U.S. patent application number 14/359105 was filed with the patent office on 2014-11-13 for alloy for high temperature tooling applications.
This patent application is currently assigned to TUBITAK. The applicant listed for this patent is TUBITAK. Invention is credited to Fahri Alageyik, Yucel Birol, Osman Cakir.
Application Number | 20140334968 14/359105 |
Document ID | / |
Family ID | 47557408 |
Filed Date | 2014-11-13 |
United States Patent
Application |
20140334968 |
Kind Code |
A1 |
Cakir; Osman ; et
al. |
November 13, 2014 |
ALLOY FOR HIGH TEMPERATURE TOOLING APPLICATIONS
Abstract
Thermal fatigue is the predominant mechanism that limits the
service life of dies in semi-solid forming of steels since the
feedstock to be shaped has a paste-like character. A novel alloy,
more resistant to these conditions than any other alloy, has been
developed. Eutectic carbides in Stellite alloys are replaced in
this novel alloy with molybdenum-rich intermetallic compound
particles between dendrites. This novel alloy offers at least 3
times longer service life with respect to Stellite 6 alloy that has
been tested under conditions that mimic the steel semi-solid
forming process and has been identified as the most suitable. The
exceptional performance of the novel alloy is attributed to its
outstanding resistance to oxidation and to softening at elevated
temperatures and to its cobalt based matrix free from the hard and
brittle carbides that have a negative impact on crack growth
process.
Inventors: |
Cakir; Osman; (Kocaeli,
TR) ; Alageyik; Fahri; (Kocaeli, TR) ; Birol;
Yucel; (Kocaeli, TR) |
|
Applicant: |
Name |
City |
State |
Country |
Type |
TUBITAK |
Ankara |
|
TR |
|
|
Assignee: |
TUBITAK
Ankara
TR
|
Family ID: |
47557408 |
Appl. No.: |
14/359105 |
Filed: |
November 16, 2012 |
PCT Filed: |
November 16, 2012 |
PCT NO: |
PCT/IB2012/056511 |
371 Date: |
May 18, 2014 |
Current U.S.
Class: |
420/440 ;
420/585 |
Current CPC
Class: |
C22C 19/07 20130101;
C22C 30/00 20130101 |
Class at
Publication: |
420/440 ;
420/585 |
International
Class: |
C22C 19/07 20060101
C22C019/07; C22C 30/00 20060101 C22C030/00 |
Foreign Application Data
Date |
Code |
Application Number |
Nov 18, 2011 |
TR |
TR 2011/11434 |
Claims
1. A cobalt based alloy with no carbon and 9-15 wt % iron in its
chemical composition.
2. An alloy according to claim 1, with a ferritic dendritic matrix
phase and molybdenum-rich intermetallic particles at dendrite
boundaries.
3. An alloy according to claim 1, the extent of softening after
5000 cycles in thermal fatigue test of which, is less than 15%.
4. An alloy according to claim 1, containing 0.72 wt % Si, 1.58 wt
% Mn, 29.22 wt % Cr, 5.50 wt % Mo, 3.77 wt % Ni, wt % 11.50 Fe, the
balance being Co.
Description
[0001] Tool materials that can withstand the process conditions are
employed in forming operations. The tool materials to be used in
tooling for high temperature forming operations must withstand the
high temperatures, the mechanical and thermal loads encountered at
these temperatures, oxidation and loss of strength. The present
invention offers a new tool material for the semi-solid processing
of steels, an attractive near-net shaping process that has not been
commercialized until today due to a lack of suitable die
materials.
[0002] The die materials intended for semi-solid forming of steels
must withstand thermo-mechanical cycles at elevated temperatures,
wear and oxidation. The tool material described by the present
invention offers a three-fold improvement in the service life with
respect to that provided by cobalt-based alloys that have been the
material of choice and widely used until today. Hence, it is a very
attractive material for the manufacturers of forged steel parts.
