material and surface engineering for precision forging dies
TRANSCRIPT
Material and Surface EngineeringFor
Precision Forging Dies
By
Sailesh Babu,Dilmar RibeiroRajiv Shivpuri
The Ohio State University
Prepared for
Precision Forging ConsortiumOhio Aerospace Institute and
National Center for Manufacturing Sciences
June 10, 1999
II
EXECUTIVE SUMMARY
This report is prepared for the Precision Forging Consortium as a deliverable under
contract from the Ohio Aerospace Institute and the National Center for Manufacturing
Sciences. The master project, supported by the Department of Energy, is entitled:
“Energy and Waste Minimization through Precision Forging for the Manufacture of
Complex Shapes.” This report is a companion report to the one produced under
contract by the team at Laval University. The focus of the latter report is on innovative
and advanced die material systems.
This report provides a comprehensive overview of the state-of-knowledge of die
materials and surface engineering for forging dies. Since hundreds of materials exist
that may have applications for forging dies, the authors have tried to select those
materials, which in their opinion, have direct relevance to precision forging. The
authors have been selective on materials types but comprehensive on the issues that
must be addressed before these materials can be used optimally in a precision forging
environment.
This report provides information on the following topics:
• Conventional die steels: physical and mechanical properties, die block
manufacturing, and heat treatment. Properties relevant to wear and failure
prevention. Suggestions on their optimal utilization.
• Advanced Die Materials and Surface Engineering: properties and wear behavior.
• Failure Mechanisms and Models that can be used for predicting wear behavior of
die materials in a forging environment. Details of the models provided in
Appendices.
• Manual for DieLit: The ENDNOTE based Database of published information on
die materials, their properties and wear behavior. Includes the references
available at OSU and the classification of this information.
• Manual for SAMS: the smart die material selector software, which has been
developed in the ACCESS environment.
While available information on materials (both conventional and advanced) is
enormous, the properties and relationships needed to optimally select or design
materials and surface engineering for increased lives of dies are incomplete at best and
possibly missing. This report is intended to provide foundation and identify gaps in
knowledge for the Phase II of the Precision Forging Project.
III
TABLE OF CONTENTS
EXECUTIVE SUMMARY…………………………………………………………………………………iiLIST OF FIGURES……………………………………………………………..………………………….vLIST OF TABLES…………………………………………………………………………………………..x
1. INTRODUCTION .............................................................................................................................. 11
1.1. GOALS OF THE PRECISION FORGING CONSORTIUM ........................................................................ 121.2. TASKS FOR PHASE I: GROUP FOR INCREASED LIFE OF DIES/ OSU .......................................... 12
2. A BRIEF REVIEW OF FAILURE OF FORGING DIES .............................................................. 14
3. MATERIAL FOR FORGING DIES................................................................................................. 17
3.1. HOT WORK DIE STEELS .................................................................................................................. 173.2. PHYSICAL AND MECHANICAL PROPERTIES OF VARIOUS TOOL STEELS ............................................ 20
3.2.1 Resistance to deformation at high temperatures................................................................... 203.2.2 Resistance to mechanical shock and fatigue......................................................................... 243.2.3 Resistance to thermal softening ............................................................................................ 273.2.4 ductility ................................................................................................................................. 29
3.3. MARAGING AND OTHER STEELS .................................................................................................... 313.3.1 Composition.......................................................................................................................... 313.3.2 Properties ............................................................................................................................. 31
3.4. SUPERALLOYS ............................................................................................................................... 35
4. DIE BLOCK MANUFACTURING AND HEAT TREATMENT.................................................. 39
4.1. DIE BLOCK MANUFACTURING: CLEANLINESS AND MICROSTRUCTURE......................................... 394.2. CAVITY MANUFACTURE: MACHINING AND EDM ......................................................................... 434.3. HEAT TREATMENT: AUSTENIZING, QUENCHING AND TEMPERING ................................................ 45
4.3.1 Austenitzation and soaking ................................................................................................... 464.3.2 Quenching............................................................................................................................. 514.3.3 Tempering ............................................................................................................................. 56
4.4. SPECIFICATIONS: DIE STEEL, HARDNESS, TOUGHNESS AND MICROSTRUCTURE ........................... 61
5. SURFACE TREATMENTS............................................................................................................... 63
5.1. CARBURIZING ................................................................................................................................ 645.2. NITRIDING ..................................................................................................................................... 655.3. CARBONITRIDING AND NITROCARBURIZING .................................................................................. 685.4. BORIDING ...................................................................................................................................... 695.5. THERMO-REACTIVE DIFFUSION (TRD).......................................................................................... 705.6. OXIDE COATINGS .......................................................................................................................... 70
6. ADVANCED DIE MATERIALS AND SURFACE ENGINEERING TECHNIQUES................ 72
6.1. CERAMICS: SIALON, SILICON NITRIDE AND SILICON CARBIDE ..................................................... 726.2. ALUMINIDES: NICKEL AND TITANIUM ........................................................................................... 736.3. WELD OVERLAYS .......................................................................................................................... 776.4. CRYOGENIC TREATMENTS............................................................................................................. 806.5. BRUSH PLATING TECHNIQUES ....................................................................................................... 806.6. VAPOR DEPOSITION: PVD AND CVD............................................................................................ 856.7. THERMAL SPRAYING ..................................................................................................................... 866.8. LASER SURFACE MODIFICATION ................................................................................................... 876.9. ION IMPLANTATION ....................................................................................................................... 88
7. MECHANISMS AND MODELS OF DIE WEAR AND FAILURE .............................................. 89
7.1. WEAR ............................................................................................................................................ 89
IV
7.2. PLASTIC DEFORMATION ................................................................................................................ 907.3. MECHANICAL FATIGUE ................................................................................................................. 917.4. THERMAL FATIGUE ....................................................................................................................... 93
8. CLOSURE........................................................................................................................................... 95
9. APPENDIX A – FUNDAMENTALS OF DIE FAILURE............................................................... 98
10. APPENDIX B – WEAR INDICES OF VARIOUS DIE MATERIALS ................................... 118
11. APPENDIX C - PROCESS EFFECT ON DIE LIFE ................................................................ 123
REFERENCES ......................................................................................................................................... 139
V
LIST OF FIGURES
FIGURE 1-1. (CSER, GEIGER ET AL. 1993) ..................................................................................................... 12FIGURE 2-1. FREQUENCY AND LOCATION OF TYPICAL DIE FAILURES IN FORGING (CSER, GEIGER ET AL. 1993)
.............................................................................................................................................................. 14FIGURE 2-2. COMPLEX INTERACTION OF FORGING PARAMETERS AND WEAR ARTINGER, (CSER, GEIGER ET AL.
1993)..................................................................................................................................................... 15FIGURE 2-3. SOME ASPECTS OF FORGING AND PROCESS DESIGN THAT AFFECT WEAR AND FRACTURE LANGE,
K IN (CSER, GEIGER ET AL. 1993) ......................................................................................................... 15FIGURE 3-1. (A) VARIATION OF HARDNESS OF DIE STEEL OF H-12 AND 6F2 WITH TEMPERING TIMES. H-12
USED WAS AUSTENETIZED FROM 1040 C WITH AS QUENCHED HARDNESS OF 584 VPN, 6F2 WAS
AUSTENETIZED FROM 850 °C WITH AS QUENCHED HARDNESS OF 601 VPN (B) VARIATION YIELD
STRENGTH OF DIFFERENT TOOL STEELS WITH TEMPERATURES. (NAGPAL 1976).................................... 20FIGURE 3-2. YIELD STRENGTH AND HARDNESS VERSUS TEMPERATURE FOR SEVERAL TOOL STEELS
(NORSTROM, JOHANSSON ET AL. 1981)................................................................................................. 21FIGURE 3-3. VARIATION OF YIELD STRENGTH WITH TEMPERATURE (SEMIATIN AND LAHOTI 1981).............. 21FIGURE 3-4. MECHANICAL RESISTANCE VERSUS TEST TEMPERATURE EXPRESSED IN TWO FORMS FOR THE H13
AND A NEW HOT WORK TOOL STEEL (QRO90) (ROBERTS AND NORSTROM 1987)................................. 22FIGURE 3-5 YIELD STRENGTH AND DUCTILITY VERSUS TEST TEMPERATURE (A5) ELONGATION (Z) AREA
REDUCTION. A) H13 AND B) QRO90 (UDDEHOLM ).............................................................................. 22FIGURE 3-6 TENSILE STRENGTH AND DUCTILITY VERSUS TEST TEMPERATURE FOR STEELS FROM H13 GROUP
(H12, H11, H10), PLUS AND HIGH ALLOY, H21, AND A LOW ALLOY GROUP 6F3 (THYSSEN ). .............. 23FIGURE 3-7. VARIATION OF CHARPY TOUGHNESS WITH DIFFERENT HARDNESS LEVELS AND TESTING
TEMPERATURES ON OF HOT WORK DIE STEELS (VALUES IN PARENTHESES INDICATE HARDNESS AT ROOM
TEMPERATURE) (NAGPAL 1976)............................................................................................................ 24FIGURE 3-8 VARIATION OF TOUGHNESS FOR SEVERAL TOOL STEELS IN FUNCTION OF HARDNESS AND YIELD
STRENGTH (CSER, GEIGER ET AL. 1993)................................................................................................ 24FIGURE 3-9. VARIATION OF TOUGHNESS FOR SEVERAL TOOL STEELS IN FUNCTION OF HARDNESS AND YIELD
STRENGTH (SHIVPURI AND SEMIATIN 1988) .......................................................................................... 25FIGURE 3-10 COMPARISON OF TOUGHNESS PROPERTIES FOR H13, H21 AND A NEW HOT WORK TOOL STEEL
QRO80M VERSUS TEST TEMPERATURE (JOHANSSON, JONSSON ET AL. 1985)....................................... 25FIGURE 3-11 COMPARISON OF TOUGHNESS KIC AND CHARPY V-NOTCH FOR SEVERAL TOOL STEELS. A) KIC
FOR THREE BAR SIZE, LONGITUDINAL DIRECTION, THE SMALL DIAMETERS REPRESENT REDUCTION FROM
THE BIGGER DIAMETER; B) CHARPY V-NOTCH LONGITUDINAL AND TRANSVERSAL C) KIC FOR H13 INFUNCTION OF THE HARDNESS, AUSTENITIZED AT 1024°C, TIME 25 MIN, AIR COOLED (HEMPHILL AND
WERT 1987). ......................................................................................................................................... 26FIGURE 3-12 TOUGHNESS VERSUS TEST TEMPERATURE FOR A) H13 AND B) QRO90 (UDDEHOLM )............. 26FIGURE 3-13.(A) VARIATION OF HARDNESS WITH TEMPERATURE FOR H-11, H-12, H-13, H-14 AND
PYROVAN. MEASUREMENTS WERE MADE AFTER HOLDING SAMPLES FOR 30 MINUTES FOR
HOMOGENIZATION (B) RESISTANCE OF HOT WORK DIES STEELS TO THERMAL SOFTENING AS MEASURED
BY THE ROOM TEMPERATURE HARDNESS (NAGPAL 1976) ..................................................................... 27FIGURE 3-14 COMPARISON OF PROPERTIES FOR H13 AND A NEW HOT WORK TOOL STEEL QRO80M VERSUS
TEST TEMPERATURE (A) STANDARD TEMPERING CURVE (JOHANSSON, JONSSON ET AL. 1985). (B)MASTER TEMPERING CURVE FOR PREMIUM H13 WERE: P= LARSOM-MILLER PARAMETER, T ISTEMPERATURE (°F), T IS TIME IN HOURS (CARPENTER )......................................................................... 27
FIGURE 3-15 THERMAL EXPANSION FOR SEVERAL TOOL STEELS VERSUS TEMPERATURE (ROBERTS, KRAUSS
ET AL. 1998).......................................................................................................................................... 28FIGURE 3-16. DUCTILITY OF VARIOUS DIE STEELS AT HIGH TEMPERATURES (NAGPAL 1976) ....................... 29FIGURE 3-17 COMPARISON OF DUCTILITY FOR H13, H21 AND A NEW HOT WORK TOOL STEEL QRO80M
VERSUS TEST TEMPERATURE (JOHANSSON, JONSSON ET AL. 1985). ...................................................... 29FIGURE 3-18 (A) MASTER TEMPERING CURVE, T IS TEMPERATURE IN KELVIN, T IS TIME IN HOURS (B) HOT
HARDNESS OF HWM COMPARED TO H- 13 (KASAK AND STEVEN 1970) (C) AGING CURVES................. 32FIGURE 3-19 PROPERTIES VERSUS TEST TEMPERATURE OF MARAGINS STEELS COMPARED WITH H13 (A)
HARDNESS (B) YIELD STRENGTH AND DUCTILITY (BAYER 1984) .......................................................... 32
VI
FIGURE 3-20 PROPERTIES ON A MARAGING STEELS 18 NI 300 AND H13. (A) HOT-HARDNESS OF MEASURED
AFTER HOLDING AT THE TESTING TEMPERATURES FOR 30 MIN. (B) TENSILE PROPERTIES VERSUS TESTING
TEMPERATURES C) TOUGHNESS OF VERSUS TEST TEMPERATURE (BARRY, WILLS ET AL. 1968) ............ 33FIGURE 3-21 PROPERTIES OF A HOT WORK MARAGING STEEL COMPARED WITH H13 GROUP AND THERMAL
FATIGUE RESULTS: A) FIELDS OF AUSTENITE AND MARTENSITE IN FUNCTION OF TEMPERATURE SHOWING
THE EFFECT OF NI% (BRANDIS AND HABERLING 1987) (B) AGING CURVES FOR 18%NI AND 12%NI C)THERMAL FATIGUE RESISTANCE FOR H13 GROUP AND A MARAGING STEEL (D) HOT YIELD STRENGTH
(GEHRICKE 1993; GEHRICKE, KLARENFJORK ET AL. 1995) .................................................................. 34FIGURE 3-22 COMPARISON OF TOUGHNESS CHARPY V-NOTCHED FOR A H13 STEEL AND A MARAGING STEEL
(MARLOCK) AT TWO TEST TEMPERATURES (DORSCH 1991).................................................................. 34FIGURE 3-23. DUCTILITY AND TOUGHNESS VERSUS TEST TEMPERATURE (SEMIATIN AND LAHOTI 1981) ..... 35FIGURE 3-24. MECHANICAL PROPERTIES EXPRESSED AS HARDNESS AND YIELD STRENGTH VERSUS TEST
TEMPERATURE (SEMIATIN AND LAHOTI 1981) ...................................................................................... 36FIGURE 3-25. COMPILATION OF SEVERAL PROPERTIES VERSUS TEST TEMPERATURE FOR SUPERALLOYS FROM
OHUCHI (OHUCHI 1990). A) HARDNESS B) THERMAL EXPANSION C-D) ULTIMATE TENSILE STRENGTH
AND YIELD STRESS. ............................................................................................................................... 36FIGURE 4-1. SCHEMATIC COMPARISON OF DUCTILITY (CHARPY UNNOTCHED) AND TOUGHNESS (CHARPY V
NOTCH) VERSUS TEMPERATURE, ROOM TEMPERATURE (RT) (NORSTROM 1989) .................................. 39FIGURE 4-2. EFFECT OF SULFUR CONTENT ON THE TRANSVERSE FRACTURE TOUGHNESS OF H-13 DIE STEEL
(ROBERTS AND NORSTROM 1987). ........................................................................................................ 40FIGURE 4-3. EFFECT OF COARSE GRAIN BOUNDARY CARBIDES FROM ANNEALED TOOL STEEL ON THE
TOUGHNESS AT ELEVATED TEMPERATURES. MEASUREMENTS WERE MADE AFTER QUENCHING AND
TEMPERING (BECKER, FUCHS ET AL. 1989) B) COMBINED INFLUENCE OF CLEAN PROCESSING AND
EXTRA FINE STRUCTURE (EFS) ON TOUGHNESS (BECKER 1984). .......................................................... 40FIGURE 4-4 (A) EFFECT OF CARBIDES ON DUCTILITY OF STANDARD H-13 (B) EFFECT OF CARBIDES AND
INCLUSIONS ON DUCTILITY. (ROBERTS AND NORSTROM 1987) ............................................................. 41FIGURE 4-5 RELATION BETWEEN SAMPLES ORIENTATION, DUCTILITY AND THERMAL FATIGUE A5
ELONGATION, Z AREA REDUCTION, VW UNNOTCHED EUROPEAN SAMPLE FOR IMPACT TEST A,B,C(ROBERTS AND NORSTROM 1987) ......................................................................................................... 42
FIGURE 4-6. EFFECT OF EDM PROCESS IN DIE SURFACE; REFROZEN LAYER AND THERMAL CRACKS (CSER,GEIGER ET AL. 1993)............................................................................................................................. 43
FIGURE 4-7. SURFACE OF DIES AFTER EDM, H13 AND MARAGING STEEL (MARLOCK) (DORSCH 1991). ...... 44FIGURE 4-8. HEAT TREATMENT CYCLE OF HOT WORKING STEELS (KRAUSS 1995). ....................................... 46FIGURE 4-9. INFLUENCE OF AUSTENITIZING TEMPERATURE IN PROPERTIES VERSUS TEMPERING
TEMPERATURE; TENSILE STRENGTH, DUCTILITY AND TOUGHNESS ARE REPRESENTED FOR H13 AND DIM
2367 HOT WORK STEEL (BECKER, FUCHS ET AL. 1989). ........................................................................ 48FIGURE 4-10. EFFECT OF AUSTENITIZING TEMPERATURES ON ASTM GRAIN SIZE AND AS-QUENCHED
VICKERS HARDNESS OF H-13 (STUHL AND BREITLER 1987). ................................................................ 49FIGURE 4-11 EFFECT OF AUSTENITIZING TEMPERATURES ON AS QUENCHED HARDNESS, GRAIN SIZE AND
RETAINED AUSTENITE A) H13. B) H13 AND H11 (PICKERING 1987) ..................................................... 50FIGURE 4-12. CCT DIAGRAMS FOR TWO AUSTENITIZING TEMPERATURES , B) EFFECT OF CARBON ON MS
(ROBERTS AND ROBERT 1980) .............................................................................................................. 50FIGURE 4-13. A) EFFECT OF BAR SIZE ON THE QUENCH RATE AND THE RESULTING PHASE STARTING WITH AN
AUSTENITIZING TEMPERATURE OF A)1000° C AND B) 1050° C (SCHMITD 1987) B) VARIATION OF
TEMPERATURES ACROSS A SECTION OF H-13 DURING QUENCHING AND THE RESULTING PHASES
(BIERMANN 1984) ................................................................................................................................. 51FIGURE 4-14 A-B) COOLING RATES USED ILLUSTRATING THE CORRESPONDING STRUCTURES IN THE CCT
DIAGRAM; B) TABLE WITH THE CONDITIONS AND THE RESULTING PROPERTIES FOR THE CORRESPONDING
COOLING RATES (WALLACE 1989). ....................................................................................................... 52FIGURE 4-15. EFFECT OF COOLING RATES ON THE PHASE CONTENT AND THE RESULTING TOUGHNESS. STEELS
WAS AUSTENITIZED AT 1020 C FOR 30 MINS AND OIL QUENCHED AT DIFFERENT RATES. (OKUNO 1987).............................................................................................................................................................. 53
FIGURE 4-16. EFFECT OF DIFFERENT COOLING RATES RESULTING FROM THE QUENCHING PRESSURES FOR A
11”X20”X30” BLOCK (ROCHE, BEATON ET AL. 1997)........................................................................... 55
VII
FIGURE 4-17. A)EFFECT OF COOLING RATE ON THE LONGITUDINAL TOUGHNESS; B) EFFECT OF VARIOUS
QUENCHING MEDIUMS ON THE DUCTILITY OF SAMPLES TAKEN IN THE THREE DIRECTIONS (1,3,4PRODUCE COARSE BAINITE AND HIGH GBC PRECIPITATION) (ROBERTS AND NORSTROM 1987) ........... 56
FIGURE 4-18. DROP IN TOUGHNESS OF H-13 WITH SECONDARY HARDENING. A) REPRESENTS TOUGHNESS IN
CHARPY V-NOTCH ENERGY, B) REPRESENTS TOUGHNESS IN KIC (PICKERING 1987) .............................. 57FIGURE 4-19 SOFTENING RETARDATION EFFECTS IN RELATION FOR SEVERAL ELEMENTS IN RELATION TO A
FE-C CARBON STEEL AT TEMPERING TEMPERATURE OF 540°C (KRAUSS 1995) .................................... 58FIGURE 4-20. EFFECTS OF TEMPERING TEMPERATURE IN TOOL STEELS DIMENSION INCLUDING CONTRACTION
AND EXPANSION (PICKERING 1987)....................................................................................................... 58FIGURE 5-1. THICKNESS OF VARIOUS COATINGS AND SURFACE TREATMENTS (SUBRAMANIAN 1996)........... 63FIGURE 5-2.A. COMPARISON OF WEAR AMOUNTS OF SURFACE TREATED UPSETTING TOOLS AFTER 1000
FORGING CYCLES WITH LUBRICANT (DELTAFORGE-31) (DOEGE, SEIDEL ET AL. 1996)......................... 64FIGURE 5-3. RESULTS FOR HOT WORK TOOL STEELS IN THE H13 GROUP PRESENTED BY KRISHNADEV
(KRISHNADEV 1997) (A) COMPOSITION (B) TOUGHNESS (C) HOT HARDNESS (D) SOFTENING OF THE
ALLOYS 2-3 WITH AND WITHOUT NITRIDING E) HARDNESS ACHIEVABLE WITH DIFFERENT COATINGS AND
THE ALLOYS CHEMICAL COMPOSITION F) CHARPY IMPACT TOUGHNESS OF H-13 AND TREATED ALLOY
NO. 3..................................................................................................................................................... 67FIGURE 5-4. RELATIVE WEAR RATES OF NITRIDED AND NON-NITRIDED TOOL STEELS USED IN EXTRUSION
FORGING (DEAN 1987) .......................................................................................................................... 68FIGURE 6-1.COMPILATION OF SEVERAL PROPERTIES VERSUS TEST TEMPERATURE FOR CERAMICS FROM
OHUCHI (OHUCHI 1990). A) HARDNESS B) THERMAL EXPANSION C) YIELD STRESS. ............................. 72FIGURE 6-2. MECHANICAL PROPERTIES A) COMPRESSIVE YIELD STRENGTH FOR NI ALLOY 718 AND NICKEL
ALUMINIDE 221M-T (AL 7.6-8.2; CR 7.5-8.2; MO 1.3-1.55; Z 1.4-2.0; B 0.003-0.01 NI BALANCE) B)TENSILE AND YIELD STRENGTH FOR 221M-T ALLOY (MADDOX AND ORTH 1997)................................ 74
FIGURE 6-3 YIELD STRENGTH OF VARIOUS GRADES OF NICKEL ALUMINIDES (BLAU 1992) ........................... 74FIGURE 6-4. COMPARISON OF CRACK GROWTH DATA FOR NICKEL ALUMINIDE COMPARED TO OTHER HIGH
TEMPERATURE ALLOYS. (FUCHS, KURUVILLA ET AL. ).......................................................................... 75FIGURE 6-5 COMPARISON OF YIELD STRENGTH OF IC-15 TO THOSE OF OTHER HIGH TEMPERATURE ALLOYS.
