Heat-treatment of an alloy for a bearing component
09732408 · 2017-08-15
Assignee
Inventors
Cpc classification
F16C33/62
MECHANICAL ENGINEERING; LIGHTING; HEATING; WEAPONS; BLASTING
F16C2204/42
MECHANICAL ENGINEERING; LIGHTING; HEATING; WEAPONS; BLASTING
International classification
C22F1/18
CHEMISTRY; METALLURGY
Abstract
A method for preparing titanium alloy that is created to be formed into a bearing component, wherein the titanium alloy comprises from 5 to 7 wt % Al, from 3.5 to 4.5 wt % V, from 0.5 to 1.5 wt % Mo, from 2.5 to 4.5 wt % Fe, from 2.5 to 4.5 wt % Fe, and from 0.05 to 2 wt % Cr. The alloy can optionally include one or more of the following elements: up to 2.5 wt % Zr, up to 2.5 wt % Sn, and up to 0.5 wt % C. The balance of the composition comprises Ti together with unavoidable impurities. The alloy is heated to a temperature T below the (α+β/β)-transition temperature Tβ and then quenched. The alloy is then aged a temperature of from 400 to 600° C.
Claims
1. A method for manufacturing a bearing component selected from the group consisting of a rolling element, an inner ring, and an outer ring, the method comprising: (i) providing an alloy composition comprising: 5 to 7 wt % Al, 3.5 to 6 wt % V, 0.5 to 6 wt % Mo, 0.2 to 4.5 wt % Fe, 0.05 to 2.5 wt % Cr, up to 2.5 wt % Zr, up to 2.5 wt % Sn, and up to 0.5 wt % C; the balance being Ti and unavoidable impurities; (ii) heating the alloy composition to a temperature T below the (α+β/β)-transition temperature T.sub.αand then quenching; and (iii) aging the alloy composition at a temperature of from 400 to 600° C.
2. The method according to claim 1, wherein after the alloy composition has been heated to the temperature T, it is worked before being quenched.
3. The method according to claim 2 wherein the working comprises rolling the alloy composition.
4. The method according to claim 3 wherein the rolling comprises multiple rolling stages with intermediate annealing stages.
5. The method according to claim 1 wherein the temperature T is greater than the (α/α+β)-transition temperature T.sub.αand less than the (α+β/β)-transition temperature T.sub.β.
6. The method according to claim 1 wherein the temperature T falls within the range of:
T.sub.β>T≧T.sub.β−50° C.
7. The method according to claim 1 wherein the temperature T is 820 to 900° C.
8. The method according to claim 1 wherein the quenching is carried out in water.
9. The method according to claim 1 wherein the quenching is performed so that the alloy composition has a microstructure comprising from 10 to 15 vol % α-phase after the quenching.
10. The method according to claim 1 wherein the aging is carried out at a temperature of 425 to 525° C.
11. The method according to claim 10 wherein the aging is carried out 25 to 35 hours.
12. The method according to claim 1 wherein the alloy composition comprises: 5 to 7 wt % Al, 3.5 to 4.5 wt % V, 0.5 to 1.5 wt % Mo, 2.5 to 4.5 wt % Fe, 0.05 to 2 wt % Cr, up to 2.5 wt % Zr, up to 2.5 wt % Sn, and up to 0.5 wt % C, the balance being Ti and unavoidable impurities.
13. The method according to claim 1 wherein the alloy composition comprises: 5.5 to 6.5 wt % Al, 3.5 to 4.5 wt % V, 0.5 to 1.5 wt % Mo, 3.5 to 4.5 wt % Fe, 0.05 to 2 wt % Cr, 1.5 to 2.5 wt % Zr, 1.5 to 2.5 wt % Sn, and 0.01 to 0.2 wt % C, the balance being Ti together with unavoidable impurities.
14. The method according to claim 1 wherein the alloy composition comprises: 5 to 7 wt % Al, 3.5 to 6 wt % V, 3 to 6 wt % Mo, 0.2 to 2.5 wt % Fe, 0.1 to 2.3 wt % Cr, up to 0.7 wt % Zr, up to 0.7 wt % Sn, and up to 0.5 wt % C, the balance being Ti and unavoidable impurities.
15. The method according to claim 1 wherein the quenching is performed so that the alloy composition has a Rockwell harness of at least 50 HRC after the quenching.
16. The method according to claim 1 wherein the alloy composition has a microstructure that comprises β-phase having precipitates of α-phase dispersed therein, the α-phase comprising 10-20 vol % of the microstructure.
17. The method according to claim 1, further comprising: machining the alloy composition into a desired shape of the bearing component, the machining being carried out between the quenching and the aging.
18. The method according to claim 17 wherein after the aging, the machining is performed to remove a layer not less than 50 μm in depth from the bearing component.
19. The method according to claim 1 wherein the bearing component is an inner ring or an outer ring.
20. The method according to claim 1 wherein the temperature T is from 835 to 880° C.
21. The method according to claim 20 wherein the quenching is carried out at a rate of at least 20° C./s to a temperature of 60° C. or lower.