The material of the present invention, whether employed as the die
material or as a hardfacing alloy, offers a very significant
opportunity for semi-solid processing of steels, an attractive
near-net shaping process that has not been commercialized until
today due to a lack of suitable die materials.
STATE OF THE PRESENT ART
[0003] Tooling issue is a very critical one since tool assemblies
are responsible for one fifth of the part cost in plastic forming
operations [1]. Dies employed in semi-solid forging of steel parts
are among tooling that are faced with the most severe conditions
[2-5]. Stresses originating from thermal cycles are essentially the
major life-limiting factors since the mechanical stresses are
limited owing to paste-like features of the feedstock to be shaped
in the forming dies [6]. Hot work tool steels are not at all
suitable for tooling to be employed in semi-solid forming of steels
in spite of the fact that they are very attractive from a cost
standpoint [7-15].
[0004] Cobalt-based alloys are attractive candidates for high
temperature tooling applications owing to an exceptional
performance in a wide temperature range [16-22]. Stellite 6, a
popular cobalt-based alloy which can be used as the die material or
as a hardfacing alloy, has been shown with the present art to offer
a three-fold improvement in tool service life with respect to hot
work tool steels [23, 24]. However, the hard and brittle carbides
dispersed in between the ferrite dendrite arms have played a key
role in crack growth and led to an increase in the crack growth
rate. Additionally, carbides were found to be more susceptible to
oxidation in high temperature wear tests and have impaired the wear
resistance of cobalt-based Stellite alloys by increasing the wear
rate [25]. The high cost of cobalt based alloys is the major
impediment that limits their wide spread use.
DETAILED DESCRIPTION OF THE INVENTION
[0005] A novel cobalt based alloy has been developed in the present
invention to be employed in the semi-solid shaping of steels. This
novel alloy has been designed without carbon with a consideration
of the negative impact of the carbides in stellite alloys on crack
propagation under thermal cycles encountered at elevated
temperatures. Another critical feature of the alloy of the present
invention that is different from its counterparts is its much
higher iron content (up to 12 wt %).
[0006] This compositional adjustment is intended to account for the
dilution of the hardfacing layer with iron from the underlying
steel substrate during hardfacing operations.
[0007] Another motivation for this change is to provide a
substantial cost reduction in the cobalt based alloy.
[0008] The alloy of the present invention does not contain carbon.
Besides, its iron content is much higher than its counterparts.
This feature not only improves the cost aspect but also accounts
for the dilution effect often encountered in hardfacing of tool
steels.
[0009] Owing to the chemical composition of the present alloy, the
eutectic carbides typical of stellite alloys are replaced by
molybdenum-rich intermetallic compound particles dispersed at
dendrite boundaries.
[0010] The alloy of the present invention offers a three-fold
improvement in the service life of dies with respect to those
manufactured from stellite 6 alloy that has been identified to be
the best die material tested until today under semi-solid steel
forming process conditions. This exceptional performance of this
novel material is attributed to its resistance to high temperature
oxidation and to temper softening as well as to a cobalt-rich
matrix phase free from the hard and brittle carbides that impair
the crack growth resistance.
[0011] Carbon-free cobalt based alloy of the present invention was
prepared by melting commercial purity elements in a 2 kg capacity
induction furnace under vacuum. The alloy melt thus obtained was
subsequently cast into boron nitride coated permanent mould between
1580-1600.degree. C. under vacuum.
[0012] The microstructure of the alloy of the present invention
consists of a ferritic dendritic solid solution matrix phase and
molybdenum-rich intermetallic compound particles at interdendritic
sites (FIG. 1).
[0013] Carbides in stellite alloys are replaced in the alloy of the
present invention with molybdenum-rich intermetallic compound
particles. The hardness of these particles is nearly two times
higher than that of the matrix phase. The hardness of the matrix
and the molybdenum-rich particles were measured to be 382.+-.52 HV
and 700.+-.143 HV respectively.