(HORTON, LIU ET AL. ) .......................................................................................................................... 76FIGURE 6-6. RESULTS OF WEAR TESTS ON VARIOUS WELDING CONSUMABLES (KOHOPAA, HAKONEN ET AL.
1989)..................................................................................................................................................... 78FIGURE 6-7. WEAR RATE VARIATION FOR DIES WITH SHARP RADII AND FILLETS, FOR DIFFERENT COATINGS.
M11, M12 AND M-14 ARE CO-MO COATINGS, W2 AND W3 ARE CO-W COATINGS (STILL AND DENNIS
1977)..................................................................................................................................................... 82FIGURE 6-8. RESULTS OF SIMULATED HOT FORGING TESTS WITH DIFFERENT COATINGS (DENNIS AND JONES
1981)..................................................................................................................................................... 84FIGURE 6-9. VARIATIONS OF WEAR AREA WITH NUMBER OF FORGINGS. THE DIES USED WERE FLAT DIES WITH
DIES HAVING SHARP RADII AND FILLETS (DENNIS AND STILL 1975) ...................................................... 84FIGURE 6-10. RATIO OF CRACKED AREA OF COATED CORNERS TO AN UNCOATED CORNER FOR VARIOUS
MATERIALS (MIRTICH, NIEH ET AL. 1981) ............................................................................................ 85FIGURE 6-11 WEAR OF DIFFERENT THERMAL SPRAYED COATINGS (MONIKA 1981) ...................................... 86FIGURE 6-12 THE BURNISHED COATING DID NOT PRESENTED CRACKS. SAMPLES 45MM DIAMETER BY 40MM
HIGH, INDUCTION HEATED DURING ~ 18S AND COOLED BY 10S BETWEEN TEMPERATURES OF 20-700°C.(MONIKA 1981)..................................................................................................................................... 87
FIGURE 6-13. EFFECT OF LASER SURFACE MODIFICATION ON WEAR PERFORMANCE OF HOT WORK DIES
COMPARED WITH NITRIDED DIES (CSER, GEIGER ET AL. 1993) .............................................................. 87FIGURE 7-1 S-N CURVE WITH PROBABILITY LINES OR S-N-P (DIETER 1986) ............................................... 91FIGURE 7-2. ILLUSTRATION OF THE METHODS FOR ESTIMATING FATIGUE BASED IN STATIC PROPERTIES
(MANSON 1972) .................................................................................................................................... 92FIGURE 9-1. APPEARANCE OF PLOUGH MARKS CAUSED BY ABRASIVE WEAR (STACHOWIAK 1993) ............. 98FIGURE 9-2. DIFFERENT MECHANISMS OF WEAR IN ABRASION (STACHOWIAK 1993) .................................... 99FIGURE 9-3 .A) A TYPICAL METALLURGICAL WELD. B) A TYPICAL ADHESION JOINT (RABINOWICZ 1995). 100FIGURE 9-4. A): HOT FORGING TOP BLOCKER PUNCH MADE FORM H13. B) CROSS SECTION OF THE PUNCH C)
MOTTLED INTERFACE D) OXIDATION INSIDE OF THERMAL FATIGUE CRACK......................................... 101FIGURE 9-5. ILLUSTRATES PHYSICAL CHANGES ON THE DIE SURFACE THAT RESULTS IN HEAT CHECKING
(NORSTROM 1991) .............................................................................................................................. 103
VIII
FIGURE 9-6. ILLUSTRATION OF GEOMETRY EFFECT ON NORMAL UNI-AXIAL STRESSES REQUIRED TO INDENT A
SLAB (SCHEY 1987)............................................................................................................................. 106FIGURE 9-7. EXAMPLES OF SEVERE PLASTIC DEFORMATION AT THE DIE SURFACE (SUMMERVILLE,
VENKATESAN ET AL. 1995) ................................................................................................................. 107FIGURE 9-8. EXAMPLE OF SURFACE PLASTIC DEFORMATION (SUMMERVILLE, VENKATESAN ET AL. 1995). 107FIGURE 9-9. EXAMPLE OF SURFACE PLASTIC DEFORMATION (SUMMERVILLE, VENKATESAN ET AL. 1995). 108FIGURE 9-10 SCHEMATIC INTERACTION BETWEEN THE PARAMETER IN HOT FORGING AND THE CRACKING
(KNORR 1993)..................................................................................................................................... 109FIGURE 9-11 ILLUSTRATION OF A CRITICAL REGION IN EXTRUSION DIES, WHERE THE FILLET RADIUS IS
SUBJECT TO TENSILE STRESS (CSER, GEIGER ET AL. 1993) .................................................................. 109FIGURE 9-12. REPRESENTATION OF THE FATIGUE CRACK PROPAGATION (DIETER 1986)............................. 110FIGURE 9-13 TULSAN (TULSYAN, SHIVPURI ET AL. 1993) PRESENTS A CURVE FORM STOREN AND OTHERS
FOR DIFFERENT TOOL STEELS AND HEAT TREATMENT. A) FRACTURE TOUGHNESS PROPERTIES AS
FUNCTION OF THE WORKING TEMPERATURES AND THE HEAT TREATMENTS B) MATERIALS AND HEAT
TREATMENT LIST ................................................................................................................................. 110FIGURE 9-14 RESULTS IN AIR AND VACUUM ATMOSPHERES, SHOWING THE AMBIENT EFFECT AT THE FATIGUE
RESISTANCE IN HIGH TEMPERATURES (SALOMON 1972)...................................................................... 111FIGURE 9-15 CORRELATION OF HIGH AND LOW CYCLE FATIGUE DATA FOR SOLUTION TREATED TYPE 304
STAINLESS STEEL AS A FUNCTION OF ALTERNATING STRESS (SOO 1972). ............................................ 112FIGURE 9-16 EFFECT OF TEMPERATURE ON FATIGUE-CRACK-GROWTH BEHAVIOR OF 2 1/4 CR-1MO STEEL
(VISWNATHAN 1989). ......................................................................................................................... 113FIGURE 9-17 VARIATION OF FATIGUE-CRACK-GROWTH RATES AS FUNCTION OF TEMPERATURE AT ∆K =
30MPA (M)1/2 (VISWNATHAN 1989)..................................................................................................... 114
FIGURE 9-18 SERIES OF CASES WITH STRESS CONCENTRATION IN FORGING DIES PRESENTED BY KNORR
(KNORR 1993). A) – B) FROM ERLMANN AT AL.; C) -D) FROM MARECZEK......................................... 117FIGURE 10-1. ABRASION RESISTANCE OF SEVERAL TOOL STEELS VERSUS STRUCTURAL PARAMETER (WEAR
INDEX) (BLAU 1992) ........................................................................................................................... 118FIGURE 10-2. VARIATION OF WEAR INDEX WITH DIE HARDNESS AT ROOM TEMPERATURE (KANNAPAN 1969;
KANNAPAN 1970)................................................................................................................................ 119FIGURE 10-3. WEAR RESISTANCE OF .55% C DIE STEEL WITH HARDNESS, % CR AND HEAT TREATMENT. 1
INDICATES (KANNAPAN 1969; KANNAPAN 1970).............................................................................. 120FIGURE 10-4. WEAR TEST RESULTS USING DIFFERENT DIE MATERIALS (BRAMLEY, LORD ET AL. 1989) ..... 121FIGURE 10-5. WEAR TEST RESULTS USING DIFFERENT DIE MATERIALS (BRAMLEY, LORD ET AL. 1989) ..... 122FIGURE 10-6. VARIATION OF WEAR INDEX WITH DIFFERENT DIE STEELS. THE GRAPHS ALSO ILLUSTRATE THE
EFFECT OF DIFFERENT FORGING STEEL (THOMAS 1970) ...................................................................... 122FIGURE 11-1. EFFECT OF MAXIMUM CAVITY DEPTH ON DIE LIFE (HEINEMEYER 1976)................................ 124FIGURE 11-2. EFFECT OF NOMINAL LOAD AND ENERGY ON AVERAGE DIE LIVES (HEINEMEYER 1976)........ 124FIGURE 11-3. EFFECT OF FORGING WEIGHT ON DIE DAMAGE (ASTON 1969) ............................................... 125FIGURE 11-4. VARIATION OF DIE DAMAGE WITH SIZE OF FORGING (ASTON AND BARRY 1972) .................. 126FIGURE 11-5. EFFECT OF FORGING WEIGHT, FILLET RADII, DRAFT ANGLES AND CONTACT AREA ON WEAR OF
FORGING DIES (ASTON 1969)............................................................................................................... 127FIGURE 11-6. EFFECT OF VARIOUS TOOL STEEL ON DIE WEAR (THOMAS 1970) ........................................... 127FIGURE 11-7. EFFECT OF BULK TEMPERATURE AND STOCK TEMPERATURE ON WEAR OF HAMMER DIES
(THOMAS 1971)................................................................................................................................... 128FIGURE 11-8. RELATIVE DIE DAMAGE OF FIVE DIFFERENT PART FAMILIES WHEN FORGED IN A HAMMER AND A
PRESS (ASTON 1969) ........................................................................................................................... 129FIGURE 11-9. EFFECT OF DWELL TIME ON THE WEAR VOLUMES OBSERVED (ROOKS 1974) ......................... 130FIGURE 11-10. DIE WEAR FOR THREE DIFFERENT DWELL TIMES FOR A) H.50 DIES AND B) NO. 5 TOOL STEEL
DIES (ROOKS 1974) ............................................................................................................................. 131FIGURE 11-11. EFFECT OF SCALING TIME ON ADHESIVE WEAR CHARACTERISTICS (THOMAS 1971) ............ 132FIGURE 11-12: OXIDE FORMATION ON 080M40 (EN8) STEEL BILLETS HEATED TO 1100°C (DEAN 1974)... 133FIGURE 11-13. SCALE FORMATION AND ADHERENCE AS FUNCTION OF HEATING TIME AND FURNACE
ATMOSPHERE (THOMAS 1971)............................................................................................................. 133FIGURE 11-14. EFFECT OF FURNACE SELECTION ON DIE WEAR OF EXTRUSION DIES (DOEGE, SEIDEL ET AL.
1996)................................................................................................................................................... 134
IX
FIGURE 11-15. EFFECT OF SCALE THICKNESS ON THE DIE SURFACE TEMPERATURE (KELLOW, BRAMLEY ET
AL. 1969) ............................................................................................................................................ 135FIGURE 11-16. EFFECT OF FORGING TEMPERATURE ON THE WEAR DEPTH AFTER FORGING 4000 PIECES
(NETTHOFEL 1965).............................................................................................................................. 136FIGURE 11-17. VARIATION OF WEAR PATTERN OF THE TOP AND BOTTOM DIES WITH LUBRICATION (SINGH,
ROOKS ET AL. 1973)............................................................................................................................ 137FIGURE 11-18. VARIATION OF WEAR RATE WITH LUBRICATION (SINGH, ROOKS ET AL. 1973).................... 137FIGURE 11-19. VARIATION OF WEAR VOLUME WITH DIE BULK TEMPERATURE FOR LUBRICATED AND DRY
FORGING (SINGH, ROOKS ET AL. 1973) ............................................................................................... 138
X
LIST OF TABLES
TABLE 3-1. AISI CLASSIFICATION AND COMPOSITION OF TOOL STEELS (ROBERTS, KRAUSS ET AL. 1998) .. 17TABLE 3-2 LIST OF HOT WORK TOOL STEELS AND OTHER MATERIALS COMMERCIALLY AVAILABLE IN US... 18TABLE 3-3. COMPOSITIONS OF SOME COMMERCIALLY AVAILABLE HOT WORK STEELS ................................. 19TABLE 3-4 LIST OF HOT WORK TOOL STEELS AND OTHER MATERIALS COMMERCIALLY AVAILABLE IN US
TYPE...................................................................................................................................................... 19TABLE 3-5. CLASS 510. CHROMIUM DIE STEEL (ROBERTS, KRAUSS ET AL. 1998). ...................................... 19TABLE 3-6. CLASS 520. CHROMIUM – MOLYBDENUM DIE STEELS (ROBERTS, KRAUSS ET AL. 1998). ........ 19TABLE 3-7. CLASS 530. CHROMIUM – TUNGSTEN DIE STEELS (ROBERTS, KRAUSS ET AL. 1998)................ 19TABLE 3-8. CLASS 540. TUNGSTEN DIE STEELS (ROBERTS, KRAUSS ET AL. 1998)........................................ 19TABLE 3-9. CLASS 540. TUNGSTEN DIE STEELS (ROBERTS, KRAUSS ET AL. 1998)........................................ 19TABLE 3-10 TABLES WITH A COMPILATION OF CLASSIFICATION AND COMPOSITION FOR SEVERAL
SUPERALLOYS AND AGING ALLOYS; A-B (SEMIATIN AND LAHOTI 1981). .............................................. 38TABLE 4-1. CLEANLINESS OF STEELS USED BY ROBERTS AND NOSTRUM (ROBERTS AND NORSTROM 1987) 41TABLE 4-2. A, B HARDENING AND TEMPERING TEMPERATURES AND PROCEDURES FOR TOOL STEELS
(ROBERTS AND ROBERT 1980). ............................................................................................................. 60TABLE 5-1. RESPONSE OF DIFFERENT TOOL STEELS TO SEVERAL SURFACE ENGINEERING TOWARDS
ENHANCEMENT OF TOUGHNESS, HOT HARDNESS, HEAT CHECKING, TEMPER RESISTANCE (KRISHNADEV
1997)..................................................................................................................................................... 66TABLE 5-2. AVERAGE MAXIMUM WEAR DEPTHS (µM) ON SURFACE ENGINEERED DIES AFTER UPSETTING 500
AISI 1040 STEEL BILLETS AT 1070° C (VENKATESAN, SUMMERVILLE ET AL. 1998) ............................ 69TABLE 6-1. COMPOSITIONS OF VARIOUS GRADES OF NICKEL ALUMINIDES (BLAU 1992).............................. 73TABLE 6-2. WEAR CONSTANTS OBTAINED THROUGH PIN-ON-DISC TYPE TESTS FOR VARIOUS GRADES OF
NICKEL ALUMINIDES (BLAU 1992) ....................................................................................................... 75TABLE 6-3. SOME PHYSICAL PROPERTIES OF IC-50 (OAK ) ........................................................................... 76TABLE 6-4 VARIATION OF YIELD STRENGTH, ULTIMATE STRENGTH AND DUCTILITY OF IC50 WITH
TEMPERATURE (OAK ) ........................................................................................................................... 77TABLE 6-5. VARIATION OF MODULUS OF ELASTICITY OF IC50 WITH TEMPERATURE (OAK ) ......................... 77TABLE 6-6. WEAR VOLUME OBTAINED AFTER 100 FORGINGS USING FLAT DIES ELECTRO-DEPOSITED WITH
SOME WEAR RESISTANT COATINGS (STILL AND DENNIS 1977) .............................................................. 81TABLE 6-7. RESULTS OF INDUSTRIAL TRIALS OF USE OF COATINGS. 17A REPRESENTS NON-ROUND SHALLOW
DIES, 17B YOKE-TYPE DIES AND 17C GEAR BLANK DIES (STILL AND DENNIS 1977)............................... 82TABLE 6-8. RESULTS OF INDUSTRIAL TRIALS ON HOT FORGING DIES BRUSH PLATED WITH CO-MO ALLOY
COATINGS (DENNIS AND JONES 1981) ................................................................................................... 83TABLE 6-9 RESULTS OF PRODUCTION TESTING OF VARIOUS SURFACE TREATMENTS (MONIKA 1981) ........... 86TABLE 7-1. TABLE SUMMARIZING DIFFERENT WEAR MODELS FOUND IN LITERATURE................................... 90TABLE 9-1. RESULTS FOR CRACK PROPAGATION TYPO PARIS DA/DN FOR THE CONSTANTS “C, N”. B)
MATERIALS COMPOSITIONS FOR THE HOT TOOL STEELS USED (SCHUCHTAR 1988). ............................ 115
11
1. INTRODUCTION
Near net shape manufacturing processes are processes that produces product shapesclose to final desired shape. Closed-die forgings have traditionally had liberal stockallowances and high tolerances. With steady demands placed on the forgingcommunity to make forgings with lower stock allowances and higher quality, demandsplaced on die and the die material to last longer have become intense. Flashlessforgings with less than .03” machining stock, spur gears, spiral bevel gears and splineswith near net tooth with as little as .005” grinding stock and net toothed bevel gearshave become part of many forger’s product line.
These precision forged parts, apart from reducing material usage, reduces machiningtimes and provides better mechanical properties. With such small machiningallowances and tolerances, there is very little room for forging process variations. Theforgers have to reduce the process variations as much as possible so that the parts theymake meet or exceed the customer’s expectations.
Awareness of the importance of process control, well understood by aerospace materialforgers, is creeping into other steel forgers also. There is a strong need in the forgingindustry to reduce process variations and improve quality at the same time reduce costof forgings. This is essential to the survival of forging plants in the long run as well asviability of the new generation of precision forged parts.
One of the most important ingredients in cost of forgings is the cost of tooling involved.Die costs range from 10 to 15% of the cost of a forging. This is illustrated in Figure 1.1(Doege, Seidel et al. 1996). This includes cost of die material, machining the dies andsubsequent heat treatment, if necessary. The indirect cost of dies is however, far moresignificant.
If tooling wears out or become unusable, the production has to be stopped to changedies. Setup times can range anywhere from under 10 minutes to over 3 hours,depending on the complexity of the setup, skill and practices used by the setup crew.This results in additional direct wages in material handling, tool rework and otheroverhead costs. Also, this may result in additional overtime premiums in the die shopand the forge shop, low resource utilization and in an extreme case, result in misseddelivery to customers. If quality and inspection systems breakdown, if dies are notchanged at the appropriate time, additional loss occurs due to scrap. The effect oftooling failure on setup costs is shown in (Figure 1.1). Though tooling cost is only 10 -15% of a forging cost, the indirect cost of tooling could be as high as 70%. Life of a forgetooling, hence, has great ramifications on the economic competitiveness of a forgingcompany. Identifying different modes of die failure and understanding dominantmechanisms are essential first steps in the path to increasing die life.
12
Figure 1-1. (Cser, Geiger et al. 1993)
To understand the problems associated with forging die failure, one must understandthe forging processes and all the system components involved.
1.1. GOALS OF THE PRECISION FORGINGCONSORTIUM
The long term goals of the precision forging consortium is to provide the warm forgingoperations in North America, precision forging capability in all aspects of forgingoperation. In the short term, towards tracking progress in the direction of improvingperformance of dies essential to the viability of precision forging, the precision forgingconsortium has set for itself the following metrics.
1. 10X improvement in tool-life2. reduce die cost / piece by 15%3. 15% reduction in raw material consumption through precision forging capabilities4. Validating cost effective transition to lower forging temperatures5. 20% reduction in overall input energy
1.2. TASKS FOR PHASE I: GROUP FOR INCREASEDLIFE OF DIES/ OSU
The cornerstone of successful precision forging is development of high performancetooling that is cost effective from a overall product cost point of view as well as easilymanufacturable. Forgers needs to understand what factors affect the quality of theirparts and cost of their product. As we indicated earlier, one of the biggest componentsof cost is tooling cost and indirect costs of bad tooling including cost of additionalsetups, rework, scrap and loss of productivity. These costs will be substantially higherin precision forging with tighter tolerances and surface finish requirements. Currently,precision hot and warm forgers experience from 10-20 times more scrap (in ppm)
13
compared to conventional forging. Understanding and reducing die failure andimproving die life is an essential part of long term survival of forging industry.
To achieve these objectives, it is also necessary to evaluate die failure under a unifiedenvironment consisting of die materials, surface treatments, coatings as well as forgingprocess variables like temperatures, lubricants, cycle times and forging steels. Theeffect of process variables on die failure needs to be understood and modeled. Effect ofheat treatment and surface treatments on physical properties that control die failureneeds to be quantified. Thermo-mechanical conditions during forging need to be studiedand used as a starting point to predict failure. The goals Ohio States University’s groupfor increased die life set for itself towards achieving this objective were• Assemble and if necessary, generate necessary data and consolidate information
into a database• Assemble necessary information on coatings and surface engineering that have
potential use in precision forging• Develop correlation between die material properties and common failure modes• Develop models based on the database, to direct advances towards new die
materials
Towards these objectives, the group has performed the following tasks.• Reviewed the state of the art in die materials, coatings and surface treatment• Collected available information on properties that are necessary to predict die failure• Reviewed different fundamental failure mechanisms and appropriate models to
evaluate failure rates
In this report, the OSU team have reviewed the state of the art in materials, surfaceengineering techniques and advanced concepts that have either been tried in forging orhave the potential to improve die lives in precision forging. Appropriate information thatwas gathered is presented. Gaps have been identified that will be fixed in future eitherthrough surveys or through laboratory testing. A frame work for incorporating theavailable data into existing models has been proposed. The team will build on thisframework to create an Intelligent Software for Prediction of Die Failure.
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2. A BRIEF REVIEW OF FAILURE OF FORGINGDIES
Depending on the conditions of the process and the characteristics of the material andsurface conditions, one could encounter various modes of tool failure. These are:
• Wear (abrasive, adhesive and oxidation)• Thermal fatigue or heat checking• Mechanical fatigue• Plastic deformation
Of these, wear (abrasive and adhesive) and mechanical failures are the most commonforms of failure (Figure 2.1). Of the two mode of wear, abrasive wear is the morecommon form of wear. Adhesive wear is not very common in hot and warm forging ofsteels because of the presence of lubricant film and/or scales and oxide layer. It doesbecome a mode of die wear when the lubricant film is non-existent either because thereis no lubricant application or when excessive sliding and deformation thins thelubricant film. Good tooling design and material selection can overcome gross crackingand mechanical fatigue. Thermal fatigue, in almost all cases, serves as a catalyst toaccelerate abrasive wear.
The main physical phenomenon that control the abrasive wear in a metallic surfacesliding past another surface are relative sliding distance, normal pressure and hardnessof the surface. Design of forging dies, choice of forging and heating equipment, diematerial selection and surface treatments used have a tremendous effect on the wearcharacteristics as these factors affect one or more of the controlling fundamentalphysical phenomena. This relationship is illustrated in Figures 2.2 and 2.3.
Figure 2-1. Frequency and location of typical die failures in forging (Cser, Geiger et al.1993)
15
Figure 2-2. Complex interaction of forging parameters and wear Artinger, (Cser,Geiger et al. 1993)
Figure 2-3. Some aspects of forging and process design that affect wear and fractureLange, K in (Cser, Geiger et al. 1993)
16
From the above illustrations, the factors affecting die failure can be subdivided into• Tooling Issues – Die material selection, heat treatment, surface engineering, die
manufacture and design• Billet Issues – Billet preparation, steel type• Process Issues – Forging temperature, lubricant type and application, forging cycle
times and other forging practices
Effects of various process parameters and billet materials are described in Appendix A.The team felt that these, by careful choice of physical constants like heat transfercoefficients, friction factors and yield strengths obtained either through past work ornew but well understood testing, one can model and recreate the process using FiniteElement Method (FEM). FEM would provide forging designers stress-strain cycling,temperature history at a die location and sliding velocities – factors that cause diefailure. The relationship between these factors and rate of die failure are discussed insection 7. Section 3,4, 5 and 6 discuss material properties, heat treatment and surfaceengineering necessary to evaluate a materials capability to resist failure.