22. The method according to claim 21 wherein the aging is performed at 425 to 525° C. degrees C. for about 25-35 hours, and then the bearing component is cooled at a rate of from 2 to 10° C./s.
23. The method according to claim 1, wherein V is 3.5-4.5 wt %, Mo is 0.5-1.5 wt %, Fe is 2.5-4.5 wt % and Cr is 0.05-2 wt %.
24. The method according to claim 1 wherein the alloy composition has a molybdenum equivalence [Mo].sub.eq that is from 10 to 12, the molybdenum equivalence being calculated according to the following formula:
[Mo].sub.eq=[Mo]+0.2[Ta]+0.28[Nb]+0.4[W]+0.67[V]+1.25[Cr]+1.25[Ni]+1.7[Mn]+1.7[Co]+2.5[Fe].
25. The method according to claim 1, wherein Al is 6-6.5 wt %.
26. The method according to claim 1, wherein V is 3.5-4.5 wt %.
27. The method according to claim 1, wherein Fe is 2.5-4.5 wt %.
28. The method according to claim 1, wherein Cr is 0.06-1.5 wt %.
29. The method according to claim 1, wherein Zr is 1-2.5 wt %.
30. The method according to claim 1, wherein Sn is 1.5-2.5 wt %.
31. The method according to claim 1, wherein C is 0.015-0.35 wt %.
32. The method according to claim 1 wherein the temperature T falls within the range of:
T.sub.β>T≧T.sub.β−30° C.
33. The method according to claim 1 wherein the temperature T falls within the range of:
T.sub.β−10° C.≧T≧T.sub.β−20° C.
34. A method for manufacturing a bearing component selected from the group consisting of a rolling element, an inner ring, and an outer ring, the method comprising: (i) providing an alloy composition comprising: 5to 7 wt % Al, 3.5 to 6 wt % V, 0.5 to 1.5 wt % Mo, 0.2 to 4.5 wt % Fe, 0.05 to 2.5 wt % Cr, up to 2.5 wt % Zr, up to 2.5 wt % Sn, and up to 0.5 wt % C; the balance being Ti and unavoidable impurities; (ii) heating the alloy composition to a temperature T below the (α+β/β)-transition temperature T.sub.β and then quenching; and (iii) aging the alloy composition at a temperature of from 400 to 600° C.
Description
BRIEF DESCRIPTION OF THE FIGURES
(1) The invention will now be discussed further with reference to the figures, provided purely by way of example, in which:
(2)
(3)
(4)
(5)
(6)
(7)
(8)
(9)
(10)
(11)
(12)
(13)
EXAMPLES
(14) The invention will now be described further, by way of example, with reference to the following non-limiting examples.
Example 1
(15) Preparation of Ingots:
(16) A 20 kg ingot having the chemical composition Ti-6.4Al-4.1Fe-1.1Mo-4.3V-2.5Sn-2.4Zr was prepared in a vacuum arc furnace with consumable electrode by double remelting. Titanium sponge TG-120 was used to make the electrode. Due to its high melting point, Mo was introduced into the alloy via the ligature AMBTi (32% V-36.8% Mo-14% Al-0.39% Fe-0.23% Si), rather than in pure form which may lead to the formation of inclusions. V, Fe and Al were partly introduced via ligature and partly in pure form. Zr and Sn were introduced into the alloy in pure form only. Titanium sponge was mixed with ligature and pure elements, which were in the form of small metal pieces or turnings. The resultant mixture was pressed into a matrix using a 2000-ton press. The pressed electrode was melted in a copper crystallizer with diameter Ø125 mm and as a result the ingot of the first remelt was obtained. This ingot was then remelted for the second time in a crystallizer with diameter Ø150 mm at current intensity 2000 A, electric voltage 30V and vacuum level 5×10.sup.−2millimeters of mercury (6.7 Pa). Chemical analysis certified by TUV has shown that the chemical composition of the alloy corresponds to that of Ti-6.4Al-4.1Fe-1.1Mo-4.3V-2.5Sn-2.4Zr. After a second remelting the ingot was subjected to machining, which involved turning along the cylindrical surface and removal of shrinkage voids. The final ingot size after machining was Ø140 mm×250 mm.
