Effect of inclusion size on the high cycle fatigue strength and failure mode of a high V alloyed powder metallurgy tool steel
Jun Yao, Xuan-hui Qu, Xin-bo He, and Lin Zhang
State Key Laboratory for Advanced Metals and Materials, School of Materials Science and Engineering, University of Science and Technology Beijing, Beijing 100083, China (Received: 3 July 2011; revised: 8 October 2011; accepted: 10 October 2011)
Abstract: The fatigue strength of a high V alloyed powder metallurgy tool steel with two different inclusion size levels, tempered at different temperatures, was investigated by a series of high cycle fatigue tests. It was shown that brittle inclusions with large sizes above 30 μm prompted the occurrence of subsurface crack initiation and the reduction in fatigue strength. The fracture toughness and the stress amplitude both exerted a significant influence on the fish-eye size. A larger fish-eye area would form in the sample with a higher fracture toughness subjected to a lower stress amplitude. The stress intensity factor of the inclusion was found to lie above a typical value of the threshold stress intensity factor of 4 MPa·m1/2. The fracture toughness of the sample with a hardness above HRC 56 could be estimated by the mean value of the stress intensity factor of the fish-eye. According to fractographic evaluation, the critical inclusion size can be calculated by linear fracture mechanics.
Keywords: powder metallurgy; tool steel; fatigue of materials; strength of materials; failure modes; inclusions; fractography
[This work was financially supported by the National Key Technologies Research and Development Program of China (No. 2007BAE51B05).]
1. Introduction
Powder metallurgy (PM) tool steels are superior to con-ventional ingot/casting tool steels because of their fine and homogeneous microstructure, good wear resistance, and excellent strength. Now they are widely used as cold- working applications and as engineering parts and compo-nents in many industries. However, these PM tool steels are always subjected to repeated stresses during contact be-tween a tool and its work material. This may result in the fa-tigue fracture of the tool material. In the case of PM materi-als, a fatigue crack is often initiated from an inclusion, and occasionally from a carbide or a pore [1-5]. Inclusions are difficult to completely suppress and are present in all com-mercial materials as a result of deoxidizing additions, impu-rities, or entrained exogenous materials. Therefore, it is very necessary to study the influence of inclusions on the high cycle fatigue response of PM tool steels.
The fatigue failure induced from a defect usually forms a so-called fish-eye. When a fish-eye comes into being, the high cycle fatigue strength of the steel is significantly af-fected by the size of the crack origin [2-3, 6-10]. Some works further found that reducing the inclusion size can im-prove the fatigue properties [2, 6-8, 11-13]. Thus, there is much more attention being paid to the critical inclusion size in tool steels, especially in PM ones. Moreover, cracks tend to nucleate in the vicinity of the crack origin and often cause internal crack initiation during 104-107 cycles. Another fa-tigue failure mode of subsurface crack initiation was re-vealed [5, 14-18]. Its crack origin always locates near the specimen subsurface, and the fish-eye often exhibits a semi-circle. However, there is less research about the influ-ence of inclusion size on the fatigue failure modes of PM tool steels compared with conventional tool steels and high strength steels in this respect.
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Therefore, the aim of this work is to study the effect of inclusion size on the high cycle fatigue strength and the fa-tigue failure mode of a high V-alloyed PM cold-working tool steel (AISI 11). The mechanical properties were ob-tained from axial loading fatigue tests, tensile tests, and fracture toughness measurements at room temperature. With the assistance of fractographic investigation, the quantitative estimation of the fatigue properties and the critical inclusion size was performed by linear fracture mechanics.
2. Experimental
The present study was performed on two materials which were produced by the hot isostatic pressing method. The nominal chemical composition is illustrated in Table 1. Steel A was produced by Crucible Material Corporation in USA (AISI 11), and steel B was under the same designation but without inclusion size control. The samples subjected to mechanical tests were wire-cut from the forged-and-soft annealed master blocks. These samples were then austeni-tized at 1120°C for 15 min in vacuum and gas-quenched in uniform N2. Tempering was done triply at 540-620°C for 2 h. As-tempered specimens were finally well prepared to ob-tain the mirror-like finish in the longitudinal direction. The microstructure of the two materials was studied on a Zeiss SUPRA-55 scanning electron microscope (SEM) and a JEM-2100 transmission electron microscopy (TEM).