17
3. MATERIAL FOR FORGING DIES
Die material selection is possibly the biggest factor that affects the life of dies in a hot orwarm forging operation. There is a large variety of tool steels available in the marketthat can be used for hot and warm forging applications. These steels could becategorized as low alloy tool steels (Groups 6G, 6F, 6H), air-hardening medium alloytool steels (A2, A7-A9), chromium hot work steel (H-10 – H-19), tungsten hot worksteels (H20-H26), and molybdenum hot work steels (H41-H43).Selection of die material grade (steel composition and microstructure distribution) andsubsequent heat treatment play a key role in failure of dies. These properties completelydefine the thermal and mechanical properties that affect the mode of failure and therate of tool failure. In this section, we will go over the main classifications of tool steelgrades and characteristics of incoming tool steel – alloying composition, physical andmechanical properties. A short section will also discuss new non-steel basedsuperalloys available in the market.A comprehensive classification of tool steels by the American Iron and Steel Institute(AISI) is presented in tables 3.1 and 3.2 (Roberts, Krauss et al. 1998). The groups arebased on alloying elements and applications. Steels that are not temperature resistantare generally not used in making the forging dies. However, they used in other parts ofthe die set like the bolster and spacers.
Table 3-1. AISI classification and composition of tool steels (Roberts, Krauss et al.1998)
3.1. HOT WORK DIE STEELS
Hot work die steels are classified into 3 different categories (Roberts, Krauss et al. 1998)based on their alloy content. These can be:Chromium basedTungsten or Molybdenum basedSteels where tungsten and chromium are approximately in equal proportion
Most hot work steels are low carbon steels with medium or high alloying elements.Table 3.2 lists some of the more commonly used hot work steels. Table 3.3 lists some ofthe common grades of chromium die steels. Table 3.4, 3.5 and 3.6 lists some commongrades of Chromium – Molybdenum, Chromium-Tungsten, Tungsten and Molybdenumhot work steels (Roberts, Krauss et al. 1998).
18
COMMERTIAL NAME AISI C Ni Si W Cr V Mo Co OTHER
THYROTHERM 2343 EFSCARTECH 882CRU HALCOMB 218
H11 0.38 1.0 5.3 0.4 1.3
THYROTHERM 2344 ESFCARTECH 833; PLUSUDD OVAR SUPCRU NUDIE V; CPM-NUDIE EZFIN DC + XTRA
H13 0.40 1.0 5.3 1.0 1.4
FIN SHELLDIE - 0.36 1.0 5.0 0.3 1.85 Mn0.75
FIN SHELLEX - 0.36 0.9 5.0 0.25 2.85 Mn0.6
THYROTHERM 2365 EFS H10 0.32 3.0 0.5 2.8
THYROTHERM 2367 EFS 0.37 5.0 0.6 3.0
CARTECH 879CRU HALCOM 425
H19 0.4 0.3 4.25 4.25 2.1 0.45 4.25 Mn0.4
THYROTHERM 2581CRU PEERLESS A
H21 0.30 8.5 2.6 0.4
CRU CHRO-MOWTHYROTHERM 2606 EFS
H12H12
0.350.36
1.0 1.31.3
5.05.3
0.350.3
1.31.5
THYROTHERM 2713 6F2 0.55 1.7 0.7 0.1 0.3
THYROTHERM 2714 6F3 0.56 1.7 1.1 0.1 0.5
FINK DURODI (VI.F3) 0.55 1.55 0.5 1.00 0.8 Mn0.6
FIN FX-XTRA (VI.F2) 0.5 0.9 0.25 1.15 0.5 Mn0.85
THYROTHERM 2307CRU 4340
~43404340
0.310.4
-1.85
2.40.8
0.2 -
0.20.25
THYROTHERM 2742 0.56 0.5 1.0 0.1 0.4
THYROTHERM 2744 0.57 1.7 1.1 0.1 0.8
FIN PRESS-DIE - 0.2 3.25 0.25 3.35 Mn0.7
CRU MARLOK - 0.01 18.0 5.0 11.0 0.3Ti
THYROTHERM 2799 0.02 12.0 8.0 8.0 0.5Ti
THYROTHERM 2885 EFSCRU WR95
H10AH10M
0.30.35
3.03.5
0.50.6
2.82.5
3.02.0
THYRODUR 2379CARTECH 610UDD SVERKER 21 PM
D2
~D2
1.55 1.0 12.0 1.0 0.7
CARTECH 880CRU CRUCIBLE A9
A9 0.5 1.5 1.0 5.0 1.0 1.4 Mn0.3
CARTECH EXTENDO-DIE - 0.44 1.0 6.0 0.8 1.9 Mn0.45
CARTECH PYROTOUGH - 0.4 4.45 0.8 2.05 Mn0.45
CARTECH DURA-FORM - 0.65 1.4 4.0 1.5 2.5 Mn0.5
CARTECH PYROTOOL V - 0.04 27.0 0.25 14.5 0.2 1.25 Mn0.25Ti 3.0
CARTECH AERMET - 0.23 11.1 3.0 1.2 13.4
CRU CPM V3 - 0.8 7.5 2.75 1.3
CRU CPM 9V - 1.78 5.25 9 1.3
UDD VANADIS 4 - 1.54 0.09 0.91 8.03 3.9 1.53 Mn0.32
FIN WF-XTRA - 0.42 0.8 0.5 2.5 0.08 1.00 Mn075
UDD QRO 90 SUPREME - 0.39 0.3 2.6 0.8 2.3 Mn0.75
Table 3-2 List of hot work tool steels and other materials commercially available in US
19
Table 3-3. Compositions of some commercially available hot work steels
Table AISI C Mn Si W Cr V Mo510 .95 0.30 0.30 4.0511 .95 0.30 0.30 4.0 0.50 0.50512 .60 0.30 0.30 4.0 0.75 0.50513 S7 .50 0.70 0.30 3.25 1.40514 .50 0.30 0.90 3.25 0.25 1.40
Table 3-5. Class 510. Chromium die steel (Roberts, Krauss et al. 1998).
Type AISI C Mn Si Cr Ni V Mo W520 H-11 0.35 0.30 1.00 5.00 0.40 1.50521 H-13 0.35 0.30 1.00 5.00 1.00 1.50522 H-12 0.35 0.30 1.00 5.00 0.40 1.50 1.50523 0.40 0.60 1.00 3.50 1.00 1.00 1.25524 H-10 0.40 0.55 1.00 3.25 0.40 2.50525 0.35 0.30 1.00 5.00 2.00
Table 3-6. Class 520. Chromium – Molybdenum die steels (Roberts, Krauss et al.1998).
Type AISI C Mn Si Cr V W Mo Co530 H14 0.40 0.30 1.00 5.00 0.25 5.00 0.25 0.50531 H19 0.40 0.30 0.30 4.25 2.00 4.25 0.40 4.25532 0.45 0.75 1.00 5.00 0.50 3.75 1.00 0.50533 0.35 0.60 1.50 7.25 7.25534 0.45 0.60 1.50 7.25 7.25535 H16 0.55 0.60 0.90 7.00 7.00536 H23 0.30 0.30 0.50 12.00 1.00 12.00
Table 3-7. Class 530. Chromium – Tungsten die steels (Roberts, Krauss et al. 1998).
Type AISI C Mn Si Cr Ni V Co W Mo540 H21 0.35 0.30 0.30 3.50 0.50 9.00541 H20 0.35 0.30 0.30 2.00 0.50 9.00542 0.30 0.30 0.30 2.75 1.75 0.30 10.00 0.25543 H22 0.35 0.30 0.30 2.00 0.40 11.00544 0.30 0.30 0.30 2.50 0.40 3.60 12.00545 H25 0.25 0.30 0.30 4.00 1.00 15.00546 0.40 0.30 0.30 3.50 0.40 14.00547 H24 0.45 0.30 0.30 3.00 0.50 15.00548 0.35 0.30 0.30 4.00 2.50 14.00 2.00549 H26 0.50 0.30 0.30 4.00 1.00 18.00
Table 3-8. Class 540. Tungsten die steels (Roberts, Krauss et al. 1998).
Type AISI C Mn Si Cr Ni V W Mo Co550 H15 0.35 0.30 0.40 3.75 0.75 1.00 6.00551 H15 0.40 0.30 0.50 5.00 0.75 1.00 5.00552 H43 0.55 0.30 0.30 4.00 2.00 8.00553 H42 0.65 0.30 0.30 4.00 2.00 6.40 5.00554 H41 0.65 0.30 0.30 4.00 1.00 1.50 8.00555 0.30 0.50 0.30 3.00 3.00556 0.10 0.30 0.30 3.50 0.50 4.00 5.00 25.00
Table 3-9. Class 540. Tungsten die steels (Roberts, Krauss et al. 1998).
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3.2. PHYSICAL AND MECHANICAL PROPERTIESOF VARIOUS TOOL STEELS
In modeling, analyzing and predicting die failure, the knowledge of the physical andmechanical properties is very important. Knowledge of these properties is necessary toboth understand the reasons for die failure as well as perform forging simulations byfinite element methods (FEM). Tool steels used for hot forming should possess thefollowing properties.
3.2.1 RESISTANCE TO DEFORMATION AT HIGHTEMPERATURES
(a) (b)
Figure 3-1. (a) Variation of hardness of die steel of H-12 and 6F2 with temperingtimes. H-12 used was austenetized from 1040 C with as quenched hardness of 584VPN, 6F2 was austenetized from 850 °C with as quenched hardness of 601 VPN (b)
Variation yield strength of different tool steels with temperatures. (Nagpal 1976)
21
a) b)
Figure 3-2. Yield strength and hardness versus temperature for several tool steels(Norstrom, Johansson et al. 1981)
Figure 3-3. Variation of yield strength with temperature (Semiatin and Lahoti 1981)
22
Figure 3-4. Mechanical resistance versus test temperature expressed in two forms forthe H13 and a new hot work tool steel (QRO90) (Roberts and Norstrom 1987)
a) b)
Figure 3-5 Yield strength and ductility versus test temperature (A5) elongation (Z)area reduction. a) H13 and b) QRO90 (Uddeholm )
23
H12 H11
H10 H21
6F3
Figure 3-6 Tensile strength and ductility versus test temperature for steels from H13group (H12, H11, H10), plus and high alloy, H21, and a low alloy group 6F3 (Thyssen
).
24
3.2.2 RESISTANCE TO MECHANICAL SHOCK ANDFATIGUE
Figure 3-7. Variation of Charpy toughness with different hardness levels and testingtemperatures on of hot work die steels (values in parentheses indicate hardness at
room temperature) (Nagpal 1976)
Figure 3-8 Variation of toughness for several tool steels in function of hardness andyield strength (Cser, Geiger et al. 1993)
25
Figure 3-9. Variation of toughness for several tool steels in function of hardness andyield strength (Shivpuri and Semiatin 1988)
Figure 3-10 Comparison of toughness properties for H13, H21 and a new hot worktool steel QRO80M versus test temperature (Johansson, Jonsson et al. 1985).
26
a) b)
c)
Figure 3-11 Comparison of toughness KIC and Charpy V-notch for several tool steels.a) KIC for three bar size, longitudinal direction, the small diameters represent
reduction from the bigger diameter; b) Charpy V-notch longitudinal and transversal c)KIC for H13 in function of the hardness, austenitized at 1024°C, time 25 min, air
cooled (Hemphill and Wert 1987).
a) b)
Figure 3-12 Toughness versus test temperature for a) H13 and b) QRO90 (Uddeholm )
27
3.2.3 RESISTANCE TO THERMAL SOFTENING
(a) (b)
Figure 3-13.(a) Variation of hardness with temperature for H-11, H-12, H-13, H-14and Pyrovan. Measurements were made after holding samples for 30 minutes for
homogenization (b) Resistance of hot work dies steels to thermal softening asmeasured by the room temperature hardness (Nagpal 1976)
a) b)
Figure 3-14 Comparison of properties for H13 and a new hot work tool steel QRO80Mversus test temperature (a) standard tempering curve (Johansson, Jonsson et al.
1985). (b) Master tempering curve for premium H13 were: P= Larsom-Millerparameter, T is temperature (°F), t is time in hours (Carpenter )
28
Figure 3-15 Thermal expansion for several tool steels versus temperature (Roberts,Krauss et al. 1998)
29
3.2.4 DUCTILITY
Figure 3-16. Ductility of various die steels at high temperatures (Nagpal 1976)
Figure 3-17 Comparison of ductility for H13, H21 and a new hot work tool steelQRO80M versus test temperature (Johansson, Jonsson et al. 1985).
Apart from these, because of practical reasons, they need to possess good machinabilityand resistance to warping during heat treatment. Die material’s resistance to plasticdeformation depends on how well it retains its hardness with temperature. It alsodepends on its yield strength. Resistance to mechanical shock relies on the materialhaving good fracture toughness commonly measured in Charpy V-notch testing units.
30
Resistance to wear depends on tempering resistance with temperature indicated byhardness measurements at elevated temperatures. Resistance to heat checkingdepends on the material having high ductility, good tempering resistance, high yieldstrength and low thermal expansion. High heat conductivity and low thermal expansioncoefficient in die materials is desirable because it reduces the temperature gradient orassociated thermal strains that is the cause of thermal fatigue and shock. It is alsodesirable that the steel retains all its properties for an extended period under elevatedtemperatures. The resistance of a die steel to thermal softening mainly depends on itsalloying constituents and its distribution. The tempering characteristics of these toolsteels obtained under laboratory condition represents very well the die material’sresistance to thermal softening.
Section below shows some critical physical and mechanical properties of hot work steelsthat impact one or more of the properties listed above. Information in tables figure 3.1to figure 3.10 have been compiled from a variety of sources.
31
3.3. MARAGING AND OTHER STEELS
Maraging steels are relatively new group of steels that was primarily developed foraerospace applications. It has high nickel, cobalt and molybdenum content but verylittle carbon. After austenitiztion and quenching the steel, the structure is soft nickelmartensite or similar soft structure with typical hardness of 30 – 40 Rc. Aging thismatrix at temperatures around 500° C results in dispersed precipitation of intermetallicphases. This precipitation is not concentrated at the grain boundary alone. Thisdramatically increases the strength without unduly affecting the toughness. Its highresistance to thermal shock and high toughness makes it a good candidate for dieswhere the mode of failure is heat checking. Maraging steels, used in die castingindustry, is not very common in forging industry.
3.3.1 COMPOSITION
Type Ni Co Mo Ti Al C * Si* Mn* S * P*I-VascoMax C-200 18.5 8.5 3.25 0.2 .1 .03 .10 .10 .01 .01II- VascoMax C-250 18.5 7.5 4.8 0.4 .1 .03 .10 .10 .01 .01III-VascoMax C-300 18.5 9 4.8 .6 .1 .03 .10 .10 .01 .01IV-VascoMax C-350 18 11.8 4.6 1.35 .1 .03 .10 .10 .01 .01HWM (+) 2 11 7.5 - - .05 .10 .10 .01 .01X2NiCoMoTi 12 8 8Thyrotherm 2799
12 8 8 .5 .5 .03 .10 .10 .01 .01
Marlock(Cr0.2) 18.0 11.0 5.0 0.3 0.01 0.1 0.01 0.01
Table 3-7 Composition of common maraging steels, VascoMax is a trade name ofTeledyne, (*) indicates maximum allowed content, (+) trademark of Crucible steel
3.3.2 PROPERTIES
a) b)(contd.)
32
c)
Figure 3-18 (a) Master tempering curve, T is temperature in Kelvin, t is time in hours(b) hot hardness of HWM compared to H- 13 (Kasak and Steven 1970) (c) aging curves
a) b)
Figure 3-19 Properties versus test temperature of maragins steels compared with H13(a) hardness (b) Yield strength and ductility (Bayer 1984)
33
a) b)
c)
Figure 3-20 Properties on a maraging steels 18 Ni 300 and H13. (a) Hot-hardness ofmeasured after holding at the testing temperatures for 30 min. (b) tensile propertiesversus testing temperatures c) toughness of versus test temperature (Barry, Wills et
al. 1968)
34
a) b)
c) d)
Figure 3-21 Properties of a hot work maraging steel compared with H13 group andthermal fatigue results: a) fields of austenite and martensite in function of
temperature showing the effect of Ni% (Brandis and Haberling 1987) (b) aging curvesfor 18%Ni and 12%Ni c) thermal fatigue resistance for H13 group and a maraging
steel (d) Hot yield strength (Gehricke 1993; Gehricke, Klarenfjork et al. 1995)
Figure 3-22 Comparison of toughness Charpy V-notched for a H13 steel and amaraging steel (Marlock) at two test temperatures (Dorsch 1991)
35
3.4. SUPERALLOYS
Nickel, cobalt and iron based superalloys are another group of die materials that hasexcellent potential in hot precision forging. This group of materials have extremely hightemperature strength and thermal softening. Like maraging steels, this group ofmaterials gets its strength from precipitation strengthening of intermetallic compoundslike Ni3Al. Tables 3.8 provide a comprehensive list of superalloys and its composition.There are 4 primary group of superalloys. They are:
Iron-based alloys. This group comprise of die steels like H-46 and Inconel 706 andcontain over 12% of Chromium. Small amounts of Molybdenum and Tungsten providethe matrix with high temperature strength. Iron based superalloys also includeaustenitic steels with high chromium and nickel content. This group can be used inapplications where dies could heat up to 1200°F.
Nickel-Iron based alloys. This group of alloy contains 24-27% nickel, 10-15%chromium and 50-60% iron along with small quantities of Molybdenum, Titanium andVanadium. The carbon content in these alloys is very small, typically less than .1%.Nickel based alloys. This group of alloys contains virtually no iron. The primaryconstituent of these alloys are nickel (50-80%), chromium(20%) and combination ofmolybdenum, aluminum, tungsten, cobalt and columbium. These grades again, gettheir strength from solid solution strengthening and can be put to service attemperatures up to 2200° F. Example of nickel-based superalloys are Waspalloy,Udimet 500 and Inconel 718.
Cobalt based alloys. This group of alloys are more ductile than the other groups.Again, these are age hardenable alloys whose primary constituents are Nickel, Iron,Chromium, Tungsten and Cobalt. These can be used in applications where it couldreach 1900° F.
a) b)
Figure 3-23. Ductility and toughness versus test temperature (Semiatin and Lahoti1981)
36
Figure 3-24. Mechanical properties expressed as hardness and Yield strength versustest temperature (Semiatin and Lahoti 1981)
a) b)
c) d)
Figure 3-25. Compilation of several properties versus test temperature for superalloysfrom Ohuchi (Ohuchi 1990). a) hardness b) thermal expansion c-d) Ultimate tensile
strength and Yield stress.
37
a)
38
b)
Table 3-10 Tables with a compilation of classification and composition for severalsuperalloys and aging alloys; a-b (Semiatin and Lahoti 1981).
39
4. DIE BLOCK MANUFACTURING AND HEATTREATMENT
4.1. DIE BLOCK MANUFACTURING: CLEANLINESSAND MICROSTRUCTURE
The first step in ensuring die quality is to make sure tool steel that is purchased hasthe correct composition and structure. Two of the main properties that dictate theperformance of a die material are its ductility and the toughness. These properties areaffected in varying degree, by the quality of tool steel. Figure 4.1 shows that thecleanliness of the steel has a very pronounced effect on the ductility of the steel. On thecontrary, the heat treatment the steel is subjected to has a big impact on the resultingtoughness. The charpy V-notch values indicate the toughness measured and theunnotched test measures the ductility of tool steel. The steel making process has astrong effect on the following:• Cleanliness of the steel produced, number and size of non-metallic inclusions
• Eutectic carbide size and number
• Microbanding and segregation of alloying elements
There are several tool steel making processes used currently like conventional orelectric arc, vacuum arc remelting (VAR) and electro-slag remelting (ESR) processes.These processes are capable of producing at different levels of cleanliness. In general,oxides and sulfides are detrimental to the toughness and ductility of tool steelproduced. Figure 4.2 illustrates the effect of sulfur content on the transverse fracturetoughness of H-13 dies. Low levels of oxygen are achieved using vacuum degassing andadvanced deoxidization methods. Any resulting oxides are reduced in size bysubsequent electro slag remelting process. Smaller inclusions are less detrimental tothe ensuing mechanical properties. Low levels of sulfur are achieved via ladle-refiningtechniques and / or electro-slag remelting (ESR);
Figure 4-1. Schematic comparison of ductility (Charpy unnotched) and toughness(Charpy V notch) versus temperature, room temperature (RT) (Norstrom 1989)
40
Figure 4-2. Effect of sulfur content on the transverse fracture toughness of H-13 diesteel (Roberts and Norstrom 1987).
Primary or eutectic carbides can reduce the transversal ductility if present in sufficientsize and amount. This carbide forms in the last stages of solidification an elongateduring the hot work reducing the ductility and toughness, specially in the transversesection. Figure 4.3a shows the effect of coarse grain boundary carbides in annealed toolsteel on the toughness of H-13 after it is heat-treated. Here, it shows that the carbideinclusions were too large to be dissolved during the austenitizing phase. Figure 4.13bshows the dramatic effect of cleanliness of die steel and mechanical processingperformed on the ingot, on the final toughness of H-13 and H-11 hot work tool steel.The figure 4.4a illustrate how the amount of carbides reduce the ductility in thestandard H13; notice that the compositions are similar and the standard H13 has lowlevel of oxygen and sulfur. Figure 4.4 (b) illustrates the effect of inclusions and carbidesfor two premium H13 grades (OMS1,2) and a standard H13. The annealedmicrostructure of all three grades (inclusion content provided in Table 4.1) wereconsidered acceptable by standards established by various organizations like Chrysler(Chrysler-NP 2080) and German tool steel specification(VDG Datasheet M82). It is againnoted that the ductility was greatly reduced by the inclusions in the transversaldirection.
a) b)
Figure 4-3. Effect of coarse grain boundary carbides from annealed tool steel on thetoughness at elevated temperatures. Measurements were made after quenching andtempering (Becker, Fuchs et al. 1989) b) Combined influence of clean processing and
extra fine structure (EFS) on toughness (Becker 1984).
41
a) b)
Figure 4-4 (a) Effect of carbides on ductility of standard H-13 (b) Effect of carbidesand inclusions on ductility. (Roberts and Norstrom 1987)
Sulphides Alumina Silicates Globularoxides
Number ofcarbidesOMS1 0 0 0 1 11
OMS2 0 1 0 1 8
H13 0.5 2.5 0 1.5 99
Table 4-1. Cleanliness of steels used by Roberts and Nostrum (Roberts and Norstrom1987)
Microbanding and segregation are features found in annealed die steel under an opticalmicroscope that indicates segregation of alloying elements. Annealed structures wereclassified by European and American associations in tables that define what isacceptable and what is not. These classifications base the criterion for rejection onpresence of acicular structure, grain size and banding. However, Roberts and Norstrom(Roberts and Norstrom 1987) and Kogler and Schindler (Kogler, Breitler et al. 1989)showed that these classifications can be inaccurate. They concluded that the “acicular”appearance is due to the carbide distribution resulting from a bainitic structure prior toannealing. They also concluded that banding that reflect segregation of the principalalloying elements Cr, Mo, and V, do not necessarily translate into poor ductility afterheat treatment. However, if eutectic carbides line up along the bands, it will reduce thetoughness and the ductility of the tool steel. Becker (Becker, Fuchs et al. 1989) showsthe effect of long-term diffusion annealing on reduction of micro-segregation. He foundthat by improving the isotropy he could achieve better toughness and ductility in thetransverse direction.The crack initiation by thermal or mechanical stress-strain loading is directlydependent on the ductility. The crack growth depends of the toughness, however betterductility also reduces the crack growth. Details of the mechanism are provided in thesection on thermal fatigue. Figure 4.5 clearly shows the relationship between ductilityand thermal fatigue. Both the grades A and B have the same chemical composition.Samples were taken in three orthogonal directions (transverse, longitudinal and secondtransverse). The table clearly shows that the lower the ductility, lower is the thermalfatigue resistance.