(17) Preparation of Rods:
(18) In the next stage, the ingot was rolled on a screw-shaped rolling press at a temperature of about 1050° C. (corresponding to the β-range) from a diameter of 140 mm to a diameter of Ø55 mm in one gate, and then machined to a diameter of Ø47 mm for the removal of cinder and α-modified layers. The rod's microstructure after rolling in the β-range was homogeneous and presented by large initial β-grains of about 500-800 microns size. Dispersed within the β-grains were α-phase plates precipitated during cooling from rolling to room temperature (see
(19) Subsequent stages of rolling for obtaining rods with diameter 20 mm were conducted in the (α+β)-range at a temperature of about 850° C. in three operations with intermediate annealing at 850° C. for 10 minutes. Dynamic recrystallisation took place during rolling leading to a decrease of the initial β-grain size to about 50-70 microns. In addition, rolling in the (α+β)-range caused α-phase precipitation during the deformation process as well as during the subsequent cooling to room temperature (see
(20) Part of the rods with diameter Ø22 mm were then used for sample production (herein Rolf Samples). For the production of another set of samples (herein Polymet Samples), rods with a diameter of Ø20 mm were then rolled to Ø14 mm in one gate on a screw-shaped rolling mill also at a temperature of about 850° C. Additional deformation and dynamic recrystallisation processes lead to a reduction of the initial β-grain size to about 28-45 microns (see
(21) Heat Treatment:
(22) The starting semiproducts, especially the rod with diameter Ø14 mm, exhibited high hardness values. Accordingly, an attempt was made to obtain the desired hardness level (preferably ≧52 HRC) by ageing the hot-rolled semiproducts without intermediate quenching. However a significant increase in hardness was not observed. The hardness of the Ø14 mm rod increased on average by 1 unit, while that of the Ø20 mm rod increased on average by about 5-6 HRC units. However, the increase in hardness was not sufficient.
(23) Accordingly, it is clear that in order for high levels of hardness to be obtained, a quenching step is desirable before ageing. Previous investigations conducted on trial samples have shown that the quenching temperature should correspond to an upper interval of the (α+β)-range. However, the higher the quenching temperature in the (α+β)-range, the less alloyed the β-phase is and the higher the strengthening increase that can be obtained by the subsequent ageing. In spite of the fact that minimal degree of β-phase alloying is achieved on quenching from the β-range, such a high temperature causes the complete removal of deformation strengthening, which is also a very important component of the final strengthening. Therefore, quenching from the β-range and subsequent ageing led to a smaller strengthening effect in comparison to a heat treatment consisting of quenching from the (α+β)-range and ageing.
(24) The polymorphic transformation temperature Tβ, the temperature of the (α+β)/β-transition, was determined for the new alloy using a trial quenching method. Samples were quenched in water in temperature intervals of about 20° C. between about 800-960° C. With the use of metallographic and X-ray analysis it was shown that for the alloy investigated the polymorphic transformation temperature Tβ was equal to about 880° C. Therefore, the quenching temperature was chosen to be about 850° C. After quenching, samples cut from rods Ø14 mm and 20 mm have practically identical microstructures consisting of small amounts of α-phase and β-phase (see
(25) In samples with diameter Ø14 mm, the primary α-phase has a predominantly globular shape (see
(26) Next, ageing of samples was carried out at temperatures of about 500° C. and about 530° C. for various lengths of time. The results are set out in Table 1.
(27) TABLE-US-00001 TABLE 1 Hardness of samples of Ti—6.4Al—4.1Fe—1.1Mo—4.3V—2.5Sn—2.4Zr alloy after ageing at 530° C. and 500° C. for various lengths of time Hardness after Diameter quenching Ageing Duration, hour (mm) (HRC) T (° C.) 1 2 6 8 10 14 18 20 20 36 500 50-50.5 49-50 51-51.5 51 51 51.5 51 51-51.5 530 51 51 50-51 51-51.5 51-51.5 51 50.5-51 50.5-51 14 37 500 51.5-52 51-52 52-52.5 52 52-53 53 52.5-53 52.5-53.5 530 51.5-52 51-51.5 51-52 51 51-51.5 51 51 51
(28) The ageing temperature determines the diffusive mobility of the alloying element atoms, as well as the rate of nucleation and growth of secondary α-phase particles. The higher the ageing temperature the higher the diffusive mobility of the atoms and, therefore, the more rapidly the precipitation process begins and finishes. At the same time, at higher temperatures, growth rate predominates over nucleation rate and therefore the precipitated α-phase particles will be larger. Consequently, the level of strengthening will be lower.
(29) For rods with a diameter of Ø20 mm the precipitation process at about 530° C. is complete after about 8-10 hours (see Table 1), but the required hardness is not achieved. When ageing at about 500° C. the precipitation process is complete after 18 hours. It is noted that the lowest required hardness level is obtained on the Ø14 mm rod.
(30) In order to achieve a hardness higher than about 52 HRC, the ageing temperature was decreased to about 475° C. While this required an ageing time of about 25-27 hours, a hardness of about 53-54 HRC was obtained (see Table 2).
(31) As shown in
(32) TABLE-US-00002 TABLE 2 Hardness of samples of Ti—6.4Al—4.1Fe—1.1Mo—4.3V—2.5Sn—2.4Zr alloy after ageing at 475° C. for various lengths of time Hardness after Diameter quenching Duration, hour (mm) (HRC) 1 3 6 10 14 20 25 30 40 20 36 50.5-51 51-51.5 51-52 50-51 52-53 52.5-53 53-53.5 53 53-53.5 14 37 51 52-52.5 52-53 53-53.5 53-53.5 53.5 53.5-54 53.5-54 53.5
(33) Preparation of Samples:
(34) Intermediates were cut from the 20 and 14 mm rods, and were then heated to a temperature of about 850° C. in an air furnace and then quenched. Once quenched, the samples had a low hardness and, therefore, it was easer to machine them. Accordingly, after quenching, samples were machined to remove of 0.30 mm per side. The machined semiproducts were subjected to ageing. In order to check the level of harness during the ageing a control sample was put with each sample batch. The control samples had a hardness of about 53-54 HRC. The samples were then subjected to final machining.