Table 1. Chemical composition of studied materials wt%
Samples C V Cr Mo O
Steel A 2.45 9.40 5.44 1.24 0.011
Steel B 2.46 9.76 5.20 1.26 0.016
The dimension of specimens used for tensile, fracture toughness, and axial fatigue tests is shown in Fig. 1. Tensile (Fig. 1(a)) and fracture toughness (Fig. 1(b), compact ten-sion (CT) samples with a width of 20 mm) tests were per-formed on a servohydraulic testing machine. CT samples were first precracked by cyclic tensile stress. Axial loading fatigue experiments (Fig. 1(c)) were carried out on a reso-nant testing machine with a working frequency of about 140 Hz and a load ratio of 0.1. Each specimen was tested to its final failure or to the maximum life of 107 cycles. According to the stair-case method, at least 8 specimens were used to obtain the value of fatigue strength. In addition, a series of high cycle fatigue tests at maximum stresses of 1650 MPa and 1050 MPa was carried out, respectively. Fractography was favored to investigate the fatigue nucleation site and the fish-eye area on the fracture surface, the size of which were
represented by an equivalent diameter of a circular-like de-fect.
X-ray diffraction (XRD) techniques confirmed that the two experimental materials are both composed of uniformly distributed carbides, tempered martensite, and no indication of significant amounts of retained austenite. Fig. 2 shows the detailed microstructure of steel A and steel B. As shown in Fig. 2(a) and Fig. 2(b), most of the primary carbides are 1-3 μm, which are identified to be of the VC type, and the finely dispersed secondary carbides are even much smaller. The TEM micrograph and electro-diffraction pattern presented in Figs. 2(c)-2(f) demonstrate that the tempered matrix of AISI 11 is twin martensite. An earlier study found that the micro-structure of the two materials, obtained from the same heat treatment, is difficult to distinguish from each other [19]. The two materials both have the same mean grain size and volume fraction of carbides.
The experimentally determined mechanical properties are presented in Table 2. It is noted that owing to the special microstructure (Fig. 2), AISI 11 is a high-strength steel with low ductility. However, when the two materials are tem-pered at a similar hardness, the tensile strength of steel A is much higher than that of steel B. The reason for such a dif-
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Fig. 2. Microstructures of tempered steels A1 and B1: (a) SEM image of the microstructure in sample A1; (b) SEM image of the mi-crostructure in sample B1; (c) TEM image of the plate-type matrix in A1; (d) TEM image of the plate-type matrix in B1; (e) diffrac-tion pattern for twin martensite in A1; (f) diffraction pattern for twin martensite in B1.
Note: a the maximum fracture stress of B1 is used to represent σb.
ference was discussed by Yao et al. [19]. The large defect, such as an inclusion, a carbide, or a carbide cluster, was found to be responsible for the lower strength properties of steel B in the monotonic condition. Moreover, many works have claimed that inclusions always harmfully affect the fa-tigue strength of tool steels, and the larger the inclusion is, the worse its influence on cyclic properties will be [2, 6, 8-13]. A fractographic examination reveals that most of the
samples are fractured from an inclusion. However, the size level is different, the inclusion origins of steel A are 10-30 μm and that of steel B are all above 30 μm. When consider-ing the significant difference in fatigue strength as indicated in Table 2, it can well be understood that the inclusion size should be one of the most important factors to affect the mechanical properties of AISI 11. Moreover, the above findings also supply us with a good opportunity to study the effect of inclusion size on the fatigue properties of PM tool steels.
3.2. Crack initiation and propagation
Figs. 3 and 4 depict the nature and shape of crack nuclea-tion sites on the fracture surface of the failed specimens. Figs. 3(a)-3(d) indicate that steel A is mainly fractured from aluminum silicate. As illustrated in Figs. 3(e)-3(h), it is noted that the majority of strength-controlling flaws in steel B are Al2O3. From these energy dispersive spectroscopy (EDS) spectra, it is noted that the crack origins of steel A and steel B are both from brittle, non-metallic inclusions.
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Fig. 3. SEM micrographs (a, c, e, g) and EDS spectra (b, d, f, h) of fatigue crack initiation: (a-d) steel A; (e-h) steel B.