42
a) b)
Tensile in S directioncenter
Absorbed energy VWCenter (ft.lbs) ductility
orientation
Heatchecking rate
N0/in amax ∑a103in
Bar Rp0.2(Ksi)
A5 Z L T S
A 174 6.6 16 169 102 52 LTS
81015
355635480
182330
618983
B 180 10 40 216 213 212 LTS
568
230355330
111917
316549
c)
Figure 4-5 Relation between samples orientation, ductility and thermal fatigue A5
elongation, Z area reduction, VW unnotched European sample for impact test a,b,c(Roberts and Norstrom 1987)
43
4.2. CAVITY MANUFACTURE: MACHINING ANDEDM
Die manufacturing techniques, to some extent impact the fatigue and wear life offorging dies. Most round dies are turned on CNC lathes. Dies that are semi-cylindricalare semi-finished on lathes and finished on mills or EDM machines. Non-round diesare either machined on mills or burned using wire or plunge EDM machines. Dies maybe ground or polished after machining.Cavity manufacturing does not have as big an impact on the performance of a precisionforging die. However, there are a few issues that need to be taken into account in themanufacture of cavity, that may affect the life and quality of dies made. They are:1. Grinding feeds and speeds2. Load or current levels used in EDM3. Machining capabilitiesThere are some studies illustrating the effect of incorrect choice of grinding parameters(Roberts and Robert 1980). Roberts notes that if the grinding thickness in a singlegrinding pass is high, it may result in thermal cracks on the surface. If these dies areput into service in a hot working environment, it may result in premature failure of dies.Malm (Malm, Svensson et al. 1979) reports that a rougher surface causes a highersurface crack density but does not result in higher crack depths or crack propagationrate in thermal fatigue tests. Grinding can also form a soft re-tempered surfacealthough literature is poor in quantifying this effect.Incorrect choice of EDM parameters causes surface defects like micro cracking, whitelayer, and melted regions on the surface of dies. These defects were shown to decreasethermal fatigue resistance in a series of work (Young 1968; Suzuki, Ishihara et al.1972; Young 1979; Wallbank and Phadke 1982; Nichols and Dorsch 1984; Becker,Fuchs et al. 1989; Centa and Wallace 1989; Dorsch 1991; Gehricke 1993; Kim andWallace 1994; Schwam and Wallace 1995; Venkatesan, Subramanium et al. 1997).Figure 4.6 illustrates the drop in hardness of the refrozen layer with increase in thecurrent used.
Figure 4-6. Effect of EDM process in die surface; refrozen layer and thermal cracks(Cser, Geiger et al. 1993)
44
Although literature is poor in quantifying this effect, as EDM surface weakens with thesurface defects, it reduces the dies’ wear resistance (Wallbank and Phadke 1982). Ithas also been proven that the EDM layer (refrozen, over-tempered, brittle) reduces thethermal fatigue resistance (Kim and Wallace 1994). As the thermal fatigue cracks form,the die surface layer becomes more prone to wear. The Figure 4-7 exemplifies damagedlayer (cracks and refrozen) caused by EDM process.Dorsch (Dorsch 1991) showed that for the H13 and H10-A the layers under the “whitelayer” is softer due to over-tempering. Rough EDM process forms a overtempered layerthat is approximately 0.003in thick. This becomes thinner with finishing EDM. Theauthor shows hardness loss in a range of 2-10HRC in the over-tempered layer.
Figure 4-7. Surface of dies after EDM, H13 and maraging steel (Marlock) (Dorsch1991).
When a maraging steel (Marlock) is aged before EDM, its subsurface also loses somehardness (about 10 HRc). However, when a maraging steel die is aged after EDM, theloss in the hardness of the subsurface is only 2 Rc. Its white layer is soft and ductilebecause it is in a solution condition. The surface cracks are fewer in number. Whenthe dies are aged after EDM, the subsurface and the white layer reach harnesses within2-3 HRc of the parent material. The surface of EDMed maraging steel dies are lotcleaner compared to a EDMed H-13 or H-10 die, Figure 4.7.
Walbank (Wallbank and Phadke 1982) shows results from fatigue tests thatdemonstrate the reduced fatigue resistance of EDM specimen. The use of multiple stepsof decreasing EDM energy, tempering, and mechanical grinding to remove the affectedrecast layer are different ways to avoid the loss in thermal fatigue resistance of EDMeddies.
High performance machining is another alternative to EDM. Traditional machiningtechnology is not capable of machining die steel hardened to 445 Rc or above. Withnewer cutting tools available in the market, it is now possible, to machine pre-hardeneddie blocks. By machining pre-hardened die blocks, one can avoid heat-treat distortionsof the die cavity. Also, this reduces the lead times involved in die making.
45
4.3. HEAT TREATMENT: AUSTENIZING,QUENCHING AND TEMPERING
Precision forging process subjects the dies to extreme conditions of pressure,temperature variations. It is our objective to achieve the best combination of properties– toughness, ductility and hardness to both resist the wear and thermal fatigue on diesat the same time reduce the chance of catastrophic failure. Once a die material ispurchased to specification and the die cavity manufactured, the heat treatment it issubjected to has a dramatic effect on the properties of the die material. Understandingthe effects of various elements of heat-treatment should guide the heat-treatspecification on the die. Our goal in this report is to discuss the various issues involvedin heat treatment from a hot forging perspective.Before selection of a die material for an application, it is imperative to know what amaterial is capable of and the performance levels we can aspire for. Alloying elementslike Chromium, Nickel, Vanadium and Molybdenum play a dramatic role in determiningthe range of physical and mechanical properties one can expect from the material. Ifproperly heat-treated, high alloy die materials are capable of delivering high wearresistance by retaining its hardness at higher temperatures. In general, if a lower alloydie material is used, for similar toughness, the wear resistance that one can expectwould be lower. Any attempt to heat treat the material to a higher hardness couldpotentially lead to catastrophic failure of dies by reducing the toughness. On the otherhand, if we do not make use of the wear resistance a die material is capable of by heat-treating it to high toughness values, we may not be utilizing the full potential of thematerial.Heat treatment of die steels involves the following steps:
1. After dies are made, it is heated to austenitizing temperature. Austenitizingtemperatures for hot work tool steels range anywhere from 1000° C - 1500° C. Duringthis phase, the structure of steel transforms from ferrite-pearlite structure to austenite.2. The dies are held at these temperatures for an extended period. This is the “soak” or“hold” time. During this stage, the structure becomes uniformly austenitic. Carbides ofalloying elements go into solution.
3. After soaking, the dies are quenched in a quench medium to temperatures below thetransformation temperature. During this phase, based on the cooling rate differentregions of die experiences, transforms into different phases. Martensite is the ideal finalstructure, however in practice lower bainite, upper bainite, pearlite or retainedaustenite can be present in the structures, specially in blocks with big section.
4. Tempering is the next stage of heat treatment. Here, martensite formed as a result ofquenching is tempered to a tougher structure. This could be done in more than onesteps to maximize the toughness achieved, without sacrificing hardness.These stages in heat treating a tool steel die is illustrated in Figure 4.8.
46
Figure 4-8. Heat treatment cycle of hot working steels (Krauss 1995).
4.3.1 AUSTENITZATION AND SOAKING
Austenitization is the phase of heat treating tool steels where structure of steel isconverted to austenite. This is typically done in one or 2 stages and is performed in aaustenitizing furnace. Issues that affect the physical and mechanical properties duringaustenitization phase are as follows-
• Number of stages and rate of heating
• Soaking temperature or the maximum austenitizing temperature
• Soaking or holding time
• Furnace and atmosphere
The first step of the hardening is austenitization, a two-phase region that containaustenite and undissolved carbides. The austenitizing temperature and the time at thetemperature will determine the amount of carbides dissolved in the austenite andconsequently its composition. Higher austenitizing temperatures result in dies withlower amounts of primary carbides. Dissolved carbides enrich the austenite structurewith carbon and alloying elements reducing primary carbides. During quenching, thesecarbides in solution can precipitate as grain boundary carbides reducing the toughness.Higher austenitizing temperatures move the carbon precipitation line to the leftincreasing their amount for the same quenching rate. Figure 4.9 illustrates thereduction in toughness with austenitizing temperatures.
Enriched austenitic matrix result in a carbon rich martensite, which is stronger.Hence, higher austenitic temperatures result in a structure with higher yield strength.Figure 4.10 shows the effect of austenitizing temperatures and times on the asquenched hardness of a H-13. Figure 4.10 also shows the grain size as a function ofsoaking temperature and holding times. It is obvious that the effect of temperature onthese parameters is more pronounced than the effect of time. Equation 4.1 presented inStuhl (Stuhl and Breitler 1987) also illustrates this point clearly.
)log24( tTHP += Equation 4-1
-HT is the hardening parameter-T is the hardening temperature in Kelvin-t is the time at the hardening temperature in minutes
47
Enriched austenitic matrix result in a carbon rich martensite, which is stronger.Hence, higher austenitic temperatures result in a structure with higher yield strength.Figure 4.9 and 11 show the effect of austenitizing temperatures of the as quenchedhardness of some typical hot work die steels. Figure 4.9 also illustrates the higher yieldobtained as a result of increasing the austenitic temperatures. The enriched matrixalso lowers the Ms temperature below which martensite is formed. This increases thetendency to form retained austenite on quenching. Figure 4.12 shows the effect ofcarbon on the Ms temperatures. Higher retained austenite may necessitate moretempering time or temperature or both. For precision hot forging applications theretained austenite is harmful because it can transform into undesirable phases duringthe die working. Retained austenite transforms into cementite and ferrite, oruntempered martensite, with the increase in the tempering temperature. These phaseslead to toughness reduction because untempered martensite is fragile and theformation to cementite and ferrite can produce elongated interlath carbides detrimentalto the toughness (Krauss 1995).
Higher temperature and longer soak times also increase the prior austenite grain size.Longer soak times also increases the risk of decarburization at the surface. Larger prioraustenite grain size will result in larger grain sizes in the heat-treated dies. Bigger grainsize, in general is considered detrimental to the strength and toughness of a matrix.Smaller austenitic grain size also increases the toughness by shifting the DBTT to theright.
48
Figure 4-9. Influence of austenitizing temperature in properties versus temperingtemperature; Tensile strength, ductility and toughness are represented for H13 and
Dim 2367 hot work steel (Becker, Fuchs et al. 1989).
The rate of heating and the number of stages used to austenitize is typically controlledso that the core and the surface heats up uniformly. This reduces distortion andmaintains uniform microstructure throughout the die block. It is desirable to equalizethe die temperature at lower temperatures where the grain growth is minimal. If oneattempts to equalize the dies after they have reached the normalizing temperatures, thegrain size at the surface could get too large compared to the core grains. Also, it isdesirable to equalize the die block during transformation (around 1500-1600° F). Again,this ensures uniform structure and low distortion.
There are several furnaces available for austenitizing. Austenitizing furnaces could beone of the following.
• Vacuum Furnaces
49
• Atmosphere-sealed furnaces• Fluidized bed• Salt bath
Vacuum furnaces can lead to uneven heating and distortion, but decarburization andoxidation is very well controlled. Because of the uneven heating rates, if one requiresthe dies to be uniform, the heat treater must employ multiple thermocouples to tracktemperatures of core and surface. Atmosphere-sealed gas furnaces uses nitrogenand/or endothermic gas to rapidly heat the dies. Because heat is transferred byconvection, heat transfer is uniform. Fluidized bed furnaces heat the dies bysuspending them in a bed of fluidized gas. Heating is uniform and fast, but the surfacecould become carburized. Salt bath furnaces suspend the dies in molten salt. Again,this process provides uniform heating but could corrode, oxidize or decarburize thesurface.
Figure 4-10. Effect of austenitizing temperatures on ASTM grain size and as-quenchedVickers hardness of H-13 (Stuhl and Breitler 1987).
50
Figure 4-11 Effect of austenitizing temperatures on as quenched hardness, grain sizeand retained austenite a) H13. b) H13 and H11 (Pickering 1987)
a)
b)
Figure 4-12. CCT diagrams for two austenitizing temperatures , b) Effect of carbon onMs (Roberts and Robert 1980)
51
4.3.2 QUENCHING
Once dies are austenitized and held at an austenitized temperature, dies are quenchedusing a quenching medium in a quenching furnace. The rate of quenching and theuniformity of temperature in the block has a large effect on the resulting microstructureand mechanical properties. The uniformity of temperatures from the surface to the corehas a profound effect on the distortion experienced by the dies. The following sectionwill discuss various issues in quenching that affect the properties of dies and itsperformance in the field.
In an ideal case, quenching a small sample of tool steel will result in the structuretransforming to martensite. Martensite is the most desirable form of microstructurethat one can aspire for in a as-quenched die. A martensitic as-quenched structurecould be tempered to give the best combination of toughness and wear resistance in adie. One could obtain a martensitic structure if the cooling rate is fast enough to avoidless desirable phases like lower bainite, pearlite and ferrite. The different pathsresulting from different cooling rates resulting in different phases, are represented inFigure 4.13. However, because of various reasons, this becomes both impractical andundesirable.
Figure 4-13. a) Effect of bar size on the quench rate and the resulting phase startingwith an austenitizing temperature of a)1000° C and b) 1050° C (Schmitd 1987) b)
Variation of temperatures across a section of H-13 during quenching and theresulting phases (Biermann 1984)
During quenching, the center of the dies experiences the slowest cooling rates. Theserates are slow enough to produce non-martensitic structure at the center of theworkpiece. Figure 4.14a-b clearly shows the differences in cooling rates experienced bythe surface and the center of a bar. At these slow cooling rates, it is not possible toobtain a completely martensitic structure. One can increase the cooling rates at thecore to higher values using other techniques. But this is not practical without creatingextreme thermal gradients across the cross-section of dies. Large thermal gradients
52
result in thermal and transformation-induced distortions which are unacceptable toprecision forging applications.
a) Cooling curves form Uddeholm tests b) Cooling curves from Case Western University
c)
Figure 4-14 a-b) Cooling rates used illustrating the corresponding structures in theCCT diagram; b) Table with the conditions and the resulting properties for the
corresponding cooling rates (Wallace 1989).
The extremely high hardness of martensitic structures is due to its high resistance tothe slip and dislocation motion. This resistance is primarily due to the trapped carbonatoms in the martensitic matrix. During an actual quenching process, because of theslower cooling rates found in certain sections, several transformations can occur beforethe material reach the room temperature. As the cooling rate experienced by a section ofdie falls, the first non-martensitic phase to form is bainite. Bainite forms in quenchingspeeds intermediate to pearlite and martensite. Its structure, subsequently, hascharacteristics similar to both ferrite and martensite. Bainite that forms at lower coolingrates close to pearlite field in a continuous cooling transformation (CCT) diagram iscalled upper bainite. Bainite in the region close to martensite line is lower bainite.
53
Upper bainite contains elongated carbides that are also bigger than in lower bainite. Ashot forging dies needs ductility and toughness the lower bainite structure with finecarbides distribution and fine structure will perform better. These transformations aretime-temperature dependent and are well understood using a CCT diagram. Typicaleffect of cooling rate on the microstructure and hence, the mechanical properties of hotforging die is illustrated in Figure 14.a-b and the attached table c). From thisillustration, it is also clear that with slower cooling rates, grain boundary carbideprecipitation increases. This results in a lower Charpy V-notch value as well as asmaller yield strength. These tests were performed on small samples under laboratoryconditions. However, the underlying mechanism of transformation and microstructure-mechanical property correlation are universally true. Similar trend has beendemonstrated by Okuno (Okuno 1987) who examined the effect of quenching rates ontoughness of H-13. Again, we see that with increased time for quenching, the toughnessdrops because of grain boundary carbides and increased pearlite and lower bainitecontent, figure 4.15.
Figure 4-15. Effect of cooling rates on the phase content and the resulting toughness.Steels was austenitized at 1020 C for 30 mins and oil quenched at different rates.
(Okuno 1987)
A typical hot forging tool steel (H13) transforms into martensite completely if it isquenched to 300°C in 1000 seconds. When the die block is too big, this is notachievable uniformly across all sections. In these instances, interrupted cooling is analternative. There are two interrupted cooling techniques common to heat treaters:martempering and austempering.In martempering, the dies are quenched to a temperature just above Ms and kept thereuntil the temperature becomes homogeneous. It is then quenched crossing the Ms
curve. In austempering, the final quenching step forms upper bainite. It is important
54
that the cooling rates used to reach the first stage homogenizing temperature is highenough to miss the grain boundary carbide precipitation line. Austempering reduceddies distortion because the temperature is equalized before transformation andsubsequent cooling is slow. The resulting bainitic structure does not need temperingoperation (Krauss 1995).
Several quenching medium are available for use in tool steels. Tool steels are quenchedusing oil, air, non oxidative gas or molten salt. Each option we have, will result indifferences in cooling process that will affect the resulting microstructure, surfaceproperties and distortion. Quenching techniques available to heat treaters are
• Vacuum Furnaces with vacuum or gas quench medium• Sealed quench furnaces with gas or oil quench medium• Isothermal Salt quench• Fluidized bed quench
In vacuum furnaces, where the austenitization is done by heating the die steel byradiation, quenching is done either by cooling in vacuum or using nitrogen. Lowercooling rate may result in carbide precipitation. In the multiple-chamber vacuumfurnaces a quenching chamber permits higher cooling rates. Higher cooling rates areachieved by using nitrogen at higher pressures. Figure 4.16 shows the effect of nitrogenpressure on cooling rates achievable in a gas quench process. Also, the figure illustratesthe effect of cooling rates on the microstructure and phases present.If oil is used as medium in sealed quench furnaces, higher cooling rates are achieved.However, because there are no controls in place to affect the temperatures, there is arisk of distortion. Oil should be used as a quench medium only for small, simple dies.To reduce distortion, isothermal quench chambers that use salt baths should beemployed. These equalize temperatures around 1000° F. Again, 2-step quenching givessmall distortion than 1-step quenching. Fluidized beds also provide good control overtemperatures and distortion. Fluidized beds and oil quenching provide cooling ratesthat are most desirable from a toughness point of view. In these cases, the grainboundary carbide precipitation is kept to the minimum. Figure 4.17a-b illustrates someof the properties we can expect. This should be used only as a reference to comparedifferent quenching processes relative to one another.
55
Figure 4-16. Effect of different cooling rates resulting from the quenching pressuresfor a 11”x20”x30” block (Roche, Beaton et al. 1997)
56
a) b)
Figure 4-17. a)Effect of cooling rate on the longitudinal toughness; b) effect of variousquenching mediums on the ductility of samples taken in the three directions (1,3,4produce coarse bainite and high GBC precipitation) (Roberts and Norstrom 1987)
4.3.3 TEMPERING
Quenched structure usually is a combination of untempered martensite, retainedaustenite and carbides retained from austenitization. The goal of tempering is toimprove the strength and toughness of the quenched die steel by stress relieving andatomic rearrangement. It does this in the following ways
• Transformation of all retained austenite into martensite and bainite.• Diffusion of carbon atoms to create stronger and tougher tempered martensite• Precipitation and dispersal of alloy carbides in the matrix so that they do not
coarsen during hot working. This phenomena causes hardening and is termedsecondary hardening.
The most important parameters that affect these changes during tempering are1. Number of tempers2. Duration of temper3. Tempering temperatures
Heat treaters use at least 2 “draws” to convert retained austenite to martensite andother stable and tough phases. The first tempering cycle transforms the retainedaustenite in cementite and ferrite or martensite. The second tempering will temper theuntempered martensite produced in first temper. The second tempering cycle alsospheroidizes interlath carbides formed by the transformation of the retained austenite.The 3rd and subsequent cycles are directed at obtaining secondary hardness gains
57
through precipitation of metallic carbides. Low alloy steels like L6 used in hammersdoes not need the second tempering. H-13 requires 2 draws or tempering cycles toobtain necessary toughness. Higher alloy tool steels like H-10A requires at least 3draws. If the die block goes through fewer draws than recommended, it could result inthe following.
• Retained austenite and untempered martensite resulting in low toughness andstrength
• Absence of metallic carbide precipitates resulting in low secondary hardening• Interlath carbides resulting in reduced toughness and higher chance of
catastrophic failure of dies
Tempering time and temperatures have similar effects on the microstructure. Usuallythe tempering curves from steel producers provide variation of hardness with temperingtemperatures for a fixed time of one or two hours. Another approach to specify effect oftempering is using charts and functions that track hardness with a combination oftime and temperature. Figure 3.7 and 3.12a shows the master tempering curves for H-19 and H-13 respectively. Using these curves, we can predict what the resultinghardness will be if we temper a H-13 block for a specific duration at a specifiedtemperature. Although this approach is better suited to evaluate softening of hotworking dies during the service, not many curves are present in the literature.
Since these processes are diffusion dependent, each draw should be at least 1 hour atthe tempering temperature. This ensures that all retained austenite converts tomartensite. Typically heat treaters use at least 2 draws of at least 2 hours each to evenout the effects of inhomogeneous heat chemistry, hardening temperatures andquenching conditions. Again, since these transformations rely on carbon diffusing outof the matrix, these transformations increase in speed at higher temperatures. Ingeneral, the tempering temperature for hot forging applications range from 500-600° C.Table 4.2 shows typical hardening and tempering temperatures for tool steels. Becauseof precipitation of metallic carbides, there is a accentuated reduction in toughness inthe region of secondary hardness peak. Toughness reduction is represented by both theCharpy and KIC tests in figure 4.18. This critical temperature is characteristic of thealloy composition and the level of carbide dissolution during the austenitizing phase.For hot forging dies, it is important to temper at a temperature that exceeds this criticalpoint. Tempering at temperatures just above the secondary peak gives the maximumhardness, desirable for good performance in thermal fatigue and wear. However, if thedies experience extreme mechanical loads, an increase in tempering temperatureincreases the toughness.
Figure 4-18. Drop in toughness of H-13 with secondary hardening. a) representstoughness in Charpy V-notch energy, b) represents toughness in Kic (Pickering 1987)
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Alloying elements have a strong effect on the dies resistance to thermal softening. Toolsteels for hot work applications depends on strong carbide formers, like V, Mo, T, Cr,etc. to provide resistance to thermal softening. These effects are illustrated in Figure4.20. For example, .13% Vanadium has the ability to flatten the slope of a temperingcurve by 50%. This means that after 2 hours of tempering (after quenching or duringhot working), a tool steel with .13% V will result in a drop of hardness of half the valuecompared to a pure iron-carbon structure. However, in order for this to happen, theseelements have to be dissolved in the austenite during austenitization. On the otherhand, the dissolution of primary carbides decreases wear resistance. There is hence, atrade-off between resistance to thermal softening and wear that needs to be understoodby the forging designers. When specifying the heat treat specification, the forger shouldalso ensure that the specified mechanical properties can be achieved without unduedistortion. Figure 4.20 shows typical distortion associated with different alloys due tonon-uniformity in transformation and internal stresses due to heat treatment.