[0014] Samples of the present alloy, 25 mm.times.25 mm.times.20 mm
in size, were submitted to thermal fatigue testing in order to
identify their performance. The minimum and the maximum
temperatures of the thermal cycle were selected to be 450
centigrade and 750 centigrade, respectively, with a consideration
of the die surface temperatures encountered in the conventional
forging and semi-solid forging operations. The former temperature
is the temperature to which the die is heated to in plastic forming
operations while the latter is the highest temperature measured at
the surface of the die during semi-solid forging of steels. The
front face of the thermal fatigue test sample was heated to 750
centigrade within 30 seconds with an oxyacetylene torch and
subsequently cooled to 450 centigrade within the next 30 seconds
with forced air blow (FIG. 2). The temperature of the front and
rear faces of the thermal fatigue test samples were measured with
K-type thermocouples fixed into 3 mm diameter holes drilled at 0.1
mm from the respective surfaces. Thermal fatigue tests were
terminated as soon as oxidation author thermal fatigue cracks were
detected on the front face. The damage introduced with thermal
cycling were evaluated qualitatively with optical and stereo
microscopes.
[0015] The lowest and the highest temperatures at the rear face of
the thermal fatigue sample were measured to be 486 and 580
centigrade, respectively, when the temperature of the front face
was cycled between the minimum and maximum temperatures of 450 and
750 centigrade. These temperature differences between the front and
rear faces have set up thermal gradients across the section of the
samples. The temperature difference between the front and rear
faces of the sample is as much as 192 centigrade within 27 seconds
of the start of a typical thermal cycle. The maximum tensile and
compressive stresses acting on the front face of the thermal
fatigue samples were estimated to be 472 MPa and 210 MPa,
respectively, with a consideration of the thermal gradients
measured across the section of the samples.
[0016] The front face of the thermal fatigue sample prepared from
the alloy of the present invention faces the highest temperatures
and thus the highest thermal stresses during thermal cycling and is
shown in FIG. 3. The thermal fatigue resistance of this sample
revealed no signs of thermal fatigue cracking, until 13000 thermal
cycles and was thus identified to be exceptional. Several minor
lines linked with the blistering and wrinkling of the surface oxide
film were detected on the front face during the regular checks
performed every 500 cycles after a total of 13000 thermal cycles
(FIG. 3a). It has taken these minor traces another 3000 thermal
cycles to develop into legitimate thermal fatigue cracks (FIG. 3b).
Finally, a surface crack of considerable size has formed by the
spallation of the surface oxides at the intersection of two surface
wrinkles after a total of 16000 thermal cycles (FIG. 3b). The
thermal fatigue test was terminated at this point and the front
face of the sample was prepared for investigation with standard
metallographic practices. The surface crack was found to go through
the grains instead of through the grain boundaries (FIG. 3c).
[0017] Thermal fatigue test had to be terminated after 5000 cycles
when the hot work tool steel sample coated with Stellite 6 alloy
hardfacing tested under exactly the same conditions. Thermal
fatigue cracks had already traversed the 2 mm thick stellite 6
surface layer entirely at this point. Thermal fatigue cracking was
found to start after 4500 cycles and the thermal fatigue crack was
found to grow nearly 2 mm during the next 500 cycles during the
regular 500 cycle inspections [24].
[0018] It takes 13000 thermal cycles for thermal cracking to start
and another 3000 cycles for the thermal fatigue crack that has
formed to grow only 1 mm. (FIG. 4).
[0019] This simple comparison evidences that the crack growth rate
in the alloy of the present invention is at least 10 times lower
than that in the Stellite 6 alloy. It is fair to claim from the
foregoing that the alloy of the present invention offers not only
an outstanding thermal fatigue crack initiation but also to thermal
fatigue crack growth resistance. This exceptional crack growth
resistance of the alloy of the present invention is attributed to
its microstructure free of brittle carbides that were shown to
encourage crack propagation in Stellite 6 alloy [24, 26].