However, the SEM micrographs show that their sizes and shapes are different. Fig. 4(a) demonstrates that the inclu-sion origin of steel A with a size of 10-30 μm is mostly spherical or elliptic-like, whereas that of steel B, with a lar-ger size of above 30 μm, is even more irregular, as illus-trated in Fig. 4(b).
Moreover, the distances of the inclusion origin to the specimen surface (din) of the A1 and B1 samples show dif-ferent trends from the inclusion size (Din). Fig. 4(a) indicates that there is no correlation between the distance and the in-clusion size of A1. Its fatigue fracture is an internal failure
mode which was considered to be affected by the compres-sive residual stress (RS) according to Sohar, Yao, et al. [17-18, 20]. However, the inclusion size of B1 shows a ten-dency to increase with increasing distance in Fig. 4(b), and two failure modes are both observed on the fracture surface: when the inclusion is lying closely to the surface in the range of approximately 30-120 µm, it induces a subsurface rupture and the fish-eye is nearly semi-circle; when the in-clusion is far away from the surface at 120-1200 µm, it ini-tiates an interior fracture. The distinct subsurface initiation of B1 in this work is attributable to its relatively larger inclu-
612 Int. J. Miner. Metall. Mater., Vol.19, No.7, July 2012
Fig. 4. Relationship between inclusion size Din and location din. The shapes of solid data points in the figure represent inclusion shapes: (a) Din-din for A1; (b) Din-din for B1.
sion. First of all, it can be seen from Fig. 4 that B1’s mini-mum crack origins are larger than A1’s maximum ones and can appear at various locations within 30-390 µm. Second, Fig. 4(b) also indicates that the inclusion with a size no lar-ger than 40 µm, located at the subsurface, is large enough to cause much higher stress concentration than that at other sites far away. Besides that, the wide range of inclusion size distribution of B1, illustrated in Fig. 4(b), is beneficial to the potential nucleation site to be determined by inclusion size in the cyclic loading condition.
Fractographic observations illustrated in Fig. 5 demon-strate that the fish-eye size (Dfe) of A1 at the maximum stress of 1650 MPa is about 70-130 µm. The Dfe values for B1, B2, and B3 at a similar maximum stress of 1050 MPa are 170-420 µm, 220-550 µm, and 710-1510 µm, respectively. It is well known that the fish-eye, as an outstanding fatigue fracture characteristic, represents the stages of crack propa-gation. Sohar et al. [21] considered that the lower the stress amplitude was, the larger the obtained crack growth zone would be. In addition, it was also thought that the fish-eye size was dependent on the fracture toughness [20-22]. Thus, the difference in Dfe of A1 and B1 is attributed to the varied stress amplitude and that among B1, B2, and B3 is attributed to their different fracture toughness. Therefore, it can be claimed that the growth of the fish-eye is deter-mined by the combined effect of stress amplitude and frac-ture toughness.
3.3. Quantitative evaluation
Fig. 6 shows the relationship between inclusion size (Din) and cycle number to failure (Nf) for specimens that failed at a certain maximum stress on a log-log graph. The detailed investigation of the Din-Nf correlation for A1 samples was presented earlier [20]. It demonstrated that the fatigue life of A1 was determined by the size of the crack origin as the stress amplitude was constant. Besides that, the critical in-
clusion size at the maximum stress of 1650 MPa could be obtained from the well-fitted line, which turned out to be 11.7 μm. It can be clearly seen from Fig. 6 that the fatigue life of the B1 samples has a tendency to increase as the in-clusion size decreases at 1050 MPa. However, the data points of B1 show more deviation than that of A1. This should be attributed to the fact that the inclusion origins of B1 are always irregularly shaped, large-sized, and differently
Fig. 5. Fish-eye size Dfe for investigated samples.
Fig. 6. Relationship between inclusion origin size Din and cy-cle number to failure Nf for specimens failed at a certain maximum stress.
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located, which is clearly revealed in Fig. 4. According to the fitted line of B1, it was found that the inclusions with sizes less than 40.7 µm may not yield the fatigue fracture at up to 107 cycles at 1050 MPa. Therefore, it can be concluded that small inclusions are beneficial for PM tool steels to obtain excellent fatigue properties, such as a longer fatigue life and higher fatigue strength.