Figure 4-19 Softening retardation effects in relation for several elements in relation toa Fe-C carbon steel at tempering temperature of 540°C (Krauss 1995)
Figure 4-20. Effects of tempering temperature in tool steels dimension includingcontraction and expansion (Pickering 1987).
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a)
60
b)
Table 4-2. a, b Hardening and tempering temperatures and procedures for tool steels(Roberts and Robert 1980).
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To know the needs of hardenability, hardness, carbide content and toughness it isnecessary to have knowledge of what are the critical failure mechanism or the criticaldie loading for each application.
4.4. SPECIFICATIONS: DIE STEEL, HARDNESS,TOUGHNESS AND MICROSTRUCTURE
In die casting, dies are very expensive and dies are subjected to much more severeprocessing conditions than forging. This has prompted several companies andorganizations to come up with criteria and guidelines for accepting die steels. Somecriterion that are widely used are DCRF Data Digest 01-83-02D, Chrysler NP-2080,Peugeot-Citroen specification E.01.10.455.G and German tool steel specification(VDGDatasheet M82). There are other criteria used by Renault, General Motors, Ford andRockwell that are similar. Since adopting acceptance criterion, several operations havereported reduced die failure rates. The DCRF criterion for acceptance of H-13, forinstance, specifies the following.
• Composition should meet ASTM A-681-76 criterion. This ensures low levels of sulfurand phosphorous.
• Annealed hardness should be less than 241 BHN. This ensures the annealing iscomplete
• Annealed microstructure should consist of ferritic matrix with spherodized carbides• Untempered hardness of air-cooled 1” slab heated to 1850 F for ½ hour should be
atleast 50 HRc. This ensures the steel is hardenable.• Untempered grain size at the surface should be finer than No. 6 after hardening.
This indirectly specifies impact and fatigue strength of the steel.• Nonmetallic inclusions should be within limits of commercial quality electric furnace
melted steel. This ensures the dies have good ductility on heat treatment.
This criterion however, does not stipulate the steel making process. Other criteria usedaddress segregation and structural uniformity in different ways. Other specificationsmay specify different criterion for heat treaters with die casting issues and modes offailure in the forefront. However, in forging industry, not enough care is exercised inspecifying die steels and heat treatment.Before specification of die material, hardness, toughness and microstructure, forgersneed to understand the modes of failure in their process. Knowledge of modes of failureand the critical forging process parameters that affect the failure rate is essential forproper specification of steels and heat treating.
As we discussed earlier, the high hardness of the martensitic structures is beneficial todies resistance to plastic deformation, fatigue and wear. Dies resistance to plasticdeformation is well represented by its yield strength and its variation with typicaloperating temperatures. Its resistance to wear can be correlated to variation of hardnesswith temperatures and times. This information is available in the form of Larson-Millergeneralized tempering curves. Wear constants also exist in literature that characterizesmaterials resistance to wear.
At some point, increasing hardness starts to decrease other essential properties liketoughness and ductility. Toughness, ductility and ductile-brittle temperature (DBTT) areinterrelated properties. Low values of these properties reduce a dies ability to supportshock. Also, low values of ductility reduces dies ability to plastically deform anddissipate the energy internally without breaking catastrophically. Toughness,represented by Charpy V-Notch or KIC fracture toughness, measures the ability of a die
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to resist crack growth. Ductility is characterized by charpy unnotched tests and tensiletests performed in laboratory conditions.
Once the process has been characterized either by experimental data or simulation,using knowledge of materials and expected batch size they need to specify the following.• Die material grade and purity• Austenitization temperatures and time• Number of tempers, tempering time and temperature
In precision forging, for instance, if the mode of failure is abrasive wear, thespecification of die steel and heat treatment should focus on improving thermalsoftening resistance and hardness. It is important to ensure that the dies are designedwith appropriate shrink fits and the forging process ensures there is no unnecessarythermal or mechanical loads. Process optimization and optimal die design complementdie steel selection and heat treat specification. It is necessary to specify material andheat treatment after optimizing design and process.
This idea has been incorporated in SAMS, a computer program being developed at theOhio State University. The program, once completed, would be capable of using thermo-mechanical history and stress-strain information from FEM program, user enteredmaterial specification, properties and charts from a database and built-in failure modelsto estimate failure rate for different precision forging applications. This will serve as anaid to precision forgers in selecting steels and specify heat treatment after a process isdesigned and optimized.
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5. SURFACE TREATMENTS
Surface-engineering of dies are techniques and processes used to induce, modify andenhance the surfaces properties giving it more wear, corrosion and fatigue resistance.These techniques do not modify the soft and tough interior of dies. Figure 5.1 shows thetypical surface depths of various surface treatments. Die coatings and surfacetreatments, used in forging industry, primarily increase the abrasive wear resistance ofdies by increasing the hardness of the surface layers of the die. Figure 5.2a shows someresults from forging experiments that clearly illustrate the efficacy of surfacetreatments. These results were obtained from forging trials performed eccentric crankpress. This section lists different surface treatments applicable to precision forgingapplications and issues one need to be aware of that may affect the die performance.Most surface treatments used in dies and tools are diffusion–based. These processesrely on diffusion of chemicals into the surface, modifying the surface chemistry and themechanical properties of the surface layer. The thickness of the surface treated layer inthese types of diffusion processes rely on the time and temperature at which thehardening is performed. The time-temperature dependence is of the form shown inequation 5.1.
TKD = Equation 5-1
where D is the depth of the hardened case K is a temperature dependent constant
T is the time of exposure
Figure 5-1. Thickness of various coatings and surface treatments (Subramanian1996)
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Figure 5-2.a. Comparison of wear amounts of surface treated upsetting tools after1000 forging cycles with lubricant (Deltaforge-31) (Doege, Seidel et al. 1996)
Different diffusion based surface hardening techniques which may be applicable toforging dies are
• Carburizing and pack cementation• Nitriding• Carbonitriding and Nitrocarburizing• Boriding• Toyota Diffusion Process• Oxide coatings• Thermoreactive diffusion• Weld Overlays
5.1. CARBURIZING
Carburizing is the process of adding carbon to low carbon steels. Not typically used forforging dies, the process relies on heating the parts to high austenetizing temperaturesof over 1500 °F and exposing the surface to a carbon rich atmosphere. Carbon diffusesinto the austenitic surface of the parts, which are then quenched to provide amartensitic structure on the surface. As discussed before, martensite has excellentwear resistance. Coupled with the soft and tough core, this surface treatment gives theparts good resistance to mechanical shock as well as wear.Case depths and hardness levels achievable are dependent on the time of exposure andthe richness of carbon at the surface. Prolonged exposure to carbon-rich atmosphereresults in a deep case. However, the surface may have excessive retained austenite andfree carbides, which in turn will result in excessive residual stresses.Based on the medium used, carburizing can be any of the following.• Gas carburizing• Vacuum carburizing• Plasma carburizing• Salt bath carburizing
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• Pack carburizingOf these gas carburizing is the most commonly used process because of ease of processcontrol and low equipment costs. Pack carburizing, the uses a solid carburizing pack, isalso widely used. Carburizing has limited application if precision forging dies because,dies materials used are medium to high alloy steels. The process has advantages offlexibility and low cost for low production. However, labor cost for cleaning andenvironmental restrictions make the gas or liquid process cheaper. The gas and plasmacarburizing process are also more controllable.
5.2. NITRIDING
Nitriding, similar to carburizing, is a process of hardening the surface by diffusingnitrogen into the surface. Nitriding processes are performed at temperatures between925 and 1050 °F (500 to 550 °C) (ASM 1964) where the structure is still ferritic. Theprocess results in formation of and outer case of Fe3N and a inner layer that isstrengthened by a solid solution of N. In some cases, a white layer of Fe4N is formed.This layer, also called the “white layer”, may easily spall during use and has to beavoided.Steels nitrided are typically medium carbon steels with strong nitride-forming elementslike aluminum, chromium, vanadium and molybdenum. It is important that temperingof the die steel be performed at a temperature exceeding the nitriding temperature priorto nitriding in order to optimize the property combination of the core and the surface ofthe dies. Also, because of the low nitriding temperatures, there is generally littledistortion from this heat treating process.Although the depth and hardness of the nitride case depends a great deal on thenitriding time, these properties (particularly the hardness) are sharply dependent on thecomposition of the steel as well. Die steels containing large amounts of strong nitrideformers such as chromium, vanadium, and molybdenum form shallow, very hardsurface layers. On the other hand, low-alloy chromium-containing die steels (such as6G, 6F2) form deeper surface layers which are tougher, but not as hard.There are many techniques for nitriding: gas-nitriding, liquid-bath nitriding, ion-nitriding, etc. Each of these will be discussed separately.Gas Nitriding. Gas nitriding is a surface hardening process in which nitrogen isintroduced into the surface layers of ferrous materials by holding them in contact with anitrogen-containing gas, which is usually, ammonia. Because a brittle, nitrogen-richlayer (the "white nitride layer") is produced by this process, the depth of the nitridedcase is usually kept small. Sometimes, a special two-stage gas-nitriding process, whichminimizes the depth of this layer, is employed (Weist 1986).Liquid nitriding. Nitriding in a liquid salt bath, or liquid nitriding, is performed at thesame temperatures as gas nitriding, approximately 925 to 1050 °F (496 to 566 °C), buttypically requires less time than conventional single-stage gas nitriding. The salt bathsconsist primarily of mixtures (in varying proportions) of sodium and potassium cyanide(from which the nitrogen is released during nitriding) and sodium carbonate, potassiumcarbonate, and potassium chloride. These baths result in cases containing bothnitrogen and carbon compounds. Modifications of this heat treating procedure includea process involving aeration. (ASM 1964) This leads to a less brittle case of Fe3Ncompared to gas nitriding process which develops cases containing very brittle ironcompounds richer in nitrogen (e.g., Fe2N). Commercially, liquid bath nitridingprocesses such as Tufftriding and the Berry-Wear process have been used onmetalworking tooling. Employing special salts containing sulfur compounds, the Berry-Wear process appears to have the advantage of producing an outer skin which serves asa dry lubricant as well as a very hard wear-resistant surface (Brochures ). This latterprocess has been very successful in cold forming applications, but its usefulnessappears not to have been documented in hot forging yet.
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Ion-nitriding: Ion-nitriding, also known as plasma nitriding and glow-dischargenitriding, is yet another form and probably the most recently developed method ofalloying the surface layers of ferrous parts with nitrogen. (Taylor 1981), (Edenhofer1976), (Brochures ) In this process, a part to be nitrided is placed in a reaction vesselinto which an atmosphere containing nitrogen and hydrogen are introduced. The partis electric resistance heated to 930 °F. It is made the cathode and the reaction vesselthe anode in an electric circuit. A glow discharge between the vessel and the partcauses ionization of the gases, causing nitrogen ions to impinge upon the surface of thepart. Because these ions have much greater energy than those in gas or liquidnitriding, penetration and thus surface nitriding is much quicker in ion-nitriding.Among other advantages of this form of nitriding is the ability to control and minimizethe extent of brittle "white-layer" formation. In fact, with proper atmosphere control,nitriding surfaces of Fe4N can be formed. This compound is very ductile, and thusparts with hard, wear-resistant, and tough surfaces may be produced. The majordrawback of this method is the need for a special reaction vessel whose size limits thesize of parts that can be treated.
Nitriding can be used in conjunction with alloy selection to selectively enhanceresistance to certain modes of failure. In table 5.1, Krishnadev (Krishnadev 1997)presents efficiency of various surface treatments in increasing resistance to mechanicalfailure, thermal softening, heat checking and wear.
Krishnadev (Krishnadev 1997) also presents toughness, hot hardness and softeningcharacteristics of nitrided and non-nitrided of H-13 and alloys of similar composition.Figures also show the different levels of hardness and toughness achievable withdifferent coatings and nitriding.
Table 5-1. Response of different tool steels to several surface engineering towardsenhancement of toughness, hot hardness, heat checking, temper resistance
(Krishnadev 1997)
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Efficacy of nitrided tools were tested by Dean and others (Dean 1987) using extrusiontype testing. Relative wear rates of nitrided and non-nitrided tool steels in extrusionthey obtained are shown in figure 5.4. Dean indicates that nitriding reduces wear asmuch as 50%.
a)
b)
c) d)
e)
f)
Figure 5-3. Results for hot work tool steels in the H13 group presented by Krishnadev(Krishnadev 1997) (a) Composition (b) toughness (c) hot hardness (d) softening of thealloys 2-3 with and without nitriding e) Hardness achievable with different coatings
and the alloys chemical composition f) Charpy impact toughness of H-13 and treatedalloy No. 3
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Figure 5-4. Relative wear rates of nitrided and non-nitrided tool steels used inextrusion forging (Dean 1987)
5.3. CARBONITRIDING AND NITROCARBURIZING
Carbonitriding and nitro-carburizing are diffusion – based surface treatment techniquesthat combine the effects of nitriding and carburizing. Carbonitriding relies on hardeningcarburized austenitic surface layers using nitrogen as an agent. It is performed attemperatures where austenite is stable. Nitrogen and carbon diffuses into the surface.Carbon, on quenching, provides the superior hardness levels. Nitrogen increases thehardenability of the surface, thus reducing the need for drastic cooling and subsequenthigh distortion. It should be noted that nitrogen, like carbon is an austenite stabilizer.Excessively rich carbonitriding medium and prolonged times may result in highnitrogen concentrations at the surface that may result in high levels of retainedaustenite. Retained austenite is detrimental to hardness and wear resistance.
Austenitic nitrocarburizing relies on formation of Fe3N (ε carbonitrides) at the surface toimprove the hardness levels. It is carried out at temperatures of 675 °C to 775 °C.Because the mechanism relies on the formation of carbonitrides and not carbon ornitrogen trapped in surface matrix, there is no need for quenching. This process, hence,results in low distortions.
Ferritic nitrocarburizing displays superior fatigue resistance and relies on bothformation of ε carbonitrides and diffusion of carbon and nitrogen into the substrate.The process is carried out at temperatures where the surface is still ferritic – around570 °C. Diffused nitrogen is trapped in a solid solution by quenching in oil. The whitelayer of ε carbonitrides improves the wear resistance while the diffused layer results inimprovements in fatigue strengths upto 120% (ASM 1964). The yield strength of thesubstrate surface also increases. For this treatment to be effective, the surface layersneed to be clean of oxides, scales, oil or other contaminants. Vapor degreasing or grit
69
blasting with fine abrasives may be necessary steps to achieve the most out of ferriticnitrocarburizing process. Various processes exist like Nitemper process, Alnat-Nprocess and black nitrocarburizing that rely on controlling the composition of whitelayer formed. For details on this processes, refer to ASM handbook (ASM 1964).
Venkatesan and others (Venkatesan, Summerville et al. 1998) have evaluated nitro-carburized dies and compared its performance with quench and tempered H-13 andASM 6F3, Nitrided H-13, Borided and Vanadised (TD) H-13. Results obtained from theirtests are provided in Table 5.2. They found that nitrocarburized dies performed verywell compared to untreated H-13. Details of the process they used is unknown.
H13 Dies 6F3 DiesTreatments Top Bottom Top BottomQ&T 46 110 156 236Nitro-carburized 4 5 5 37Nitrided 10 12 11 9Borided 5 6 0 0Vanadised 0 0 0 0
Table 5-2. Average maximum wear depths (µm) on surface engineered dies afterupsetting 500 AISI 1040 steel billets at 1070° C (Venkatesan, Summerville et al. 1998)
5.4. BORIDING
Boriding or boronizing process relies on diffusion and subsequent absorption of boronatoms in the metallic lattice and formation of interstitial boron compounds to hardenthe structure. Diffusion treatment can be carried out in either a gas, molten salt, orpack media at a temperature between 700 °C to 1000 °C, depending upon the processand the material to be borided. Extremely hard-surface layers ranging from 11450 to5000 HV that has a low coefficient of friction are formed if the base metal forms borides.The process does not require quenching. If the base material has to be heat treated, theheat treatment can be done after boriding, although care is required to reducequenching stresses to prevent spalling of the borided layer. Borided layers resistthermal softening better than nitrided layers. They also exhibit moderate to highresistance to oxidation. However, boriding provides marginal increase in fatigueendurance limits.
Boriding of steels is also done electrolytically. Boron atoms are electro-deposited ontothe metal from a bath of molten salt containing fluorides of lithium, sodium, potassiumand boron. The dies are borided in the 1470 F (800 C) to 1650 F (900 C) temperaturerange in an atmosphere of argon or a mixture of nitrogen and hydrogen. Thickness ofcoating is from 0.0005 to 0.002 in (0.013 to 0.05 mm), and treatment lasts 15 minutesto 5 hours (Fiedler 1972).
It has been stated that boriding results in undesirable interaction with alloyingelements of hot-work die steels (H series) and develops a soft layer (Burgreev 1972).Porosity in the borided layer can develop for steels, which require post-boriding heattreatment. For this reason, it is preferred to limit boriding to those alloys that do notrequire further high-temperature treatment. For example, A6 (075 C, 20.0 Mn, 0.3 Si,1.0 Cr, 1.35 Mo) air-hardening steel can be hardened from the boriding temperature bycooling in air, and only requires tempering. This steel, therefore, can be safely borided.
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Boriding process is also known to improve the wear resistance by forming borides withthe subsurface die steel. Shivpuri and Semiatin (Shivpuri and Semiatin 1988) reportwork by Vincze who borided dies at 900° C for 3 hours followed by quench and temper.Vincze reported an increase in wear resistance of borided dies by 70% compared tountreated dies. Burgreev and Dobnar (Burgreev 1972) also report large increase inhammer die forging die life when boriding is used.As reported earlier, Venkatesan and others (Venkatesan, Summerville et al. 1998) alsoreport enhanced wear resistance of borided tools compared to untreated dies. Boridinglow alloy steels result in a jagged boride layer that are deeper than boride layers formedin high alloy steels. This is because, alloying elements reduce the diffusion of boron intothe substrate. This may explain results obtained by Venkatesan who found thatperformance levels of borided H-13 are comparable to nitrocarburizing. However, theyfound that the efficacy of boriding was higher with low alloy die steels like 6F3.
5.5. THERMO-REACTIVE DIFFUSION (TRD)
Thermo-reactive processes also called Toyota diffusion process (TD process) is anotherdiffusion type process that relies on forming hard carbides of V, Cr and Nb on thesurface of dies. The process is performed by placing preheated dies in a molten boraxbath at temperatures from 850 °C to 1050 °C for times ranging from ½ to 10 hours. Thebath also contains strong carbide forming elements like Niobium, Vanadium, Titanium,Chromium. Unlike other processes discussed earlier, TRD process results in buildup ofsurface of carbides. Also, Vanadium and chromium diffuse into the steel substrateforming iron-chromium and iron-vanadium solid solution layers beneath the carbidelayer. After treating, the die is quenched in air, salt or oil and tempered. Drasticquenching may cause unacceptable distortion and needs to be avoided. To reducedistortion, it is preferable to pre-machine and grind the dies before TRD processing.Best results are obtained for steels with atleast .3% C.
Dies processed by this method have excellent wear resistance and resistance tocorrosion and oxidation. Arai (Arai and Iwama 1981; Arai 1992; Arai and Komatsu1993; Arai 1995; Arai, Fujita et al. 1995; Tsuchiya, Kawaura et al. 1997) has done a lotof work in validating the use of TRD process in cold and hot forging. By coating 3”flashless dies made of SKD 62 with Cr Carbide using the TD process, he was able todouble the life of dies. The life of dies increased from 5000 to 10,600. Arai also reportsthat TD treated steels did not peel or crack under repeated blows with a pointedhammer. Under similar conditions, TiC coated by PVD or CVD cracked after 50,000pieces and peeled after 1000,000 pieces.
Venkatesan and others (Venkatesan, Summerville et al. 1998) found that vanadiseddies showed the least wear of the tested specimen compared to niitrocarburizing andboriding. They also noted that vanadisation dies showed no traces of wear irrespectiveof the type of substrate used. They also noted that nitro-carburizing and nitridingresulted in similar wear rates for both types of die steels used in the study.
5.6. OXIDE COATINGS
Oxide coating improves the performance of hot-work tool steels. The oxide scale formedduring the heat treatment is abrasion resistant and helps to hold die lubricants. In casethis protective layer in case it is removed by final die finishing it should be re-createdwhen the die does not have other surface heat treatment. The oxide layer can berecreated by die exposure to steam (~564°C) or by heating in liquid sodiumhydroxide/sodium nitrate salts for 5 to 20 min at 140 °C. The resulting layer is about
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5µm thick and is very adherent (Roberts, Krauss et al. 1998). Quinn and co-worksshowed, in a series of studies in fundamental wear and oxidation mechanism, thatunder hot conditions a layer of oxide prevents the occurrence of adhesive wear (Quinn,Sullivan et al. 1980; Sullivan 1981; Quinn 1983).
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6. ADVANCED DIE MATERIALS AND SURFACEENGINEERING TECHNIQUES
6.1. CERAMICS: SIALON, SILICON NITRIDE ANDSILICON CARBIDE
Ceramics are chemically inert compounds that can retain its properties at hightemperatures. Sialon and Silicon Carbide and Silicon nitride are some potential ceramicelements that can be used in tooling for precision forging.
a) b)
c)
Figure 6-1.Compilation of several properties versus test temperature for ceramicsfrom Ohuchi (Ohuchi 1990). a) hardness b) thermal expansion c) Yield stress.
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Hot pressed Silicon Nitride is a ceramic that has extremely high hardness, hightoughness and wear resistance. Currently it is used in applications like nozzle rings,bearings, rotors and cam followers in internal combustion engines. It also has goodthermal shock resistance and good hot hardness and maintains its temperature andoxidation resistance at 1200° C. Applications of Silicon Nitride as a coating, is however,limited because of its poor adhesion with the substrate.
Silicon Aluminum Oxynitride (Sialon) is a new group of solid solution compositions thatalso possesses excellent thermal shock resistance. Sialons have similar properties asSilicon Nitride but Sialons have a superior resistance to oxidation at high temperatures.Silicon Carbide with extremely high hardness is normally used in grinding wheels aswell as various internal combustion engine parts like valve seats and flame cans. Datacomparing
There has been some interest shown on this type of material primarily for automotiveapplications. Miyoshi plasma-deposited amorphous silicon nitride films at both low andhigh frequency applications. The author found that high frequency deposited layer hasbetter resistance to the shear (better adherence to the substrate). The coating alsoresulted in low adhesion (soldering) and low friction in pin-disc type tests until 700°C.
There are several ceramics available with comparable harnesses, thermal expansionand yield strength of these 3 compounds are shown in figure 6-1 a-c. Several ceramicsand carbides exist that have application in precision forging either as inserts or ascoatings. It should be realized that these compounds typically lack tensile strength andneeds to be constrained in a shrink ring. Also, they may not be applicable as coatingsbecause of their lack of adhesion to the substrate and their dissimilar thermalexpansion coefficients that may lead to cracking.