[0020] The potential of the alloy of the present invention under
thermal cycling conditions at elevated temperatures is appreciated
more when the performance of the hot work tool steels under the
same conditions are considered. Hot work tool steels could
withstand the same conditions only for 1000 thermal cycles due to
surface degradation linked with severe oxidation [23, 24]. Deep
cracks filled with voluminous oxides were noted on the front face
of the hot work tool steel sample when the thermal cycling was
terminated after 15000 cycles.
[0021] Stellite 6 alloy that has been identified to be the most
successful high temperature alloy until today, whether employed as
the die material itself or as weld overlay surface layer, could
withstand these cycling conditions for only 5000 cycles without
cracking on the front face [23]. It is thus fair to claim that the
alloy of the present invention offers at least 3 and at least 10
times longer service life with respect to Stellite 6 alloy and hot
work tool steel, respectively.
[0022] The outstanding thermal fatigue performance of the alloy of
the present invention evidences its exceptional resistance to high
temperature oxidation. This oxidation resistance is attributed to
chromium in the alloy composition that oxidizes preferentially and
forms a slowly growing, stable and protective oxide film on the
surface [23, 24]. The presence of such an oxide (Cr2O3) film on the
surface was confirmed with low angle x-ray diffraction analysis.
This stable and protective Cr2O3 layer served as a barrier to
excessive oxidation under thermal cycling test conditions. The much
lower tendency of the alloy of the present invention to oxidation
with respect to hot work tool steels is evidenced also by thermo
gravimetric analysis and by the metallographic analysis of sections
of samples submitted to thermal cycling. The surface oxide of the
alloy of the present invention is only 3-4 .mu.m thick after 16000
thermal cycles, much thinner than the oxide on the surface of hot
work tool steel samples, measured to be nearly 50 .mu.m in after
only 500 thermal cycles. Cr2O3 film on the surface is claimed to be
resistant to thermal stresses and retains its integrity.
[0023] Another feature that imparts to the alloy of the present
invention its high thermal fatigue resistance is its resistance to
loss of strength at elevated temperatures. Resistance to softening
is a key feature required in tool materials to ensure adequate
resistance to crack initiation as well as to crack growth. A hard
alloy protects its surface oxides and the surface oxides, in turn,
protect the underlying die steel. Surface oxide films offer this
protection by resisting against plastic deformations produced by
mechanical and/or thermal stresses, against blistering, expansions
and finally to cracking and spallation. Hence, surface
deterioration and damage are avoided and cracking is thus
delayed.
[0024] The resistance of the alloy of the present invention to loss
of strength at elevated temperatures has been confirmed with
hardness measurements across the section of the samples submitted
to thermal cycling. The softening in the present alloy is very
small and is confined to the surface during the first couple of
thousand cycles (FIG. 5). The average hardness measured to be 560
HV before thermal cycling, is still 500 HV after the first 5000
cycles. This hardness level implies a much higher softening
resistance of the present alloy with respect to hot work tool
steels, the hardness of which is measured to be as low as 250 HV
after the same number of thermal cycles [23]. Hardness of the alloy
of the present invention has continued to decrease steadily with
increasing number of thermal fatigue cycles. However, the hardness
of the alloy of the present invention even after 16000 thermal
cycles is 400 HV, and is thus higher than the hardness measured in
Stellite 6 alloy after 5000 cycles. It is fair to conclude in view
of the foregoing that the high resistance to softening of the alloy
of the present invention makes a favourable impact on its high
thermal fatigue resistance.
DESCRIPTION OF FIGURES
[0025] FIG. 1: Microstructure of the alloy of the present
invention: (a) dendritic matrix phase, (b) Molybdenum-rich
intermetallic phase.
[0026] FIG. 2: Change in temperatures measured on the front and
rear faces of the sample during a typical thermal cycle.