(35) Testing of Resistance to Rolling Contact Fatigue (RCF):
(36) Twenty sets of inner and outer rings of bearings were produced and heat treated. All rings exhibited a hardness value of about 52-53 HRC, with a narrow standard deviation. The dimensions of the bearings were as follows: Bearing name: SKF 6309 bearing 6309 size Brand: SKF Outer diameter D (mm): 100 Diameter d (mm): 45 Thickness B (mm): 25
(37) No coatings were applied to the rings. The rings were assembled into bearings with a polymer L-shaped cage and ceramic balls. The rolling contact fatigue characteristics of the bearings were investigated under the following conditions: Test rig: TLE G19.08 (R2) Load measuring device: TLE G19.20 Bearing axis: horizontal Lubrication: Shell Spirax 80W90 oil bath at ambient temperature Radial load: 700 daN (C/P=8) on fixed outer ring Inner ring speed: 3450 rpm (n.dm=250.000) Test duration: 1.5 million rev Contact stress: 2.4 GPa
(38) The test was able to be carried out for 123.6 hours and 25585200 revolutions without failure. Accordingly, it is clear that the rings exhibited the required mechanical stability (and surface stability since the rings were not coated) to be able to hold the stresses without plastic collapse. No known Ti alloys are capable of surviving such a stress level. The testing load is representative of loads used for testing high strength bearing steels. For safety, many industrial applications, such as windmills, require stress levels to be maintained below 2 GPa. Although the test cannot be considered as a life test, it gives a good indication of the alloy characteristics achieved by the combination of the heat treatment and alloy composition.
(39) Results of further tests of the rolling contact fatigue characteristics of the rings are set out in Table 3 below.
(40) TABLE-US-00003 TABLE 3 Results of rolling contact fatigue tests on samples of alloy Ti—6.4A1—4.1Fe—1.1Mo—4.3V—0.07Cr—2.5Sn—2.4Zr—0.02C Test Track dimensions Temperature Stress Stress Width Depth Rider (° C.) (GPa) cycles (mm) (gm) Comments Hardened Steel Discs 20 2.5 5 × 10.sup.7 0.65 Not measured Suspended AISI 52100 at 62 2.8 5 × 10.sup.7 0.70 Not measured Suspended HRc 3.2 5 × 10.sup.7 0.78 Not measured Suspended (Polymet Samples) 4.0 5 × 10.sup.7 0.82 Not measured Suspended 4.5 1.8 × 10.sup.7 0.91 ≦0.2 Spalled 5.0 4.5 × 10.sup.7 0.98 .sup. 2.5 Spalled Si.sub.3N.sub.4 balls 75 1.5 1.3 × 10.sup.8 Not measured Suspended (Rolf Samples) 2.1 1.3 × 10.sup.8 Not measured Suspended 3.0 1.3 × 10.sup.8 Not measured Suspended 4.0 1.3 × 10.sup.8 Not measured Suspended 4.5 5 × 10.sup.7 ≦0.15 Spalled
(41) Any “suspended” test is a successful one, i.e. it is suspended without failure or spalling. The Polymet Samples feature a longer bar than the Rolf Samples. Also, instead of being in contact with ceramic balls (Rolf Samples), the Polymet Samples are in contact with two rotating discs of hardened steel (AISI 52100 at 62 HRc). As a consequence, the Rolf Samples are characterised by a point contact, whereas the Polymet samples are characterised by a line contact. In every turn, the Rolf Samples make 3 loading cycles, whereas the Polymet Samples make 2.
Example 2
(42) 200 g Ingots having the following compositions were smelted in a laboratory vacuum arc furnace with a non-consumable tungsten electrode by fivefold remelting: Ti-6Al-4V-1Mo-1Cr-3.5Fe Ti-6Al-2V-0.5Mo-2.5Cr-3.5Fe Ti-6Al-2V-0.5Mo-1Cr-5.5Fe Ti-6Al-4V-1Mo-1Cr-3.5Fe-2Sn-2Zr Ti-6Al-4V-1Mo-1Cr-3.5Fe 2Sn-2Zr-0.15C Ti-6Al-4V-1Mo-1Cr-3.5Fe2Sn-2Zr-0.3C Ti-6Al-4V-1Mo-1Cr-3.5Fe2Sn-2Zr-0.5C
(43) The resulting ingots had a height of approximately 14 mm and a diameter of approximately 40 mm. Visual inspection of the ingots and analysis of their microstructures allowed selection of three alloy compositions for further investigation. Structures of alloys containing more than 4.5% β-eutectoid stabilizers had a large microchemical heterogeneity, which was difficult to eliminate by homogenization annealing. This phenomenon is very widespread for alloys containing a large quantity of β-eutectoid stabilisers (therefore their typical content does not exceed 4.4-5% in β-titanium alloys). It was not possible to eliminate microchemical heterogeneity under fivefold remelting. For alloys containing 0.3% and 0.5 wt % C, cracks were observed appearing on the ingots after their cooling to room temperature. This is presumably connected with the emergence of greater tensile stress due to formation of α-interstitial solid solution and carbides, since carbon has low solubility in the β-phase, the volume fraction of which is more than 80% after cooling.