6.2. ALUMINIDES: NICKEL AND TITANIUM
Nickel aluminides are relatively new intermetallic die materials that exhibit better hightemperature properties compared to conventional hot work steels and nickel basedsuperalloys like IN 718. Typical composition of some nickel aluminides is provided inTable 6.1. Although this is a relatively new compound, it is the same compound (γ’compound) that gives superalloys like 718, its strength on aging.
The application of Nickel Aluminide in hot and warm forging is very new. Although softat room temperatures, nickel aluminides retain their yield strengths at highertemperatures. Figure 6.2 shows typical yield strengths and tensile strengths of NickelAluminides. Figure 6.3 shows the increasing yield strength of these intermetalliccompounds with temperature up to 500 °C beyond which the yield strength drops.
Table 6-1. Compositions of various grades of Nickel Aluminides (Blau 1992)
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a) b)
Figure 6-2. Mechanical properties a) compressive Yield strength for Ni alloy 718 andNickel aluminide 221M-T (Al 7.6-8.2; Cr 7.5-8.2; Mo 1.3-1.55; Z 1.4-2.0; B 0.003-0.01
Ni balance) b) Tensile and yield strength for 221M-T alloy (Maddox and Orth 1997)
The relatively high hot hardness gives these intermetallics, very high wear coefficients.Table 6.2 shows the results of pin-on-disc type tests on Ni3Al. Although these areconducted at room temperatures, these numbers give us some idea of its wearresistance compared to conventional hot work steels. Tests that have been doneindustries show up to 10X life increase for preforming dies. The high yield strength alsogives nickel aluminide relatively late crack initiation as indicated in Figure 6.4. Figure6.5 compares the variation of yield strength of this class of materials with hastealloyand stainless steel. Table 6.3-6.5 show some physical properties available.
1
Figure 6-3 Yield strength of various grades of nickel aluminides (Blau 1992)
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8
Table 6-2. Wear constants obtained through pin-on-disc type tests for various gradesof Nickel Aluminides (Blau 1992)
Figure 6-4. Comparison of crack growth data for Nickel Aluminide compared to otherhigh temperature alloys. (Fuchs, Kuruvilla et al. )
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Figure 6-5 Comparison of yield strength of IC-15 to those of other high temperaturealloys. (Horton, Liu et al. )
Table 6-3. Some physical properties of IC-50 (Oak )
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Table 6-4 Variation of yield strength, ultimate strength and ductility of IC50 withtemperature (Oak )
Table 6-5. Variation of modulus of elasticity of IC50 with temperature (Oak )
The main deterrent in the use of Nickel Aluminide is that it is not very easilymachinable. Machining nickel aluminides require use of positive rake tooling and lowspeed and machining feeds. For this reason, the material is received in as-castcondition. However, the material machines well using electro-discharge machine.
6.3. WELD OVERLAYS
The hardfacing is a coating process that applies a surface deposit that metallurgicallybonds to the base material. In the past, the process was used primarily for repair andmaintenance of dies and molds. Now, it is increasingly being used as an inexpensivemeans for depositing a hard layer on localized wear-prone die areas. Nugent (Nugent1986) reports in his study that weld deposits of Alloy 625 increased forging die life by400%. Kohappa (Kohopaa, Hakonen et al. 1989) reports comprehensive test results(Figure 6.6) that compare wear resistance of various weld consumables.
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Wear of weld consumables
100
29
83
1
75
43
108
56
1
40
020406080
100120
H13
OK 853
8
OK 855
8
OK 856
5
OK 832
855
50 535
OK 932
5
OK 930
660
6
Weld consumable type
Wea
r
Figure 6-6. Results of wear tests on various welding consumables (Kohopaa,Hakonen et al. 1989)
The weld deposits used could be one of the following.• Deposits of identical material onto a die block to repair it or to allow resinking• Deposits of higher alloy steels (e.g., chromium hot-work steels) onto the die surface
of low-alloy steels to improve the service performance of the dies (e.g., wear and heatresistance).
• Deposits of hard or high-temperature materials (usually cobalt or nickel-basedalloys) onto low-alloy or hot-work steels to improve the service performance of thedies.
Different hardfacing processes.Before discussing specific alloys, the processes by which they are deposited will bebriefly reviewed. The first step in any of the hardfacing processes should be theannealing of the die block into which the rough impression has been sunk. This relievesresidual stresses and helps prevent cracking during welding of the surface layer. Afterannealing, the die block should then be reheated to a temperature of 600 to 1200 °F(325 to 649 °C), which is also necessary to minimize cracking due to thermal gradientsset up between the surface and interior during welding. The application of the surfacelayer can then be performed by one of a number of welding processes (Knotek 1979):1. gas torch welding (combustible gas welding)2. manual arc welding3. submerged arc welding4. gas shield arc welding (TIG or MIG5. open arc welding6. thermal spraying7. fusion treatment8. plasma spraying (plasma-arc welding)9. transferred arc plasma10. flame plating11. Deposition process (Electroslag welding)
Together with the solidification conditions, the amount of melted base material andbase material dilution is important for wear properties. For repair of dies, the shielded-metal-arc method is preferred. It allows high productivity and has the advantage of lowheat input and thus minimal distortion of the die cavity.
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After welding, the die block must be cooled to room temperature to prevent cracking ofthe welded deposit. The die impression is then finished, machined and ground. Heattreatment (austenitizing, quenching, and tempering) of the die block is performed last.Once again, differences in thermal properties between the base metal and surfacedeposit are critical insofar as thermal cracking is concerned.
Hardfacing Alloys: For hardfacing, welding alloys are generally base on iron, cobaltand nickel metals. Hard phases are formed by addition of carbon (in Fe) or boron (in Ni).The preferred application methods for various alloys are: iron alloys deposited bysurface weld methods, cobalt alloys by welding and powder surfacing and hard nickelalloys ing the form of powder. The volume fraction for hard phase is very important forthe wear resistance in the weld deposit. Often there is no proportional dependence andthe best wear resistance is not achieved by the highest hard phase concentration.
Various ferrous alloys are used to repair steel dies or to lay down deposits of better wearand heat resistance in the welded deposit. Often there is no proportional dependenceand the best wear resistance in not achieved by the highest hard phase concentration.Different microstructural combinations are used to increase wear resistance of toolsteel, these include transformation behavior (bainite, eutectic) and the use of carbideforming elements where chromium is used as alloying element. Austenitic andaustenitic-ferritic material are preferred for wear resistance under higher loads.
Hard-faced tool steels have to be heat treated before use. However, hyper-eutectic castor carbide sinter alloys are not suitable for heat treatment and the weldment fromcarbide filler rods exhibit the required material properties directly after welding. Inrespect of the economic importance, hard facing with iron base alloys predominates incomparison with nickel and chromium alloys (Farmer 1979; Knotek 1979). This is morerelevant with the increasing automation of hard surfacing and the use of robots inwelding systems.
Nickel- and cobalt-base alloys are the usual choices for hard-facing of dies. Questionsconcerning the transformation or primary phase instability during hard surfacingprocess can be considered of secondary importance in hard cobalt or nickel alloys. Thematerial properties are present after solidification from melt. Use of these alloys inhardfacing offers a considerable saving over die blocks of these alloys. In a typicalhardfacing operation, one or two layer of alloy, each about 0.010 to 0.050 in. (0.25 to1.27 mm) thick are deposited in the die. If a very large amount of buildup is desired orrequire, however, it is advisable to apply layers of stainless steel, high nickel alloy, orlow-alloy filler metal first rather than many layers of hardfacing material (Haynes )(Acros )
Hard nickel alloys are processed generally as metal powders (P/M) and to a lesserextent as cast rod and electrode. Some nickel base alloys are applied as layers. Withseveral alloys, the hardening during loading is used to increase the wear strength, e.g.for cladding cutting tools and die blocks for hot working.
Evaluation of hardfacing alloys: The properties of the cobalt alloys with chromium,tungsten and carbon have been investigated by researchers at the University of Aachen,West Germany (Knotek 1978; Knotek 1979). The influence of the cobalt matrixcomposition and the carbide content on the impact strength, thermal shock resistance,coefficient of linear expansion, tensile strength, ductility and hardness, as a function oftemperature are used for evaluating the wear behavior of the coating. Hard cobalt alloysare processed as cast rods, electrodes, filler rods and metal powder. Another recentlaboratory investigation into the elevated temperature properties of cobalt basedhardfacing alloys (Crook 1983) concluded that for the STELLITE group of cobalt alloys,
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the higher the cobalt content, the better the resistance to metal-to-metal wear in thetemperature range 0 to 1382F (0 to 750C). For this group of alloys, wear rate decreaseswith increasing temperatures, in the range 1562 to 1832F (850 to 1000C). Figure 3.23(Semiatin and Lahoti 1981) shows the effect of temperature and heat treatment in somehard facing alloys.
For a given matrix chemistry, increased hard phase volume friction may be of somebenefit in resisting metal-to-metal wear The cobalt-chromium and cobalt-iron-chromium alloys exhibit a maximum metal-to-metal wear rate around 482F (250C.).This study also reported that the high nickel alloys have relatively poor self-mated anti-galling properties at room temperature, yet exhibit low wear rates under low load/highspeed conditions, versus case hardened 4620 steel in air at room temperature. The lowwear rates of nickel-rich hardfacing alloys have been attributed to their oxidationkinetics and the nature of their oxide scales. Low wear rates and the formation of veryshiny oxide scales, termed glazes, characterize the high temperature behavior of somenickel-chromium alloys.
Electro-spark deposition. Electro-spark deposition (ESD) a variation of hard-surfacewelding, has been used extensively in Europe for improving the galling resistance ofmaterial (Sheldon 1985). Electrodes of WC, TIC and Cr3 C2 materials have been usedfor deposit on 316 stainless steel and other substrata. ESD has been found to beeffective in fusing metallurgically bonded coatings to the substrate at low heat with thesubstrata remaining near the ambient temperature.
6.4. CRYOGENIC TREATMENTS
The ideal hardening treatment would transform 100% of the austenite to martensiteprior to tempering. However, in practical cases some percentage of the austeniteremains untransformed. Cold or cryogenic treatment can improves the percentage ofthe transformation increasing the strength, dimensional and microstructural stability,wear resistance and reducing the tendency for grinding cracks.Quenching to cryogenics temperatures following the room temperature quenching woldbe offer the maximum austenite-martensite transformation. However the class of steeland the geometry complexity usually demands to apply the following cycle: quenchingto room temperature; immediate tempering, cryogenic quenching followed for the finaltempering. The die geometry and design will determine in last instance the applicabilityof the treatment.
The cold treating can be applied using dry ice in a container that reaches ~ -75°F, orcommercial units with circulating ar that reach -125°F. The liquid nitrogen can reach ~-320°F but it dos not have much use due to the cost. The advantages of the cryogenictreatment are still in discussion, although when comparing tool steels (D2,A2,M2,O1)treated to –120°F and –310°F the cryogenic treatments showed approximately twice thewear resistance (ASM 1993).
6.5. BRUSH PLATING TECHNIQUES
Brush plating is another coating technique that is used to coat cobalt-nickel, cobalt-molybdenum, cobalt-tungsten and chromium coatings. There has been some referencein use of this technique for enhancing performance of forging dies. Dennis and others
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(Dennis and Still 1975; Dennis and Jones 1981; Dennis and Mahmoud 1987; Dennisand Sagoo 1991; Dennis, Turner et al. 1991; Dennis and Such 1993) have done a seriesof studies in characterizing electro-deposited cobalt coatings. They attempted toevaluate cobalt coated forged dies with dies coated with other techniques. They forgedcylindrical test specimens similar to ones used by Thomas (Thomas 1970),
Table 6-6. Wear volume obtained after 100 forgings using flat dies electro-depositedwith some wear resistant coatings (Still and Dennis 1977)
Rooks (Rooks 1974) and others between flat coated dies and measured wear usingsurface roughness measurements. Using this setup, they evaluated cobalt-nickel,cobalt-molybdenum, cobalt-tungsten and chromium coatings. They found that cobalt-molybdenum and cobalt-tungsten coatings provide the least wear. Table 6.2 provides acomparison of wear resistance of different coatings obtained forging cylindrical mildsteel billets using a flat No.5 die. To get a more realistic comparison, after thepreliminary cylindrical specimen testing, Stills and Dennis used a more complicatedforging with flash, to measure wear resistance of the different Cobalt coatings. Wearmeasurements were made near the flashland, where the wear was highest. Figure .12shows the variation of resulting wear area with number of forgings made. These dieswere made from No. 5 die steels or Chromium steels. Table 4.3 shows the result of useof Cobalt based coatings on industrial dies. Table 4.4 provides some results of brush
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plating coatings on simulated testing. The results of industrial trials on hot forging diesbrush plated with Co-Mo alloy coatings [Dennis, 1981 #201 is given in
Figure 6-7. Wear rate variation for dies with sharp radii and fillets, for differentcoatings. M11, M12 and M-14 are Co-Mo coatings, W2 and W3 are Co-W coatings
(Still and Dennis 1977)
Die life (N0 of forgins)Alloy Die tipeUnplated Plated
% Increase indie life
W2 A ~ 9,600 ~ 11,900 ~ 24W2 B ~ 4,000 ~ 8,000 ~ 100M 12 B ~ 4,000 ~ 8,000 ~ 100W2 C ~ 3,000 ~ 6,000 ~ 100W2 C ~ 3000 ~ 6,000 ~ 100a) results related to nitried dies
b) It was estimed that about a further 1000 forgings could have been produced fromthese dies but the production run of billets.
Table 6-7. Results of industrial trials of use of coatings. 17a represents non-roundshallow dies, 17b yoke-type dies and 17c gear blank dies (Still and Dennis 1977)
Die Type No of Results and comments Die Type No of Results and comments
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setsPlated
setsplated
Gearselector
01 120% improvement in die lifecompared with an unplated one
Turbineblade
01 2200 forged. This is normal lifebut die will be used again as stillin tolerance
Round 01 34% improvement in die lifecompared with unplated one
Controllinkage
01 Plated die ran well, but suddenlyfailed, possibly due to forging coldbar stock
Coupleflange
03 20, 26 and 21% improvement indie life compared with unplatedone
Turbineblade
01 Used to forge Nimonic alloys had'normal' life. On removal fromforge found to be in tolerance andsuitable for further use.
T Piece 01 Production ceased, but estimated itwould have increased die life by25%
Connectinglink rod
01 32% increase in life comparedwith unplated die. Improved metalflow and lower forgingtemperature noted
Link pin 01 Production ceased, but estimated itwould have increased die life by25%
Suspensioncup
01 Brush plated over welded areas ofdie - no problems encountered.Improved die life by 56% overunplated die life
Gear blank(nitrided)
01 18% improvement in die life,reduction of sticking of theworkpiece to die
Ford Transitfront axledies
02 Improved life but not sufficientlyto be viable economically
Slackadjuster
01 No improvement, but failure isnormally due to cracking and noterosion
Turbineblade
01 An increase of 100% over theprevious maximum life wasachieved
Rocker arminsert
01 13% improvement in die life,compared with unplated die
Heading 01 Comparison of equivalentpressings from plated and unplatedtooling shows the effect ofreduced die wear using a plateddie
Largeuniversaljoint
02 1 pair 77% increase in die life. 1pair average life (this pair wasTufftrided giving a poorelectrodeposit)
Extrusioninserts ½ 2½ indiameter
19 Used for extrusion of titanium forturbine blading. Glass lubricantused. Life not significantlyincreased but great improvementin surface finish so that scrap ratereduced
Largeuniversaljoint dies
04 All 4 sets produced less thannitrided dies used as a control
Turbinebladepreform
02 The plated dies produced 80-250pieces but normal life was as lowas 25 pieces
Bolsterchisel
01 32% increase in lifecompared with unplated dies
Open endedSpanner die
01 Average life
Table 6-8. Results of industrial trials on hot forging dies brush plated with Co-Moalloy coatings (Dennis and Jones 1981)
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Figure 6-8. Results of simulated hot forging tests with different coatings (Dennis andJones 1981)
Figure 6-9. Variations of wear area with number of forgings. The dies used were flatdies with dies having sharp radii and fillets (Dennis and Still 1975)
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Subsequently, Lodge and others (Lodge, Still et al. 1979) brush-plated cobalt alloyelectro-deposits on to 24 different dies making parts ranging from coupling flanges,turbine blades, gear selectors, track rod pins to steering links. The dies ranged from 4kg to 500 kg. Results showed increase in die life ranging from 20 to 100%. They reportthat the coatings did not peel off or crack under forging conditions.
6.6. VAPOR DEPOSITION: PVD AND CVD
In physical vapor deposition or PVD coating processes, the coated material istransformed to gaseous state, which condenses in vacuum on the substrate surface toform atomistic bond with the substrate surface.
Very widely used along with PVD, CVD or chemical vapor deposition relies on volatilecoating compound reacting with other gases or vapors to form atomistic layer of coatingon the hot substrate. Temperatures range from 200 °C to 2200 °C and pressures rangefrom 60 Pa to .1 Mpa.
Both these processes are very versatile in compositions one can coat. Also, the coatingscan be produced with high purity and fine microstructures. Very thin coatings can beproduced with extremely high adhesion and generally, the substrate does not need anypost finishing.
Several work exists in the literature on the use of PVD and CVD techniques. This reportwill only mention a few for sake of completeness. Mirtich evaluated thermal fatigue ofseveral coating applied by sputter deposition coating on a H-13 base using ion beamtechnology. They found that 1 micron thick tungsten, molybdenum and platinumcoatings improved the resistance to thermal fatigue until the coatings are fractured[figure 6.10).
Figure 6-10. Ratio of cracked area of coated corners to an uncoated corner forvarious materials (Mirtich, Nieh et al. 1981)
Krishnadev also reports hardness and toughness of several specimen that have beenPVD coated with TiN (Figure 5.3). The figure presents the hardness achievable withdifferent coatings. However, no wear test results are available for these coatings.
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6.7. THERMAL SPRAYING
Thermal spraying, as the name indicates, is spraying of material coated by propelling itin a molten state on to the substrate. The main advantage of this technique is that thesubstrate is kept at relatively low temperature, avoiding any distortion ormicrostructural changes. Techniques like plasma arc spraying, electric arc sprayingand detonation type spraying are part of this category.
Experimental work by Monika (Monika 1981) made using the following coatings:nitriding,, suphurizinng, diffusion chromizing, Chromium plating, plasma spraying withnon-metallic coating of Al2O3, plasma spraying with metallic coatings of Cr, WFe, WCtypes and burnishing. The wear tests used hot work die steel (900-1100°C) againstcarbon die steel (20-500°C) with pressure of 10 Mpa in dry condition. The plasmasprays (W2C + 2%Co, Cr, Wc) resulted in approximately three times less wear than theothers (Figure 6.9). The plasma spraying coatings also presented good results inthermal fatigue, although the burnished coating did not presented cracks. Increasingthe top temperature from 500°C to 700°C doubled the cracks depth. The results fromindustrial tests are presented in table 6.9. The plasma sprayed coating Cr and WCpresented good results for hot forging applications.
Table 6-9 Results of production testing of various surface treatments (Monika 1981)
Figure 6-11 Wear of different thermal sprayed coatings (Monika 1981)
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Figure 6-12 The burnished coating did not presented cracks. Samples 45mmdiameter by 40mm high, induction heated during ~ 18s and cooled by 10s between
temperatures of 20-700°C. (Monika 1981)
6.8. LASER SURFACE MODIFICATION
Laser alloying is a form of reactive coating where the laser treated alloy enters thesubstrate matrix. These techniques are very useful when it is necessary to selectivelyenhance the properties of certain regions of the dies. The technique relies on applyingpowdered alloy of desired chemistry at critical regions of interest and applying highpowered laser to melt and diffuse the compound into the substrate. Alloyingmolybdenum and vanadium carbide enhances the hardness retention and hot wearresistance.
There are few results that support the use of laser modification. Table 5.1 and figure5.3 (Krishnadev 1997) compares effect of laser welding on wear performace of H-13 andother alloys to other surface enhancements. Cser and Lang present results of die lifeimprovement using laser surface alloying, figure 6.13.
Figure 6-13. Effect of laser surface modification on wear performance of hot work diescompared with nitrided dies (Cser, Geiger et al. 1993)
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6.9. ION IMPLANTATION
A relatively new technique, ion implantation relies on directly impregnating thesubstrate material with atoms of any compound. This is done using high energy ionbeams. Since this is not a diffusion or thermo-dynamics-based process, it opens upseveral possibilities. Using this technique, it is possible to implant, non-metallic (B,N,C,P) or metallic (Cr, Ti, Ni, Fe etc) onto metals, cermets , ceramics or polymers. Ionimplanted surface are in general more wear resistant, have better fatigue characteristicsand have better corrosion resistance.
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7. MECHANISMS AND MODELS OF DIE WEARAND FAILURE
The main forms of failure in hot and warm precision forging are abrasive and adhesivewears, oxidative wear, thermal fatigue, mechanical fatigue and gross cracking of dies. Inseveral cases, two or more of these mechanisms act together to wear down the die. Forfundamental insights into the different wear mechanisms, refer to appendix A.
7.1. WEAR
Several work exists in the literature that tries to characterize and model wear in hotforging. Some are based on process variables like forging area, weights and energy whilesome have taken a more fundamental approach to modeling. These models are providedin Table 7.1. With advances in finite element models and computing, it is possible touse fundamental material properties and process variables derived from FEM softwaresto model wear more universally. With the technological capabilities in mind, and withavailable data, it is possible to use Archard’s model provided in equation 7.1 to modelwear as a function of thermo-mechanical history of dies during a forging process andthe working hardness of the die material.
dtH
Vpkwear
i
ii∫×
= Equation 7-1
where p is normal pressure at a die locationV is the sliding velocity at any timeH is the hardness of the die locationAnd k is a constant dependent on several factors like billet material and scaleformation.