[0027] FIG. 3: Surface blistering detected on the front face of the
thermal fatigue test sample of the cobalt-based alloy of the
present invention after 13000 thermal cycles (a), (b, c)
development of this surface blister into a thermal fatigue crack
after 16000 cycles. (a and b) scanning electron and (c) optical
microscope micrographs.
[0028] FIG. 4: Section view of the thermal fatigue crack after
16000 thermal cycles.
[0029] FIG. 5: Change in hardness across the section of the thermal
fatigue sample with number of thermal cycles.
REFERENCES
[0030] [1] E. Doege, H. Nagele, U. Schliephake, J. Eng. Manuf. 208
(1994) 111. [0031] [2] K. Bobzin, E. Lugscheider, M. Maes, P.
Immich, Solid State Phenomena, 116-117 (2006) 704. [0032] [3] R.
Kopp, J. Kallweit, T. Moller, I. Seidl, J. Mater. Process Tech.
130-131 (2002) 562. [0033] [4] Y. Birol, Steel Res. Int., 80 (2009)
588. [0034] [5] Y. Birol. Ironmak & Steelmak, 36 (2009) 397.
[0035] [6] Y. Birol, Steel Res. Int. 80, (2009) 165. [0036] [7] S.
Muenstermann, K. Uibel, T. Tonnesen, R. Telle, Solid State
Phenomena, 116-117 (2006) 696. [0037] [8] E. Lugscheider, K.
Bobzin, C. Barimani, St. Barwulf, Th. Hornig, Adv. Eng. Mater. 2
(2000) 33. [0038] [9] M. Z. Omar, E. J Palmiere, A. A. Howe, H. V.
Atkinson, P. Kapranos, Mater. Sci. Eng. A395 (2005) 53. [0039] [10]
E. Lugscheider, Th. Hornig, D. Neuschutz, O. Kyrylov, R. Prange,
in: G. L. Chiamietta, M. Rosso (Eds.) Proc. 6th Int. Conf. On
Semi-Solid Processing of Alloys and Composites, Turin, 2000, pp.
587-592. [0040] [11] R. Telle, S. Muenstermann, C. Beyer, Solid
State Phenomena, Vol. 116-117 (2006) 690. [0041] [12] Y. Birol,
Ironmak & Steelmak, 36 (2009) 555. [0042] [13] Y. Birol, Steel
Research Int., 80 (2009) 165. [0043] [14] Y. Birol, Int. J. Mater.
Form. 3 (2010) 65. [0044] [15] Y. Birol, Ironmak. Steelmak. 37
(2010) 41. [0045] [16] J. C. Shin, J. M. Doh, J. K. Yoon, D. K.
Lee, J. S. Kim, Surf Coat. Tech. 166 (2003) 117. [0046] [17] I.
Radu. D. Y. Li, Wear 259 (2005) 453. [0047] [18] J. N. Aoh, Y. R.
Jeng, E. L. Chu, L. T. Wu, Wear 225-229 (1999) 1114. [0048] [19] H.
Kashani, A. Amadeh, H. M. Ghasemi, Wear 262 (2007) 800. [0049] [20]
M. X. Yao, J. B. C. Wu, Y. Xie, Mat. Sci. Eng. A 407 (2005) 234.
[0050] [21] C. Navas, A. Conde, M. Cadenas and J. de Damborenea,
Surf. Eng. 22 (2006) 26. [0051] [22] B. F. Levin, J. N. DuPont, A.
R. Marder, Wear 181-183 (1995) 810 [0052] [23] Y. Birol, Mater.
Sci. Eng. A527 (2010) 1938. [0053] [24] Y. Birol, Mater. Sci. Eng.
A 527 (2010) 6091. [0054] [25] Y. Birol, Wear, 269 (2010) 664.
[0055] [26] Y. Birol, A. Kayihan, Metal. Mater. Trans., 42 (2011)
3277
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