(44) Therefore, three compositions of Ti alloys were chosen for further investigation: Alloy 1: Ti-6Al-4V-1Mo-1Cr-3.5Fe (basic alloy) Alloy 2: Ti-6Al-4V-1Mo-1Cr-3.5Fe-2Sn-2Zr Alloy 3: Ti-6Al-4V-1Mo-1Cr-3.5Fe-2Sn-2Zr-0.15C
(45) The cast structure of these alloys is uniform and characterized by large β-grains and fine particles of α-phase precipitating from the β-phase during cooling of the ingots to room temperature (
(46) The amount of α-phase precipitation depends on the chemical composition of an alloy. A rather large amount of α-phase is precipitated in the Ti-6Al-4V-1Mo-1Cr-3.5Fe alloy (not shown). The addition of Sn and Zr (Alloy 2) results in stabilisation of the β-phase and reduction of the critical cooling rate (rate under cooling with which the diffusion of component atoms is suppressed). Accordingly, there is a negligible amount of separate α-phase particles in the structure (
(47) The difference in hardness of cast alloys mainly depends on the amount of precipitated α-phase and the degree of its dispersion. Results of hardness measurements are given in Table 4.
(48) TABLE-US-00004 TABLE 4 Hardness of pilot titanium cast alloy Alloy chemical composition HRC Alloy 1 Ti—6A1—4V—1Mo—1Cr—3.5Fe 48 Alloy 2 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr 35.5 Alloy 3 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr—0.15C 41.5
(49) Alloy 2 has a minimal hardness because in this case the hardness is caused by solid solution strengthening of the β-phase. For Alloys 1 and 3, the higher values of hardness are caused by precipitated α-phase particles. In Alloy 1, smaller particles give higher hardness. Ingots of the three chosen compositions were divided in half and were then forged at about 900° C. On average the initial blanks were approximately 20×14 mm in cross-section. They were then deformed to obtain a square cross section. The final size of samples with approximate section 12×11 mm and length 100 mm were obtained by forging in the longitudinal direction. The reduction ratio was about 2, which is determined as a ratio of the cross sectional area before and after deformation. This indicates than during forging the materials was deformed to only a small degree.
(50) The obtained 100 mm billets were cut up into samples with heights of 1 to 15 mm for conducting structure studies in the initial state and after extra thermal treatment, and also hardness measurements.
(51) Microstructures of Alloys 1, 2 and 3 after forging are represented in
(52) Since the material was subjected to small deformation during the forging process and since the temperature was rather high, relaxation processes had time to run. Therefore there is a hardness decrease in the Alloy 1 after forging (see Table 2). For the Alloy 2, there is a hardness increase, due to precipitation of disperse α-phase particles. There are almost no changes of hardness in the Alloy 3.
(53) TABLE-US-00005 TABLE 5 Hardness of pilot titanium alloys after forging at 900° C. Alloy chemical composition HRC Alloy 1 Ti—6A1—4V—1Mo—1Cr—3.5Fe 40.0 Alloy 2 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr 40.0 Alloy 3 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr—0.15C 42.0
(54) The temperatures of the α+β/β transformation (Tβ) for each of Alloys 1, 2 and 3 were determined using test quenches. For Alloy 1, it was about 940° C., for Alloy 2 it was about 900° C., and for Alloy 3 it was about 1000° C. A part of each sample was quenched from the β-area, and a part was quenched from the (α+β)-area from temperatures about 50° below Tβ, namely about 890°, 850° and 950° C. for Alloys 1, 2 and 3, respectively.
(55) All alloys quenched from the β-area had the same structure consisting of β-grains (
(56) The hardness of alloys after quenching depends on the heating temperature. After quenching from the β-area, hardness is minimal and it is determined only by the alloying level of the solid solution, which is increased from Alloy 1 to Alloy 3 (see Table 6). After quenching from the (α+β)-area, hardness is also determined by the degree of dispersion of primary α-phase particles.
(57) TABLE-US-00006 TABLE 6 Hardness of pilot titanium alloys after quenching in β- and (α + β)-areas HRC Alloy chemical composition of β-area of (α + β)-area Alloy 1 Ti—6A1—4V—1Mo—1Cr—3.5Fe 29.0 32.5 Alloy 2 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr 34.0 38.0 Alloy 3 Ti—6A1—4V—1Mo—1Cr—3.5Fe—2Sn—2Zr—0.15C 35.0 36.5
(58) Since the cooling rate under quenching is faster than the cooling rate after forging, the alloying level of β-phase will be different: the higher the heating temperature and cooling rate, the lower the alloying level of β-phase will be and all of the conditions being equal, the effect of strengthening should be greater.
(59) The alloying level of β-phase could be estimated by the change of its lattice parameter calculated from X-ray analysis data. Values of aβ are given in Table 7.