To obtain these in real time, we need the following pieces of information.• Hardness of dies with temperature; material data• Tempering curves; material data• Sliding velocities and distances; process data• Die pressures; process data
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Aston andBarry, 1972
HammerMean damage (x10-3) = 0.00686 x forging area + 0.0272 hammer energy - 0.1855 x forgingwt1/3 + 0.335 x spread - 0.011 x flash land area + 0.129 x flash metal escape – 0.557(84%)
Aston andBarry, 1972
HammerMean damage (x10-3) = 0.000261 x forging area + 0.000763 hammer energy - 0.00265 xforging wt1/3 + 0.012 x spread ration - 0.000694 x flash land area – 0.00266 (82%)
Aston andBarry, 1972
PressMean damage (x10-3) = 0.0284 x forging weight – 0.062 x die material - + 0.141 (83%)No. 5 steel = 1; 3Ni, 3Mo, 0.5Cr steel = 2; En40 not nitrided = 3
Aston andBarry, 1972
Mean damage (x10-3) = 0.000164 forging area + 0.000712 x flash land/gap - 0.00431(70%)
Aston andBarry, 1972
Hammer and pressMean damage (x10-3) = 0.00405 x forging area + 0.226 x forging wt1/3 - 0.019 x flash landarea + 0.00287 x flash weight (%) + 0.0184 x flash land/gap + 0.0666 (m/c factor: hammer =1, press =2) – 0.42 (72%)
Archard’smodel
dtH
Vpkwear
i
ii∫×
= for volume
Wear = k pd/H for depthk = constant,p = normal pressureV = velocityH = hardnessd = sliding distance
BudinskiWear controlhandbook
Axew41021.021023.0
−−−×=w = abrasion rate cm3/min
A = Structure parameter for a given tool steel (carbide size (µm) x volume fraction x carbidehardness (kg/mm2))
Thomas1969
R = 204 - (70 (%C) - 4 ( %Si) - 15 (%Cr)1/2 - 80 (Mo*)1/3
Mo* = %Mo + 0.5%V + 2%V + %NbR is the wear rate relative to H13 steel
(*) Model used by Bariani, 1996, Batit 1983, includes an exponent to the hardness Hm, Eriksen 1997, Painter1996. (Archard model). Use of computer simulation (Tulsyan, Shivpuri et al. 1993; Painter, Shivpuri et al.1994)
Table 7-1. Table summarizing different wear models found in literature
7.2. PLASTIC DEFORMATION
Plastic deformation occurs in areas of dies that experience intense pressures but doesnot have enough strength or hardness at the working temperatures to geometricallyresist the metal flow. This mode of failure occurs especially at sharp corners that havevery high surface to volume ratios, that experience high heating and subsequenttempering.
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Storen and others (Tulsyan, Shivpuri et al. 1993) provide a good criterion, given inequation 2.1, to avoid plastic deformation in forging dies. They say that one can avoiddeformation related die failure if we follow the following criterion.
BZ H×75.0pσ Equation 7-2
where HB is the Brinell hardness of the die material at the maximum
temperature, Zσ is the local normal pressures. In hot or warm forgings, the hardness
levels changes over the course of a run because of tempering effects. Also, theroughness of the surface, to some extent, affects the heat transfer coefficient, heatingand hence the hardness.
To model plastic deformation, , we need the following pieces of information.• Hardness of dies with temperature; material data• Tempering curves; material data• Heat transfer rates and contact times; process data
7.3. MECHANICAL FATIGUE
Appendix 11 presents an overview of low cycle fatigue that is the common form ofmechanical fatigue in forging operations. Low-fatigue test results is to plot the plasticstrain range ∆εp against number of cycles N. The plot of strain against number of cyclesusing a log scale for N results in a straight line that is known as Coffin-Manson law.These curves in a plastic strain range are very relevant for hot and forging applications.Figure 7.1 (Dieter 1986) illustrates a probabilistic S-N curve typically found inliterature.
Figure 7-1 S-N curve with probability lines or S-N-P (Dieter 1986)
The first model for strain controlled fatigue is known as Coffin-Manson law:
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2/1
2−=∆ fp N
Cε or
mfp N −
= =∆ε Equation 7-3
Were pε∆ is the plastic strain, Nf is the number of cycles to failure, C and m are
material constants.
Manson found later a graphic method to evaluate fatigue based on static tensile tests. Amethod called universal slopes was also presented by Manson, and it also included amodel that as following:
6.06.02/15.3 −− +=∆ ffu
p NDNE
σε Equation 7-4
were the first term is the elastic strain and the second term is the plastic strain.
uσ is conventional ultimate strength
E is the elastic modulus Nf is the number of cycles to failure
fε (represented by D) is conventional logarithmic ductility
The graphical representation of this equation is shown in figure 7.2, and it is a veryuseful way to evaluate fatigue using static tensile test data.
Figure 7-2. Illustration of the methods for estimating fatigue based in static properties(Manson 1972)
To evaluate mechanical fatigue, we need the following• Plastic strain; process data• Elastic modulus; material data• Ultimate tensile strength; material data• Ductility; material data
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7.4. THERMAL FATIGUE
Thermal fatigue is caused due to non-uniform temperature distribution between thesurface of dies and the interior. Any temperature differences between the surface andinterior results in strain differentials due to varying thermal expansion. When theresulting stress at the surface exceeds the yield strength of the material we haveyielding of surface layers. Extended cycling will result in crack initiation andsubsequent growth of thermal cracks. If the maximum and minimum temperatures adie location experiences, then low cycle fatigue occurs if
2
21
21
11
21
_2 EE
TTσνσν
α
−
+−
> Equation 7-5
where α is the mean coefficient of thermal expansionν is the poisson ratioσ is the stressindices 1 and 2 indicate the maximum and minimum values
Crack Initiation occurs when the following criterion (from Coffin-Manson) is met.
fCp
nF
N εε = Equation 7-6
where N is the number of cycles to crack initiationn is a material constant from 0 to 1
pε is the plastic strain range
C is a constant that is between 0 and 1
fε is the true deformation to fracture – a material property.
Crack growth occurs at a rate given by Equation 7-7.
]q
EETTa
qpa
dNda
2
221
1
111
12
σνσναρερ
−
−−
−−==Equation 7-7
wherea is the crack lengthN is the number of cyclesρ and q are positive constants dependent on material.
Any physical or process factors that impact the strain difference, impacts heatchecking.
To model thermal fatigue, we need the following pieces of information.• Ductility of dies with temperature; material data• Hardness of die; material data• Yield strength of dies; material data• Poisson ratio; material data
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• Thermal conductivity and thermal expansion coefficient; material data• Heat transfer rates and contact times; process data
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8. CLOSURE
The yield strength of the die material at the surface, exposed to the high contacttemperatures, is one fundamental property influencing die failure. It affects theoccurrence of the following failures:
• Plastic deformation• Thermal fatigue (crack initiation)• Mechanical Fatigue (crack initiation)• Wear
As the hot and warm forging dies reach high temperature during the working cycles, itis necessary that the hot yield strength stays stable during the hot work. The yieldstrength is the property that directly resists the working pressure, and keeps the diesworking in the elastic field macroscopically. This working condition will provide forgedpieces inside the geometrical tolerance range. However, critical regions of dies can besubjected to high stresses that can lead to plastic deformation. The thermal stress-strain, or the mechanical stress-strain state can cause thermal fatigue or mechanicalfatigue, respectively. As the amount of plastic strain is the driven cause for fatiguecrack initiation, high yield strength will reduce the amount of plastic strain, retardingor avoiding the crack initiation. The wear resistance is direct proportional to the yieldstrength, represented in the models by the hardness.
The ductility or plasticity of the hot work tool steel is other important property.Although the dies should work in the elastic regime, localized plastic deformations canoccur. The plastic deformations from thermal origin are difficult to avoid, especially inregions of high thermal load. In this case, the number of cycle to crack initiation will bedirect proportional to the ductility limit (area reduction in tensile test). Also, crackpropagation is believed to be controlled by plastic deformation in the low cycle fatigueregimes typical in warm and hot forging.
The toughness or fracture toughness is the materials ability to resist crack growth. Thisproperty will allow the dies to work at a higher stress-crack size without reach thecondition for fragile fracture that will leads to the die catastrophic failure. Lowtoughness also increases the crack growth rate.
From the fundamentals of wear, we can safely say that apart from hardness andsubsequent softening of die materials, pressures and the amount of sliding also affectthe failure rates. As discussed in previous chapters, pressure is primarily dictated byforging material, forging temperature, lubricant used and the geometry of the dies. Italso depends on the die closure, flash and the type of forging equipment used. Thepreform shape, lubricant and flash, control sliding distances dies experience duringforging. Surface hardness depends on, apart from the alloy composition andmicrostructure, coating or surface treatments used, lubricant, thermal cycling and toan extent, preheating.
Unfortunately, the interaction of the forging parameters and the wear and failure ratesis too complex to draw any direct correlation. For instance, forging temperature reducesthe wear resistance of the surface. Also, higher temperatures typically produce thicker,but not necessarily more adherent, scale. Scales, if adherent, increases wear. However,the loads felt by the die is lower because of the lower flow stress of the material at hightemperatures. Wear rates, here, would be controlled the relative magnitudes of theseeffects and can be predicted only by analyzing all these factors together.
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It is also important to note that dwell times, heating times and cycle times also havecontrasting effects on wear of dies. Increased heating time would increase scaleformation. In a 2 or 3-step operation involving descaling, this would not be a big factor.But process designs that employ single blows should pay special attention to heatingtimes and heating atmosphere. On a side note, thicker scales act as insulators andkeep the billet hotter. Sometimes, scales also act as lubricant, reducing loads. Byincreasing dwell time, when the die is in contact with the billet, the lower diesexperience substantial softening. But this also cools the billet in contact with the lowerdie, reducing sliding. Preheating, though effective in reducing the chance ofcatastrophic failure and thermal fatigue, increases wear by reducing the hardness andwear resistance.
As we see, there are many controlling factors that affect wear and die failure. It isimportant to evaluate wear as a cumulative result of all these process variables. Thereare several models proposed in the literature that try to capture some of theserelationships.
There are several interrelated parameters that affect the performance of forging dies. Inworking to improve hot and warm forging dies performance the fundamental step is toidentify what is the dominant failure mechanism. Only with this information, it ispossible to improve the correct properties, and optimize the correct process parametersthat will result in better die performance. During this process, it should be noted thatsolutions to reduce wear are different from those that reduce thermal fatigue andmechanical fatigue. Use of computer simulations (Tulsyan, Shivpuri et al. 1993;Painter, Shivpuri et al. 1994) could be a necessary first step to evaluate the conditionsat the die-billet interface before a good solution can be obtained.
Although there are several modes of die and tooling failure, because of safety reasons,the main concern of a tool designer is catastrophic failure of dies. In very few cases,designers employ predictive model to design tooling to avoid catastrophic failure. Thetendency, in forging industry, is to use material with low hardness and high toughness.However, beyond a point, toughness does not bring any benefit to the die life. Becauseof lock of good understanding of fatigue failure, the design and material choice is donevery conservatively.
Low hardness, because of lower alloying content, reduces the wear resistance. The wearresistance is function of the tool steel hot hardness, and the carbides in the matrix(amount, size, and hot hardness). However, generally these carbides reduce the ductilityand toughness. Carbides, necessary to resist wear, can be detrimental to resist diefracture (toughness) and fatigue (ductility). The wear resistance needs the material topossess hot yield strength. Thermal fatigue resistance is improved by both criticalductility is at room temperature and hot hardness. The alloying contents commandboth hot hardness and ductility. By carefully tailoring the microstructure and alloycontent to the application, it is possible to balance the different failure mechanismssuch that the tool life is highest. The alloying content in the matrix can be modified bythe heat treatment that controls the dissolved alloys in carbide form. However, higherthe undissolved carbides, higher is the wear resistance. But the thermal fatigueresistance reduces with higher amounts of undissolved carbide. As can be seen, there isa complex inter-relation between failures mechanisms and properties that need to beunderstood and applied correctly to improve the life of the dies for hot and warmforging.
We understand that the wear needs a more detailed evaluation under the conditionsusually found in hot and warm forging, and the same is valid for the interactionsbetween wear and the thermal fatigue. Based in these needs we developed a new test for
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applications to die forging at high temperature that can evaluate simultaneously thewear-thermal fatigue failure mechanisms.
It is essential to understand the mechanisms of die failure completely before we canattempt to increase die lives. As mentioned before, die wear is the major mechanism ofdie failure in high temperature forging, followed by mechanical fatigue. We emphasizethat the wear failure initiation can caused or increased by thermal fatigue, as indicatedby: micro observation of the die cavity, and the higher wear rate in dies with moresevere temperature cycle. However, in most cases, several modes of failure act inconjunction. In this section, we summarize the effect of various criterions on die failure.
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9. APPENDIX A – FUNDAMENTALS OF DIEFAILURE
Abrasive wear
Wear is the progressive loss of substance from the operating surface-of a body occurringas a result of relative motion at the surface (Czichos 1978) and also due to the tribo-chemical reactions (Gahr 1979). Predominant wear mechanisms present in metalforming can be classified as sliding wear mechanisms and non-sliding wearmechanisms (Stachowiak 1993). Sliding wear mechanisms include abrasive wear,adhesive wear, and delamination wear. Non-sliding wear mechanisms include solutionwear, diffusion wear, electro-mechanical wear and oxidation wear, or tribo-oxidationwear (Gahr 1979)
Abrasive wear arises when a hard, rough surface slides against a softer surface, digsinto it, and plows a series of grooves. The material originally in the groves is normallyremoved in the form of loose fragments, or forms ridges along each groove. The materialin the ridges is then vulnerable to subsequent complete removal from the surface(Stachowiak 1993)In hot forging conditions, abrasive wear may be compounded by the presence of hardthird phase particles in the interface. These particles may be hard oxides or scales,external-contaminating particles or other hard carbides dislodged from the die surface.Abrasive wear results in the displacement of die material from the surface. This istypically, either caused by the presence of hard particles between the die and thedeforming billet or protuberances embedded in the billet. The hardness of the particlethat causes the initial groove has to be equal to or greater than the hardness of the die.Figure 9.1 (Stachowiak 1993) illustrates a typical abrasive wear groove.
Figure 9-1. Appearance of plough marks caused by abrasive wear (Stachowiak 1993)
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Figure 9-2. Different mechanisms of wear in abrasion (Stachowiak 1993)
In abrasive wear, there are again different mechanisms present. Figues 11.2 show thesurfaces produced by the different mechanisms of abrasive wear - micro cutting, micro-fracture, micro-fatigue and grain pull out.
Adhesive Wear
The tendency of contacting surfaces to adhere arises from the attractive forces that existbetween the surface atoms of the two materials. If two surfaces are brought togetherand then separated, either normally or tangentially, these attractive forces act in such away as to attempt to pull material from one surface onto the other. Whenever materialis removed from its original surface in this way, an adhesive wear fragment is created(Rabinowicz 1995).
Adhesive wear occurs between two sliding surfaces, and the material is transferred fromone surface to another due to a process of solid-phase welding. The early experimentson adhesive wear were carried out with metals, were the process of adhesion wasreferred to as “welding”. By contrast, it is preferable in all cases to use the term“adhesive wear” rather than “wear by welding”. Figure 9.3 illustrates the welding andadhesion phenomenon.
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Figure 9-3 .a) A typical metallurgical weld. b) A typical adhesion joint (Rabinowicz1995).
Adhesive wear in hot forging can be very similar to welding since the interfacetemperatures can be as high as 1200°C. This phenomena is generally manifested in thedie picking up portions of the billet material and is accelerated when nascent metallicdie surface comes into contact with the hot billet. This may occur after the followingsequence of events:
• The lubricant layers and oxidation layers in the both die surface and billet surfacehave been removed by abrasive wear
• The base metal of the billet makes contact with the base die steel
• The reduced sliding of the billet material with respect to the die material is minimalbut the pressure is very high.
• The part is ejected form the die. Either a portion of the die material is removed withthe billet or a portion of the billet material adheres to the die. This second possibilityis more common since the die material is generally several times stronger than thebillet material.
Oxidation Wear
Research about wear in forging suggest that the main wear mechanisms in forging areabrasion and adhesion. Both mechanisms are classified as mechanical-sliding.Oxidation can affect wear in hot forging dies because the following reasons.The oxide film can influence the tool-workpiece interface, especially critical in adhesion
- the thermal fatigue cracks can start in oxidized points, and the cracks arefilled with oxides
- The dies can lose material by oxidation, due to the temperature range atwhich the dies surface operates
Summerville and Subramanian (Summerville, Venkatesan et al. 1995) shows (Figure9.4 a) an example of a hot forging punch with severe wear. The punch central region,
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that has low sliding, is more affected by oxidation and thermal fatigue. Figure 9.4 (b)and 11.4 (c) shows how critical the temperatures at the dies surface can be indetermining plastic deformation, phase transformation, oxidation and melting. Whilethe bulk dies temperature is usually around 350°F. The dies sub-surface temperatureusually reaches 1100°F, although the peak temperature at surface can reach as high as1650°F in certain applications. The Figure 9.4(d) shows thermal fatigue cracks filledwith oxides (Ribeiro 1998).
a) (Summerville, Venkatesan et al.1995)
b) (Doege 1994)
c) (Summerville, Venkatesan et al.1995)
d) (Ribeiro 1998)
Figure 9-4. a): Hot forging top blocker punch made form H13. b) Cross section of thepunch c) mottled interface d) Oxidation inside of thermal fatigue crack
Corrosive and oxidative wear occurs in a wide variety of situations both lubricated andunlubricated. Oxidative wear is the wear of dry unlubricated, or even lubricated, metals
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in the presence of air or oxygen. When thick oxide films are in the worn surfaces, lowwear prevails. When oxide films are absent or broken down, severe wear occurs, andadhesive wear might be the dominant wear mechanism. The fundamental cause ofthese forms of wear is a chemical reaction between the worn material and a corrodingmedium that can be a chemical reagent, reactive lubricant or air (Holmes 1972),(Quinn, Sullivan et al. 1980; Quinn 1983; Quinn 1991)In the hot forging temperature range the oxide films form very quickly, and the surfaceoxidation of the hot forging die can be detected even visually, as the die is submitted towork. Considering that an oxide film exists on the surface of the die cavity, it could bebeneficial or detrimental to the wear:The detrimental results will occur when this oxide layer forms and is removed in eachforging operation, causing oxidative wear. The oxide layer detachment could also permitadhesion, although the thick interlayer die-oxide-billet is still present. The beneficialresults will occur when the oxide acts as an insulation layer between the billet and thedie, preventing adhesive wearThe general result of pin disc test shows that the oxide film formation reduces the wearrate, it agrees with the expectation because the oxide film layer does not permitadhesion to occur. Note that in the pin-disc test the oxide layer before wear has only afew nanometers (Stachowiak 1993), and even during the tests it grows only a fewmicrometers (Quinn 1991); resulting in reduction of the wear rate. In the hot forgingprocess the die-billet interlayer is bigger than in the wear tests, is because it includes: arelatively thick billet-oxide layer, a relatively thick lubricant layer, an oxide layer in thesurface, due to its exposure to high temperatures.
The thickness of the hot forging inter-layer far exceeds the few atoms thick layernecessary to prevent adhesion, in basic pin-disc test; although the hot forging inter-layer is submitted to severe conditions (high pressure, velocities, and temperatures)that can brake and take way the oxide layer. What the pin-disc testes show is that anoxide layer reduce the wear, consequently we have to look for conditions that permit toform and keep some oxide layer in the hot forging dies surface. The lubricant-oxidelayers that separate the billet and the die can be considered thicker than 50 µm.Remember that the usual oxide thickness that provides wear reduction in pin-disctests, is less then 5 µm thick (Quinn 1991)
Colombier (Columbier 1965) presents many oxidation rates for allowing elements likeCr, Al and Si. They are presented in function of the allowing contents, and ortemperature. The behavior' analysis can help to project the surface modification bycoating or heat treatments. Chromium is per excellence the element to be used toobtain high scaling resistance; its effects begin to appear at around 5%. These 5% Crsteels are resistant to temperatures of to order of 600°C-650°C. The addition of 2% Al toa 6% Cr steel virtually suppresses scaling at 800°C over a test period of 100 hours.
Thermal Fatigue
The appearance of a fine network of cracks in the hot and warm forging dies is knownas heat checking. The hot and warm forging processes have a typical cycle that causesheating and cooling of the dies surfaces. The billet at high temperature is compressedinto the die cavity causing a drastic increase in the surface temperature. Thetemperature increase at the surface of the die causes its expansion. At the same time,the lower temperature of the die block constrains the expansion, generatingcompressive stress. Next, the part is ejected from the die and the dies are lubricated.During the cooling or lubrication, the process is reversed causing tensile stress. Theusual thermal cycle in hot and warm forging can result in thermal expansion thatresults in strain reaching plastic limits. When cracks are formed by repetitive change in
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temperatures the phenomenon is thermal fatigue. Figure 9.5 illustrates thisphenomenon.
Figure 9-5. Illustrates physical changes on the die surface that results in heatchecking (Norstrom 1991)
When the die surface start heating up, the parent metal that is still cold restricts thethermal expansion resulting in a compressive stress and hence a compressive strain.This compressive tress is initially in the elastic regime(A). If the temperaturedifferential is high, the stresses become plastic. This results in permanentcompressive strain on the surface. This state is indicated as state B in Figure 9.5.After the forging process, the dies are lubricated. Because of the coolant, the surfacecools faster than the bulk resulting in State C – where the surface stresses becometensile and the strains reach elastic limit. Beyond C , all the induced thermal strain isplastic. The next part made continues to thermally cycle the surface resulting in slowdeterioration of the surface.
The main factors that affect thermal fatigue are forging temperature, heating-coolingrates, time-temperature history, hot resistance of the die steel, temper resistance of diesteel, ductility and initial hardness of die steel, toughness, cleanliness and homogeneityof die steel and its heat treatment.
Temperature is the main parameter that controls thermal fatigue. There are differentways temperature influences thermal fatigues. It not only increases the thermal
(-)
YIEL
STRENGTH
(+)
STRAIN
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gradient that causes thermal expansion and stress but also reduces the materialstrength by causing metallurgical transformations. High temperature decreases the hotyield strength, and causes softening by tempering effects. The Larson-Miller curve, usedextensively to evaluate, represents the hardness variation as function of time andtemperature. High temperature also makes the surface weaker and more prone tooxidation. A drastic cooling rate in intricate dies can also lead to an excessive stress-strain state and result in gross crack by thermal shock.The yield strength at maximum cycle temperature is directly proportional to the amountof plastic strain in the die surface during the thermal cycles, and the plastic strain isthe cause of the heat checking. If the material suffers softening by temper effect, theplastic strain will increase in the same proportion. A correct die prediction has toconsider the softening and use the instantaneous material properties along the thermalcycles.
Higher values of yield strength or hot hardness will reduce the percentage of plasticdeformation that is ultimate cause of damage. Consequently high hot yield strength canreduce or even avoid thermal fatigue. However, for practical application in hot diecasting dies the increase in hot yield strength is limited if it is accompanied by:decrease in ductility, toughness or thermal shock resistance. When those properties arelow, they bring a risk of catastrophic failure or gross cracking by thermal shock.The die’s material needs to be able to resist the pre-heating temperature and the cycles’temperature without excessive lost of hardness. The effect of the cycle temperature isthe sum of the time’s cycle in the range of the maximum temperature. This property isrepresented by the temper curve, especially Larson-Miller type, and by the creep curves.The thermal fatigue tests confirm that the materials more temperature resistant presentbetter resistance when the other properties are similar.
