. The forging tool, Fig. 15, was designed as an electric heated die with changeable die inserts for the first and the final forging stage. Closed loop temperature cOiltrOl systems and temperature measurement systems were used to monitor the temperatures of the dies.
K. Siegert et al. / Jotrrttal of’Materials Promsittg
98
Tedtttology 71 (1997) 91-99
wrM@aa~Dm~s2.m (Al l,si Sk3.Scu C.Ch!+cl.~ oii mlww 233s fX3ZCrMOV33.wmpnd) II
t*)r
I
=a30
r3.340%
.IT.i04lK
l-hWd~dikN=+ALl Am=m_.A&--.A&,, ” _._.._
+,$.
-0 -me
.I&
-
-560 ._._._
workpkce
Temperature
lncmsse
d&rp in
-sdm
K
Fig. 14. Temperature influence on thermal tween workpiece and die.
expansiondifferencebe-
CAD designed die inserts were CNC-milled of a prehardened tool steel, whereas high temperature steel X32 CrMoV 33 was used for the die inserts. For the first forging stage, the die inserts were designed with a symmetrical parting line and conventional [20] draft angles and transition radii. The tool for the fin,al forging stage consists of punch and die witil a la.teral tolerance of + 0.01 mm and draft angles of 0”. Two ejectors were placed in the die near the bearing areas of the connecting rod to release the forged workpiece. The preforms were closed die forged with flash from turned extrusions in a temperature controlled upper and lower die. After separating the flash the workpiece was precision forged in a closed die without flash. For preforging of the connecting rods, the workpiece temperature was varied between 300 and 500°C, die temperature was adjusted to 300°C. Graphite-water emulsion was used to reduce friction between workpiece and the cavity surface. The
Fig. 16. Height measurements ent forging temperatures.
for connecting rod preforms for differ-
For a combined forging and heat treatment process, the workpiece preheating temperature is determined from the annealing temperature of the material. Thus, the preheating temperature of the final forging stage was varied between 460 and 5OO”C,the die temperature was 450°C. Conventional furnace heating and also induction heating were used for preheating the initial and preforged workpieces. The volume of the finished part is directly determined by the volume of the preforged workpiece. Hence, the complete filling of the confined cavity requires a minimum volume of the workpiece. As a measure for the volume of the preforged workpiece, the height of the bearings was examined for different forging temperatures. Because of the elastic tool deformation, the height of the preforged parts is also affected by the flow stress of the workpiece. Fig. 16 shows the influence of tlhe workpiece temperature on the height measurements. for preforged connecting rod and the reference lines for the minimum heights required for complete filling of the cavity. Both curves decrease with increasing temperature. Between 380 and 460°C the influence of the workpiece temperature on the height measurements is reduced to 0.05 mm. For the finished forged connecting rods height measurements showed a scatter of about &-0.1 mm. At the end face of the rods a small vertical flash with a thickness of x 0.2 mm occurred. With a subsequent cold sizing operation a further reduction of the tolerances may be achieved. Quenching and arfificial aging of the forged connecting rods produces a significant increase in hardness of z 50% compared with the extruded material. Compared with conventional heat treatment the distortion of the workpiece produced by annealing and quenching can be reduced. References
Fig. IS. Connecting rod forging tool. Right side: final forging stage (without flash); and left side: first forging stage (with flash).
[I] M. Hirschvogel, Transmission shaft forgings-technical and economical aspects of new developments, Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May, 1995.
PI R. Geiger, M.