(60) TABLE-US-00007 TABLE 7 Change of lattice spacing β-phase (a.sub.β) in different states a.sub.β, nm State Alloy 1 Alloy 2 Alloy 3 After forging 0.3221 0.3232 0.3236 After quenching from β-area 0.3238 0.3240 0.3250 After quenching from (α + β)-area 0.3235 0.3237 0.3247
(61) It can be seen from the data presented in the Table 7 that a minimum alloying level of β-phase corresponds to quenching from the β-area, and maximum alloying levels correspond to the state after forging. Based on the alloying level of β-phase for the same alloy, one can indirectly estimate the amount of α-phase in the structure. The smaller the aβ lattice parameter, the more α-phase in the structure. In other words, there is more α-phase in the forging state in the structure, which is a reason for the higher values of hardness (see Tables 5 and 6). Of course, higher hardness after forging is also caused by deformation strengthening. But as deformation during forging is not large, it is precipitation of α-phase during deformation and cooling that is thought to make the main contribution to strengthening.
(62) Next, samples from alloys 1, 2 and 3 were aged at a temperature of about 500° C. for different lengths of time in three states: after forging, quenching from β- and (α+β)-areas. The results of are given in the Table 8. As can be see from the data in Table 8, quenching from the β-area results in less strengthening. However, ageing after quenching from the (α+β)-area and immediately after forging provide similar results. In all cases, maximum strengthening was obtained after isothermal exposure for about 6 hours.
(63) Samples of Alloys 1 and 2 were quenched from the (α+β)-area and aged at about 500° C. for about 6 hours. The structure of these samples is represented in
(64) TABLE-US-00008 TABLE 8 Hardness of pilot titanium alloys after ageing at about 500° C. for different periods of time HRC State Alloy 1 h 1.5 h 2 h 3 h 4 h 5 h 6 h 7 h 8 h 9 h 10 h After Forging 1 49.5 50 50 51 51 52 52 52 50 50 51 2 51 51 51 52 51 52 53 52.5 52.5 52 52 3 51.5 51.5 52 52 52.2 52.5 53.7 52.5 52 52 52 After 1 50 50.5 50.5 51 51 51 51 51 50 50 49 quenching 2 50 50 50.5 51 51 51.5 51.5 51 51 51 51 from β-area 3 50.5 50.5 50.5 51 51 51.5 51.5 51.5 51 50.5 51 After 1 49.5 50 50 51 51.5 52.5 52.5 52.5 52 52 52 quenching 2 50 50 51 51.5 52 52.5 53 53 52.5 52.5 52 from α + β)- 3 51 51 52 52 52.5 53 53.5 53.5 53 52 52 area
Example 3
(65) Forged semiproducts were prepared having the composition set out in Table 9 below. The Forged semiproducts were cut into billets of approximate 120×12×40 mm size. A number of billets were cut into small samples of approximate 15×15×20 mm size for further investigation.
(66) TABLE-US-00009 TABLE 9 Chemical composition of semiproducts Element Al V Mo Fe Cr Zr Sn C N.sub.2 O.sub.2 wt % 6.5 3.8 1.1 3.0 0.8 1.9 2.2 0.02 0.03 0.14
(67) The microstructure of the semiproduct after hot rolling was investigated (see
(68) The hot-rolled semiproduct had a relatively high hardness of about 43-44 HRC. Following low-temperature ageing the hardness only increased up to about 47 HRC (see
(69) In order to select the optimal quenching temperature, it was necessary to determine the β-transition temperature (Tβ). A number of trial quenches were carried out. Samples were heated up to about 800 to 900° C. in steps of 5 to 10° C. and subjected to isothermal annealing for from about 40 minutes to about 2 hours depending on the temperature. After isothermal annealing the samples were cooled in water. The temperature of the α+β.fwdarw.β-transformation (Tβ) was determined by metallographic and X-ray analysis, and was found to be about 860° C. Quenching from a temperature equal to or higher than about 860° C. leads to the formation of β-phase microstructure (see
(70) Due to its chemical composition, the Ti-6Al-4V-1Mo-3.5 Fe-1Cr-2Zr-2Sn alloy belongs to the pseudo-β group of titanium alloys. This group is characterized by micro-chemical heterogeneity inside β-phase grains, i.e. there is some difference in chemical composition in adjoining β-grains. This is why reducing the temperature down to about 855° C. results in heterogeneous microstructure formation: some grains are presented by β-phase only, while others also contain primary α-phase (see
(71) TABLE-US-00010 TABLE 10 Influence of quenching temperature on Ti—6A1—4V—1Mo—3.5Fe—1Cr—2Zr—2Sn alloy hardness Quench- ing Temp (° C.) 880 860 855 845 830 800 Phase β β β + 5% α β + β + >20% α β + compo- (hetero- 10-15% α >20% α sition geneous micro- structure) Hard- 32-33 32-33 32-36 37-38 39-40 43-44 ness (HRC)
(72) A single-phase β-microstructure after quenching provides only minimal levels of hardness (about 32 HRC). Heterogeneous microstructure formation (quenching from about 855° C.) causes the hardness to fluctuate between about 32 and about 36 HRC. Reducing the quenching temperature leads to a gradual increase in hardness, which is explained by an increase in the amount of primary α-phase (see Table 10).