The crack initiation is directly proportional to the ductility; as can be seen in the itemwith the mathematical models, and It also influences the crack growth. Other aspect isthat if ductility became too low it can cause crack growth rapidly leading to failure withfew cycles, more like and thermal shock or gross cracking. Experimental results showsthat tool materials with low ductility (less than 30% area reduction) presented morethermal fatigue, and even thermal shock, either when the other properties where in thesame range (Malm and Tidlund 1979), (Rostoker 1969), (Roberts and Norstrom 1987),
The Charpy-V notch impact test is a more common way to measure toughness. Due toits simplicity, there are many results available. The other test is the fracture toughnessor Kie that has fewer results available because is more difficult to perform and moreexpansive. Although, kic has the advantage of to be used in a quantitative way toevaluate catastrophic failure. The knowledge of Kic for a working condition permits tocalculate the admissible combination of stress and crack depth that do not causecatastrophic failure. The toughness and the thermal shock resistance seem to be muchrelated, because both express resistance to crack growing. The first case the mechanicaleffort drives the stress and in the second case temperature drives the stress (Norstrom,Johansson et al. 1981). The thought material permits a die to work with high level ofheat checking without the risk of fragile fracture or gross crack. The toughness it self(Kic or Charpy-V notch) is not included in the models for crack initiation or crackgrowth. The experimental results also do not show direct relation between toughnessand heat checking resistance.The cleanness, homogeneity and chemical composition are the most commons’ pointsrelated with property improvement and better performance in process. The properties intransversal direction tend to be lower, and as the dies suffer efforts in all directions, theimprovement in transversal properties will produce direct effects in dies’ performance.The ductility limit is the critical property for thermal fatigue resistance that is improvedby cleanness and homogeneity. The other property that improves is the toughness;
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related with gross crack and thermal shock. Several authors showed that ductilityimproves by cleanness and homogeneity, and that it also results in better thermalfatigue performance (Johnson and Hamaker 1968; Beck and Santhanam 1976; Okuno1987; Roberts and Norstrom 1987; Schmitd 1987; Nichols 1988).
The focus of this discussion about dies heat treatment is directed to its effects inthermal fatigue resistance. The yield strength, or hardness, and the ductility limits arethe controlling variables for thermal fatigue resistance, considering the material aconstant. As a lack in toughness can cause dies’ failure by gross cracking, the influenceof heat treatment in toughness is also considered. Nostrom (Norstrom 1989) considersthe ductility is influenced primarily by steel manufacturing quality (metallurgicaltreatments, etc.) while toughness is influenced chiefly by the final heat treatment of thedie itself (cooling rate in hardening, etc.).The low quenching rate is the most critical parameter that we found to cause lowtoughness or low ductility. The low cooling rates can be result of operational features ofthe largely used vacuum furnaces or can precaution against the risk of thermal shockcrack in intricate dies. There are two problems associated with low cooling rates, asfollowing:• Grain boundary precipitation
• Bainite formation
Becker (Becker, Fuchs et al. 1989) shows an example of reduction in toughness due tocarbides’ precipitation in grain boundary. Wallace, Roberts, and Norstrom (Wallace1989) (Roberts and Norstrom 1987) made a systematic evaluation of the toughness infunction of martensite, bainite, and grain boundary precipitation. The conclusions areclear and important and shows the following:
• The pure upper bainite has the same toughness range as martensite.
• The grain boundary precipitation reduces both martensite and bainite toughness.
• The use of 300° F preheating for dies will not increase the toughness of the poorlytreated steels.
Finally, the emphases in materials properties to resist thermal fatigue are:
• To have a high hot yield strength (to avoid or reduce the plastic deformation)
• To have a high tempering resistance (to keep the hardness along the work)
• To have a high ductility (to resist the plastic deformation)
Plastic Deformation
Plastic deformation is a die failure mechanism that occurs at regions of the die that issubjected to extreme pressure and temperatures. This occurs when the local stressesresult in die stresses exceeding the local yield strength of the die material. Typical areasof the die that are prone to plastic deformation are sharp corners of the dies and thinprotuberances that trap a lot of heat during the forging process.Since extreme pressures and temperatures cause this mode of failure, increased localforging stresses will increase the chance of plastic deformation. Consequently, all designand process criterion that impact stresses and die temperatures have an effect on theplastic deformation of dies. Of these, the forging temperature, size and geometry of theforging, lubricant used, forging cycle times, type of equipment used and the type offorging (whether it is conventional or flashless) are the most important factors. These
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parameters either increase local stresses or reduce the strength of the die by thermalsoftening or a combination of both.
From basic metal forming theory, it is well known that the hardness of the material isabout 3 times the yield stress of the material. However, when the thickness of the platereduces, the measured hardness measured drops to less than 3 times yield stress. Forthese sections, the hardness measured could be as low as 1.15 times the local yieldstress. Figure (6) by Schey (Schey 1987) illustrates this phenomena. Similarly, in metalforming, features like sharp corners and projections that geometrically “thin” andpossess less rigidity tend to deform plastically first. Also, these features tend to heat upquickly because of high exposed surface area, resulting in reduced local hardness.Simulation of the forging process and analyzing the die stresses and comparing it to thehot strength of the die material may be the most accurate way to predict plasticdeformation.
Figure 9-6. Illustration of geometry effect on normal uni-axial stresses required toindent a slab (Schey 1987)
Hence, for a specific process, die material selection becomes very important. Hotstrength or hot hardness of the die is the most important property necessary towithstand plastic deformation. Die steels with high hot strength will resist plasticdeformation better than steels whose strength drops drastically with temperature. Forinstance, for dies whose mode of failue is plastic deformation at high temperatures,nickel based superalloys like Nimonic and Inconel 718 may be good substitutes.Plastic deformation is also very prominent at microscopic levels. Plastic deformationoccurs at the interface of die and billet. Because of the high temperatures typical in theinterface, die material on the surface becomes extremely soft and pliable. Figures 11.711.8 and 11.9 by Summerville and coorwers (Summerville, Venkatesan et al. 1995)shows various levels of plastic deformation found in the die-billet interface.
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Figure 9-7. Examples of severe plastic deformation at the die surface (Summerville,Venkatesan et al. 1995)
Figure 9-8. Example of surface plastic deformation (Summerville, Venkatesan et al.1995)
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Figure 9-9. Example of surface plastic deformation (Summerville, Venkatesan et al.1995)
Mechanical Fatigue and Cracking in Hot and Warm Forging Dies
Several factors interact in a very complex nature to affect the performance of toolingduring a forging process (Figure 9.10). One of the failure modes very common in forgingtooling is mechanical fatigue and cracking. Mechanics of mechanical fatigue is welldocumented. The process can be divided in three steps:• Crack initiation
• Crack growth
• Catastrophic failure
The forging dies are subjected to high pressures in order to fill the die cavity. In highvolume batch production like a forging operation, dies are subjected to repeated loadingand unloading (Figure 9.11). Similar to thermal fatigue, mechanical fatigue is causedby alternating stresses that cause strains at crack tips regions exceed the plastic limit.Repeated loading of the crack results in the advancement or propagation of cracksresulting in gross cracking.
There are several models that are present in the literature that model the fracturebehavior of materials. Appendix D details these models and provides different ways ofmodeling fatigue crack initiation, propagation and catastrophic failure. Figure 9.12(Dieter 1986) represents the different phases of crack growth in a fatigue failure. Thesemodels have enabled researchers to quantify objectively, the fatigue resistance ofmaterials at low and high stress levels. These models try to identify the number ofcycles necessary to initiate a crack or necessary to propagate a crack by a givenamount.
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Figure 9-10 Schematic interaction between the parameter in hot forging and thecracking (Knorr 1993)
Figure 9-11 Illustration of a critical region in extrusion dies, where the fillet radius issubject to tensile stress (Cser, Geiger et al. 1993)
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Figure 9-12. Representation of the fatigue crack propagation (Dieter 1986)
Figure 9-13 Tulsan (Tulsyan, Shivpuri et al. 1993) presents a curve form Storen andothers for different tool steels and heat treatment. a) fracture toughness properties asfunction of the working temperatures and the heat treatments b) materials and heat
treatment list
Using these models, there has been a lot of testing done to characterize high and lowcycle mechanical fatigue. However, most of the testing have been performed on alloysthat are not commonly used in high temperature forging. Most of the testing have beendone at room temperature. The hot and warm forging dies works in temperatures higherthan room temperature. Dies are preheated to improve toughness. Also, the die surfaces
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heat up due to the contact with the hot billet. It is necessary to consider the workingcondition when looking at the fatigue curves.
The fatigue test at high temperature showed that environment effects the fatigueresistance. The tests involving air and vacuum atmospheres showed that thephenomena in dependent of the temperature and the frequency (Figure 9.13) (Tulsyan,Shivpuri et al. 1993). Studies by Storen have indicated that oxidation also increases thefatigue damage. In hot and warming forging operations the dies are also in contact withspray that contains water and lubricant, and that could cause also affect fatigue.
Fatigue data that could be applied towards metal forming applications should considerthe following• Tool steel, heat treatment, surface treatment, and coatings applicable to hot and
warm forging
• Fatigue test temperature at dies working temperature range• Cycle time and frequencies compatible with forging equipment and forging
production rate• Evaluation of the effect of water spray, and water-lubricant spray
The following figures (Figure 9.14, Figure 9.15, Figure 9.16, Table 11.1.) illustrates thetemperature influence.
Figure 9-14 Results in air and vacuum atmospheres, showing the ambient effect atthe fatigue resistance in high temperatures (Salomon 1972)
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Figure 9-15 Correlation of high and low cycle fatigue data for solution treated type304 stainless steel as a function of alternating stress (Soo 1972).
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Figure 9-16 Effect of Temperature on Fatigue-Crack-Growth behavior of 2 1/4 Cr-1Mosteel (Viswnathan 1989).
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Figure 9-17 Variation of fatigue-Crack-growth rates as function of temperature at ∆K= 30Mpa (m)1/2 (Viswnathan 1989)
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Table 9-1. Results for crack propagation typo Paris da/dN for the constants “C, n”. b)Materials compositions for the hot tool steels used (Schuchtar 1988).
From the models, we see that toughness is essential in avoiding gross cracking.Toughness is the ability of the material to withstand large plastic strains. Whileincreasing yield strain increases the fatigue resistance of dies, it frequently reduces thehardness of the dies and its wear resistance. To avoid catastrophic failure, one mustreduce the mechanical stress, increase the material toughness and avoid die-manufacturing processes that induce stress cracks. Geometrical features of the partand die, that increase stresses also reduce the fatigue life of the tooling. These could be
• Sharp fillets and corners
• Thin rib-like sections (that increase forging pressures)
• Geometry of part and flash
• Forged material
The increase in toughness is generally associated with reduction in yield strength – afunction of alloying elements and microstructure. As yield strength is essential in givingthe material its wear resistance and thermal fatigue resistance, increasing mechanicalfatigue resistance is detrimental to the die’s resistance to other modes of die failure.A good approach to selecting the right material would be to evaluate the die life usingthe stress-strain state in the dies cavity. Analytic methods like slab or upper bound, orFinite Element Methods (FEM) can give stresses and strains required in the analysis.The FEM is applicable to more complex geometry and gives field distribution. Thesecond step is to use the contact pressures to calculate de stress-stain distribution in
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the dies, with analytic or FEM methods. Finally, a more precise evaluation shouldconsider the strains due to mechanical stress and due to thermal stress. Using thestress and strain variations and the appropriate constants, the fatigue life needs to becalculated.The final goal when design the forging process is to obtain a sound forged piece withoutdies’ catastrophic failure. However, it is also necessary to keep a good die life byavoiding excessive die wear and thermal fatigue. Some examples of actions to reducemechanical fatigue are:• Increase corner radii, if possible• Use inserts and prestress dies in rings• Use correct die block heat treatment (low quenching rates tends to cause carbide
precipitation in grain boundary that reduces the toughness). The inserts can alsohelp to have small sections to quench.
• Use surface heat treatments, coatings or surfacing that allow to have high blocktoughness combined with higher properties at the cavity surface
• Use die preheating that increase the toughness• Avoid die overloading due to process variationsSeveral die geometry effect in the stress state were presented by Knorr and Shivpuri(Knorr 1993), from Mareczek and Stute-Schlamme, Erlmann et al., and some arepresented in the figures 11.16 presents examples of this effects.
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a) Die deflection due improper support a-flat, b) convex c- concave
b) a- concave support 1mmb- convex support 1mmc- flat supportDie bottom surfaceFillet radiusDie wallCorner radius at flash
c) Influence of die geometry on stressd) Influence of die geometry on stress
Figure 9-18 series of cases with stress concentration in forging dies presented byKnorr (Knorr 1993). a) – b) From Erlmann at al.; c) -d) From Mareczek
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10. APPENDIX B – WEAR INDICES OF VARIOUSDIE MATERIALS
Figure 10-1. Abrasion resistance of several tool steels versus structural parameter(wear index) (Blau 1992)
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Figure 10-2. Variation of wear index with die hardness at room temperature(Kannapan 1969; Kannapan 1970)
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Figure 10-3. Wear resistance of .55% C die steel with hardness, % Cr and heattreatment. 1 indicates (Kannapan 1969; Kannapan 1970)
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Figure 10-4. Wear test results using different die materials (Bramley, Lord et al. 1989)
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Figure 10-5. Wear test results using different die materials (Bramley, Lord et al. 1989)
Figure 10-6. Variation of wear index with different die steels. The graphs alsoillustrate the effect of different forging steel (Thomas 1970)
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11. APPENDIX C - PROCESS EFFECT ON DIELIFE
Die Design
Die or cavity design plays a very important role in the failure of dies. The variousaspects of cavity that impact the die wear and failure are:
• Cavity size / diameter or width of part• Flash design and flash less designs• Corner radii and draft angles• Preform design
Prime cause of die wear is sliding and high normal stresses on the die. As a result, anyaspect of die design that impacts these will affect die failure. Aspects of the die cavitythat impact the die pressures are the cavity depth, size of the part, flash thickness, sizeof flash gutter, part geometry and the preform or blocker design. Heinemeyer(Heinemeyer 1976) studied the relationship between the die life and the cavity depthusing 160 part geometries and 2300 production runs. The trend he obtained from hisstudies is shown in Figure 13.1. Heinemeyer also reports the effect of energy and loadon die wear (Figure 11.2). However, Heinemeyer’s deduction that nominal load increasedie life may be inaccurate. It is well known fact that higher the loads, higher are thenormal pressures on the dies and higher is the ensuing wear.
Aston and Muir (Aston 1969) and later, Aston and Barry (Aston, Hopkins et al. 1972)did a series of analysis of data from forge shops in England. They derived empiricalrelationships between damage and a series of design and process variables. From theirstudy, they found that damage increases with the forging weight (Figure 11.3). It ispossible, there were other factors that changed the nature of relationship between thenominal load and wear. Also, the damage increases with the size of forging (Figure11.4). Aston and others also noted that damage increased with weight, draft angle, anddropped with increasing radius and increasing contact area (Figure 11.5). Again, this isfallacious because, it is well known that, if all other factors are kept constant, thetonnage required is directly related to the size of the forging.
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Figure 11-1. Effect of maximum cavity depth on die life (Heinemeyer 1976)
Figure 11-2. Effect of nominal load and energy on average die lives (Heinemeyer 1976)
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Figure 11-3. Effect of forging weight on die damage (Aston 1969)
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Figure 11-4. Variation of die damage with size of forging (Aston and Barry 1972)
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Figure 11-5. Effect of forging weight, fillet radii, draft angles and contact area on wearof forging dies (Aston 1969)
Flash design is extremely important in defining the loads in a forging process.Consequently, flash design becomes a factor that affects the die life also. A highernominal load on the press directly translates into higher stress on the die. This coulddamage the die in many ways. High stress cycling could result in mechanical fatigueand cracking. Also, higher normal pressures on the die surface will result in higherabrasive wear. However, restricting flash size is essential to ensure good fill.
Billet alloy
The forging pressure is directly proportional to the wear damage. Billet materials thathave high yield strengths at high temperatures, will result in high die pressures thatwill result in more wear. Figure 11.6 by Thomas illustrates this effect. He reports thatas the carbon content increases, the die stress and the die wear increases. High carbonalso forms more carbides that results in higher abrasion of dies.
Alloy steels like stainless steels also result in high wear because they form verydestructive oxide layers that are very adherent. These oxide layers can not be brokeneasily and increase the wear. However, no quantitative results exist in this area. Otherexamples of material with high yield stress that causes accentuated wear are alloys forturbine blades and engine valves (Tulsyan, Shivpuri et al. 1993; Painter, Shivpuri et al.1994).
Figure 11-6. Effect of various tool steel on die wear (Thomas 1970)
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Part tolerances and surface requirements
Part tolerances and machining allowances built into a forging invariably play animportant role in deciding when forging dies need to be pulled from service. Forexample, the die life in forging a part with .100” stock using conventional dies will bemuch more than the die life expected in forging a similar sized toothed part with netshaped surfaces. Part tolerances, hence, should always be specified with the goodunderstanding of the process and the forging application.
Billet Temperature
Loosely speaking, forging at a lower billet temperature has the same effect as using ahigher carbon alloy. Forging at lower temperatures increases the flow stress of the steeland the load required to forge a part. It also decreases the formability of the steel.However, forging at lower temperature reduces the tempering the die steel experiencesby reducing the surface temperature of the dies. Increase in billet temperatures canalso dramatically increase the friction (Ribeiro 1993). For a specified flow,, frictionincreases die wear by increasing the loads or normal pressures on the die surface.
Thomas (Thomas 1971) shows that the increase in bulk die temperature increases thewear (Figure 11.7). The results presented by Thomas (Thomas 1971) illustrate very wellthe variation of wear in function of temperature. Thomas also found that increasing thestock temperature reduces the wear after a critical point. The effect of stocktemperature on wear is complicated. It is the combination of competing effects offriction, load and die tempering. Figure 11.8 illustrates this very well.
Doege (Doege 1994) found wear reduction when reducing the forging temperature form1373 K (1100° C) to 1173 K (900° C). This is in agreement with findings of Thomas.
Figure 11-7. Effect of bulk temperature and stock temperature on wear of hammerdies (Thomas 1971)
Forging and Heating Equipment
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Choice of forging equipment plays a decisive role in determining life of a die. Forgingequipment – hammer, mechanical press or hydraulic press – determines the strain ratesand loads experienced material flow, forging duration and the incidental die temperingeffects. Aston’s (Aston 1969) findings presented in Figure 4.25 illustrates this effect.Aston found that the average life of dies in hammer forging of 5 different part familieshe studied, is much higher the die lives in press forged parts.
Figure 11-8. Relative die damage of five different part families when forged in ahammer and a press (Aston 1969)
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a)
b)
Figure 11-9. Effect of dwell time on the wear volumes observed (Rooks 1974)
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Figure 11-10. Die wear for three different dwell times for a) H.50 dies and b) No. 5 toolsteel dies (Rooks 1974)
Presses have higher dwell times compared to hammer because of the forging speeds.Rooks (Rooks 1974) studies also illustrate the effect of dwell times on the wear depths.Rooks found that at lower dwell times, the wear depths were higher for lubricated diesand lower for non-lubricated dies (Figures 13.10). Note that the trends are different forthe lubricated dies and non-lubricated dies.
By the same analogy, because of tremendous die tempering that is found in hydraulicpress forgings, die wear would be most severe when these presses are used. However, insome applications like forging extrusions, where low speeds and long strokes areessential to the viability of the process, hydraulic presses become a necessity.The wear of dies on the press was about three times as great as on the hammer for thesame number of identical forgings [Blau, 1992] (Bishop 1957). However, it should benoted that the high contact time that usually causes severe hardness loss inmartensitic steels could increase hardness in precipitation hardening steels (Nagpal1976).Press or forging speed also increases the velocity or forging strain rates. In hot forming,this increases the die stresses as well as the sliding velocities. Also, as the slidingvelocity increase, the heat generation at the interface increases. Dwell time is defined asthe time that the dies and the billet are in contact under pressure.
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Figure 11-11. Effect of scaling time on adhesive wear characteristics (Thomas 1971)
Selection of heating equipment also affects the die life in a subtle manner. Heatingduration and the presence of inert atmosphere affects the type and amount of scalesformed. Thomas found that the percentage of adherent scale drops with increase inheating times (Figure 11.11). Figure 11.11 also shows that a heating atmosphere whichis richer in oxygen reduces the adherent scale because of higher oxidation found. Thishas the same effect as increased heating times. Adherent scale increases wear bymaking the descaling process less effective. Box furnace and slot furnaces increase theheating duration, thus helping reduce the adherent scale and reducing wear. Inductionheaters, on the other hand, heat the billets fast and may produce a very thin adherentlayer of oxide that may be detrimental to the life of dies. This could however be reducedby the use of inert atmosphere. The use of controlled atmosphere in the heating furnacecan have practical applications in hot-warm forging to reduce the oxidation rate.
The effect of furnace selection on die wear of extrusion dies is presented in Figure 11.13(Doege, Seidel et al. 1996). It should be noted that, in extrusion, there is no de-scalingprocess that results in more scales acting as abrasives. It should also be noted thatscales act as a thermal barrier between the dies and the work piece. Figure 4.31 showsthe variation of oxide thickness on the die temperature (Kellow, Bramley et al. 1969).
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Figure 11-12: Oxide formation on 080M40 (En8) steel billets heated to 1100°C (Dean1974).
Figure 11-13. Scale formation and adherence as function of heating time and furnaceatmosphere (Thomas 1971)
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Figure 11-14. Effect of furnace selection on die wear of extrusion dies (Doege, Seidelet al. 1996)
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Figure 11-15. Effect of scale thickness on the die surface temperature (Kellow,Bramley et al. 1969)
Die Preheating
Preheating dies help reduce the chance of gross cracking by increasing the toughness ofdies. Appendix B provides toughness information on several die steels at differenttemperatures. Preheating also reduces thermal fatigue by reducing the thermal gradientbetween the surface layers and the bulk of the die steel. However, by increasing thebulk die temperature, the bulk hardness and the surface hardness of the die drops.This results in increased die wear. Netthofel (Netthofel 1965) shows this effect in Figure11.16. Netthofel’s results obtained from wear testing experiments explained, alsoprovides some insights into the effect of forging temperature on wear.
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Figure 11-16. Effect of forging temperature on the wear depth after forging 4000pieces (Netthofel 1965)
Die preheating also affects phase transformation at die surface. The phase transformsfrom martensite-austenite during the heating followed by the transformation back tountempered martensite (Okell and Wolstencroft 1968). The dies’ surface micrographsshows white layer, micro cracks, and micro plastic deformation (Summerville,Venkatesan et al. 1995; Doege, Seidel et al. 1996).
Lubrication
Graphite-water and graphite-oil are the most effective lubricants in hot die forging.Graphite-free lubricants cause die life reduction when compared with graphite baselubricants, specially for high sliding distances. (Doege, Seidel et al. 1996) Lubricationdecreases the friction and hence pressure as well as increases sliding. For a givenmetal flow, lubrication decreases pressures and die wear. Thomas (Thomas 1971) foundthat the wear in three times less for lubricated conditions. However, there are studies inthe literature that show the contrary. This is because, decreasing the friction increasesliding, if unrestricted, thus increasing wear (Singh, Rooks et al. 1973). Thisphenomena is clear from figures 13.17, 13.18 and 13.19.
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Figure 11-17. Variation of wear pattern of the top and bottom dies with lubrication(Singh, Rooks et al. 1973)
Figure 11-18. Variation of wear rate with lubrication (Singh, Rooks et al. 1973)
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Figure 11-19. Variation of wear volume with die bulk temperature for lubricated anddry forging (Singh, Rooks et al. 1973)
Forging Cycle Times
Doege (Doege, Seidel et al. 1996) found three times less wear when an additional diecooling was applied before the conventional die lubrication. Additional cooling times,between two blows act as additional cooling. Additional cooling times between forgings,decrease temperatures in critical corners dramatically and help retain hardness. Longcontact time under pressure conditions, as in low speed forging or when the pieceadhere to the cavity, increases the die surface temperature and tempering effects. Bothphenomena increase wear.
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