tinsel, From near net-shape to net-shape cold forging--state of the art. Proc. 9th Pm. Cold Forging Congr.. Solihull, UK, 22-26 May, 1995. [31 M. Kammerer. Th. Werle. K. Piihlandt, Entwicklung von Stadienpllnen zur Fertigong schwieriger Fliessressteile, in: K. Siegert (Ed.), Ncuere Entwicklungen in dcr Massivumformung 1995 DGM Informationsges. mbh, Oberursel. 1995. [41 M. Kammerer, Werkzeuge zum Fliesspressen von Aluminium, in: K. Siegert (Ed.), Fliesspressen von Aluminium, Vol. 1, DGM lnformationsges. mbh, Qberursel. 1995. 151 T. Alum, K. Sweeney, V. Vazquez, H. Kim. M. Knoerr, Cold forging of complex shaped parts to close tolerance-application of metal flow simulation to process and tool design, Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May, 1995. Konstruktion und [61 K. Siegert, R. Neher, Computerunterstiitzte Fetigung sowie massliche Kontrolle mit Datenriickfiihrung bei der Herstellung von Schmiedegesenken und ..teilen, in: K. Siegert (Ed.), Neuerc Entwicklungen in der Massivumformung 1991 DGM Informationsges. mbk, Oberursel. 1991. Hochgeschwindigkeitsfrasen von Neher, .I. Boxier. [71 R. Graphitelektroden. in: K. Siegert (Ed.), Rechneranwendung in der Umfonntechnik. DGM Informationsges. mbh, Oberursel, 1992. [El G. Du. Ein wissernbasiertes System zur Stadienplanermittlung beim Kaltmassivumformen. Berichte aus dem Institut fur Umformtechnik Nr. I 11. Springer-Verlag. Berlin, 1991. Stirn[91 F. Schmieder, Beitrag zur Fertigung von schragvenahnten radern durch Querpressen. Berichte aus dem lnstitut fiir Umformtechnik Nr. 118. Springer-Verlag. Berlin. I99 I. HOI V. Szentmihalyi, Beitrag der Prozesssimulation zur Entwicklung komplexer Umforrnteile. Berichte aus dem lnstitut fur Umformtechnik Nr. I2 I. Springer-Verlag, Berlin. 1994.
[I I] E. Doege. H. Nagele. FEM-gcstlutzte Auslrgung van Prazisionsschmiedepro/e+icn ~.rlnlornrtechnik 29 ( I ) ( I9951 ?? 10. [I’] M. Lamer. Untersuchungen ubrr d&s KaltHicssp~e~~rn geradund schragverzahnter Stirnrider. VDI Fortschrittsberichtc Rcihe 3 Nr. 221, VDI-Verlag. Dusseldorf, 1991. p, Ihl. [l3] K. Lange. V. Szentmihalyi. Optimized cold fot,ginp of helical gears by FEM-simulation. Proc. 9ih Inl. Cold Forging Congr.. Solihull, UK, 22-26 May. 1995. 1141 T. Kepplcr-Ott. Rechnerunterstiitzte Optimierung der Werkzeuggeometrie fur die Herstellung von gerad- und schragverzdhnlen Stirnradern durch Verfahren der Massivumformung, cnpublished report, 1996. [l5] F. Dohmann, N. Lintel, Konlaktspannungen beim Verzahnungspressen, Umformtechnik 29 (I) (19S5) 46-51. [I61 H.W. Wagner, J. Wolf. Kaltmassivumforrinmg von MMCs in Form der Matrixlegierungen AlMgSit und AlSil2 mit dem Ver-stiirkungswerkstoff SIC. Verbundwerkstoffe und Werkstoffverbunde, Symp. Deutschen Gesellschaft fiir Materialkunde, Chemnitz. 17-19 Juni. 1993. DGM Informtionsges. mbH. Oberursel. 1993, S.43350. [I 71 K. Siegert, D. Ringhand, Formability of dispersion strengthened P/M aluminium alloys. Proc. 9th Int. Cold Forging Congr., Solihull, UK, 22-26 May. 1995. [I 81 K. Sicgert. K.-J. Fann. D. Engelmann. Simulation des Yoll-Vorwarts-Fliesspressens von Aluminium, in: K. Siegert (Ed.), Rechneranwendung in der Umformtechnik. DGM Informationsges. mbh. Oberursrl, 1992. [l9] K. Siegert. D. Ringhand. Flashless and precision forging of connecting rods from P,:M aluminum alloys. J. Mater. Proc. Technol. 36 (1994) 157-167. [ZO] DIN l?49, Gesenkschmiedestiicke aus Aluminium. 1211 H. Meyer-Nolkemper. Gesenkschmieden von Aluminiumwerkstoffen I-IX. Aluminium 55 (1979).