(73) In order to provide preferred hardness values of about 52 HRC or more, optimal temperature and ageing times were investigated. Three quenching temperatures (about 860° C., 845° C. and 830° C.) and ageing temperatures (about 475° C., 500° and 525° C.) were chosen. Ageing time was varied from about 4 up to about 70 hours. The results are given in Table 11.
(74) Maximum strengthening, i.e. the greatest difference between initial and final hardness levels, was achieved on samples which had single-phase β-microstructure after quenching. However, the level of hardness obtained did not exceed about 49 HRC. In addition, ageing for more than about 120 hours at the chosen temperatures results in abrupt material brittleness: hardness indentation causes the appearance of small cracks on the surface. It appears that the α-phase, which is formed during low-temperature ageing, has a semicoherent boundary with the matrix, which results in high internal stresses.
(75) The results indicate that ageing for about 70-hours at a temperature of about 475° C. is generally insufficient to achieve the required hardness level. It is likely that ageing for up to 100-150 hours would provide the necessary levels, but this is not reasonable from an economic perspective.
(76) TABLE-US-00011 TABLE 11 Influence of ageing temperature and its duration on Ti—6A1—4V—1Mo—3.5 Fe—1Cr—2Zr—2Sn alloy hardness Initial Quenching hardness Ageing HRC Temp (° C.) (HRC) Temp (° C.) 4 h 8 h 10 h 20 h 30 h 50 h 70 h 860 33 525 48.0 48.5 48.5 48.0 — — — 500 — — 48 49.0 — — — 475 — — 47.0 48.0 — — — 525 47.5 47.5 48.0 48.5 48.5 48.0 47.5 845 37 500 — — 50.5 50.5 52 51.5 51.0 475 — — 49.5 49.5 50.5 50.7 51.0 830 39 525 47.0 47.0 46.5 47.5 46.5 46.0 45.0 500 — — 49.0 49.0 49.0 48.5 48.0 475 49.0 49.0 49.5 49.5 49.5
(77) Ageing at about 525° C. also does not provide the optimum result due to an increase of diffusion processes. As a result, α-phase particles have a larger size in comparison with ones that are formed at lower temperatures.
(78) The results indicate that the optimum ageing temperature is approximately 500° C. Ageing at this temperature for up to about 30 hours resulted in hardness values of about 52 HRC for the samples that were previously quenched from about 845° C. (see Table 11). It should be noted that α-phase particle precipitation during ageing is highly dispersive: even on images, obtained with the help of scanning electron microscopy with magnification up to 8000, these particles cannot be clearly seen. This is also caused by a semicoherent interphase α/β-boundary.
(79) In summary, the results indicate that particularly high hardness levels can be achieved by quenching from a temperature that is about 15-20° C. lower than Tβ. Such a quench typically provides about 10-15% volume fracture of primary α-phase in the microstructure. Quenching from temperatures that are lower than Tβ allows partial retention of crystalline defects (for example dislocations) accumulated during plastic deformation and in this way contributes to strengthening of the alloy after subsequent ageing. In addition, the presence of a small quantity of α-phase particles allows partial retention of the toughness and the prevention of spontaneous crack formation.
(80) Increasing the quantity of α-phase, caused by reducing the quenching temperature, results in reduction of the strengthening effect due to an increase of β-phase stability and a decrease in the quantity of α-phase, which is formed during ageing. As a result, for a titanium alloy that contains 6.5Al-3.0Fe-0.8Cr-1.1Mo-3.8V-19Zr-2.Sn, high strength can be obtained by quenching from 845° C.±2° C. followed by ageing at about 500° C. for about 30 hours.
(81) Since titanium alloys exhibit low thermal conductivity, the hardness at various depths was investigated. For that purpose hardness was measured on a central part of a semiproduct previously quenched in water. The cutting plan and obtained hardness values are shown in
(82) Semiproducts approximately 120×120×40 mm in size were placed into a furnace and heated to a temperature of about 845° C. (in approximately 37-40 minutes). The semiproducts were kept in the furnace for another hour and were then quenched in water. In the next stage the semiproducts were machined on a lathe with a programmed numerical control. Rough turning was implemented by a carbide tool (plate) named VK8 (analogue to HT10); speed of rotation n=100 rpm (revolutions per minute) and feed f=0.1 mm/rev. Finish turning of the rings was implemented at n=120 rpm, f=0.05 mm/rev by carbide tool (plate) HT 10 (Mitsubishi) using lubricant-coolant liquid. In addition two semiproducts 120×120×40 mm in size were used to make samples for tensile and impact tests. Some samples were tested after quenching, while others were additionally aged at about 500° C. for about 30 hours. Results of the mechanical tests are presented in Table 12. The mechanical properties were measured according to a uniaxial tensile test (ASTM E8/E8M−11).