However, the critical inclusion size under the fatigue strength for each sample could not be obtained from the above method (Fig. 6, Din-Nf graph). In order to further un-derstand the effect of inclusion size on the fatigue strength of AISI 11, classic linear fracture mechanics is used. The stress intensity factor ΔK (ΔKin, ΔKfe, and ΔKth) can be esti-mated from the following equation [1, 23]:
a22πDK σΔ = (1)
where D is the equivalent diameter of a defect, 2σa is the stress range, and the effect of location on ΔK is ignored. As indicated in Fig. 6, the shapes of the inclusions have an in-fluence on the fatigue properties; however, this effect is al-ways ignored by Murakami, Meurling, Sohar, et al. [1, 6, 21]. Accordingly, the inclusions are all considered as inte-rior circular-like defects in this work. The stress intensity factor ranges of the inclusion origin (ΔKin) are calculated using Eq. (1), and the corresponding results are shown in Table 3. Meurling et al. [1] has recently demonstrated that ΔKin lies above ΔKth for PM tool steels and 4 MPa·m1/2 is a typical value of their threshold stress intensity factor. Therefore, with this consideration, the minimum value of ΔKin for each sample is seen as ΔKth. It was found that the ΔKin values for the samples at HRC 62-58 are all above 4 MPa·m1/2, based on which the rough estimation of ΔKth equal to 4 MPa·m1/2 for A1, A2, A3, and B1 samples was ob-tained. Similarly, ΔKth for B2 and B3 is estimated to be 4.2 and 4.4 MPa·m1/2, respectively. Generally, an accurate value of ΔKth is usually measured in crack growth rate experi-ments. However, for samples such as AISI 11 exhibiting high strength and low toughness, the tests are difficult to perform and the experimental results contain a certain error. Thus, much emphasis is laid upon the usage of fractography in this work. In addition, the stress intensity factor ranges of the fish-eye area (ΔKfe) are calculated as well. As shown in Table 3, it can be seen that the mean values of ΔKfe for A1, A2, A3, B1 and B2 samples are closely related to their own fracture toughness but that of B3 is found to be much higher than its measured value. It is noted that the fracture tough-ness of AISI 11 above HRC 56 can be evaluated by the mean value of the stress intensity factor of the fish-eye.
Table 3. Calculated stress intensity factor ranges of the inclu-sion origin and fish-eye area for specimens failed at 105-107 cy-cles.
Note: there is an insignificant difference in ΔKin and ΔKfe between A2 and A3, so the two values of A3 are not listed here.
A defect is defined as critical if its stress intensity factor exceeds the threshold value, ΔKth. Therefore, on the basis of Eq. (1), the critical inclusion size Dinc in a cyclic loading condition is obtained through
2
thinc
a
π2 2
KD
σ⎛ ⎞Δ
= ⎜ ⎟⎝ ⎠
(2)
When 2σa reaches 0.9σf (R = 0.1), the critical inclusion sizes of A1, A2, and A3 by Eq. (2) are estimated to be about 13.1, 15.3, and 16.3 µm, respectively, and those of B1, B2, and B3 are roughly considered to be 31.0, 34.0, and 36.0 µm, re-spectively. These calculations imply that the fatigue strength should depend greatly on the inclusion size and the temper-ing temperature. However, when the inclusions are as large as 30.0 µm, the fatigue strength would decrease rapidly and show little relationship with tempering just like the steel B samples.
4. Conclusions
The effects of inclusion size on the high cycle fatigue strength and the fatigue failure mode of AISI 11 have been investigated. The following results are presented.
(1) Brittle inclusions with sizes less than 13.1, 15.3, and 16.3 µm could not yield fractures under the fatigue strength of 1538, 1425, and 1380 MPa, respectively. However, in-clusions larger than 30.0 µm could cause fatigue strength independent of tempering temperature and maintain it at about 1000 MPa, and subsurface initiation could also be in-duced by them.
(2) The sample with a lower fracture toughness subjected to a higher stress amplitude was demonstrated to obtain a larger fish-eye. The critical inclusion size at the maximum stress of 1050 MPa for steel B tempered at 550°C was cal-culated to be 40.7 µm according to the correlated line be-tween inclusion size and fatigue life.
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(3) Based on linear fracture mechanics, the critical inclu-sion size, ΔKth and KIc can be evaluated by fractography as well. The stress intensity factor of the inclusion was found to lie above 4 MPa·m1/2. The mean value of the stress inten-sity factor of the fish-eye can be used to estimate the frac-ture toughness of the sample above HRC 56.
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