(83) TABLE-US-00012 TABLE 12 Mechanical properties of titanium alloy Ti6A1—4V—1Mo—3.5 Fe—1Cr—2Zr—2Sn (yield strength, tensile strength, elongation and impact strength) KCU Conditions of thermal treatment σ.sub.0.2 (MPa) σ.sub.B (MPa) δ (%) (kJ/m.sup.2) 845° C., 1 hour, water cooling 920 1060 4.2 130.0 845° C., 1 hour, water cooling + 1400 1400 — 42.0 500° C., 30 hours, air cooling
(84) Samples of Ti6Al-4V-1Mo-3.5 Fe-1Cr-2Zr-2Sn alloy after quenching have average levels of strength and ductility, whereas after aging they have high strength and practically zero ductility. The alloys were compared with steels used for bearings manufacturing (Russian alloy Fe-1.0C-0.25Si-0.3Mn-1.45Cr). After quenching and annealing, steel has higher hardness (61-64 HRC) and strength (2100 MPa) level, but the impact toughness of bearing steel (50 kJ/m2) is rather close to that of the titanium alloy of the present invention (42 kJ/m2).
(85) Fracture surfaces of destroyed samples were studied. Large facets caused by intergranular failure were observed on the fracture surface of quenched samples. However, it should be noted that all facets have tracks of micro plastic deformation, which can be concluded from their pit structure. Secondary cracks formation as well as flat facets with small pits could also be observed on the fracture surfaces.
(86) The fracture surface of quenched and subsequently aged samples has a brittle character of failure. The fracture surface is also characterised by facets having feebly marked relief. Secondary cracks were almost not observed. Some facets have “river” features, while others have dispersive pit destruction due to precipitation of dispersive α-phase particles inside the grains.
(87) Bearing rings are typically manufactured from quenched material when its hardness is not too high. To achieve the required hardness level (preferably≧52 HRC) it is desirable to conduct ageing at about 500° C. for up to about 30 hours.
(88) Titanium alloys have high oxidation susceptibility. In addition, increasing the volume fraction of β-phase (at the same temperature and duration) results in intensification of the oxidation processes. Ti-6Al-4V-1Mo-3.5Fe-1Cr-2Zr-2Sn is a pseudo-β titanium alloy and has more than 50% β-phase after annealing. Accordingly, the depth of the “alpha”-layer (i.e. layer enriched with oxygen and having higher brittleness) was measured on oblique sections (the load on the indenter was P=50 g) on the Ti6Al-4V-1Mo-35 Fe-1Cr-2Zr-2Sn alloy samples after quenching and ageing at about 500° C. for about 30 hours. This layer was revealed to be about 50 microns thick, as indicated by a drop in harness of from around 6000 MPa at the surface to around 4000 MPa at a depth of 50 microns, after which the hardness remained fairly constant down to 200 microns in depth. Accordingly, final machining of the rings after ageing should preferably remove a layer not less than about 50 microns in depth.
Example 4
(89) Two hot-rolled rods of approximate diameter 30 mm and 25 mm were prepared having the chemical composition set out in Table 13.
(90) TABLE-US-00013 TABLE 13 Chemical composition of hot-rolled rods Rod Alloying elements, wt % diameter Al Mo V Cr Fe Base 30 mm 5.4 4.6 5.9 1.1 1.0 Ti 25 mm 5.5 4.6 4.9 1.1 1.2 Ti
(91) Samples were quenched from either about 870° C. (just below Tβ) or about 900° C. (just above Tβ) at a rate of about 30° C./s. They were then aged for varying lengths of time at either about 450° C., 500° C. or 550° C. before being cooled at a rate of about 3° C./s. Hardness values of the samples are set out in Table 14.
(92) TABLE-US-00014 TABLE 14 Hardness values achieved for various quenching temperatures, ageing temperatures and ageing times HRC after Quenching quench- Hardness, HRC T (° C.) ing (Time of ageing, hours) Ageing at 450° C. (2) (4) (5) (6) (10) 900 (25 mm) 32 47 48 48.5 49.5 48.5 870 (25 mm) 35 49 49 50 50 50 900 (30 mm) 30 46 46 47 47 47 870 (30 mm) 33 47 47 48 48.5 48.5 Ageing at 500° C. (2) (3) (4) (5) (6) 900 (25 mm) 32 47 47 48 48 46 870 (25 mm) 35 48 49 49 48.5 47 900 (30 mm) 30 46 46.5 46.5 46.5 45 870 (30 mm) 33 47 47 47.5 47.5 47 Ageing at 550° C. (1) (2) (3) (4) 900 (25 mm) 32 48 48.5 47 45 870 (25 mm) 35 49 49 48.5 47.5 900 (30 mm) 30 45 46 46 44 870 (30 mm) 33 47 47.5 47.5 46
(93) It can be seen that the ageing treatments result in an increase in hardness.
(94) The foregoing detailed description has been provided by way of explanation and illustration and is not intended to limit the scope of the appended claims. Many variations in the presently preferred embodiments illustrated herein will be apparent to one of ordinary skill in the art, and remain within the scope of the appended claims and their equivalents.